Lecture Notes in Civil Engineering

Wenhui Duan
Lihai Zhang
Surendra P. Shah Editors

Nanotechnology
in Construction
for Circular
Economy
Proceedings of NICOM7, 31 October–02
November, 2022, Melbourne, Australia
Lecture Notes in Civil Engineering

Volume 356

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Sheng-Hong Chen, School of Water Resources and Hydropower Engineering,
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Ioannis Vayas, Institute of Steel Structures, National Technical University of
Athens, Athens, Greece
Sanjay Kumar Shukla, School of Engineering, Edith Cowan University, Joondalup,
WA, Australia
Anuj Sharma, Iowa State University, Ames, IA, USA
Nagesh Kumar, Department of Civil Engineering, Indian Institute of Science
Bangalore, Bengaluru, Karnataka, India
Chien Ming Wang, School of Civil Engineering, The University of Queensland,
Brisbane, QLD, Australia
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Wenhui Duan · Lihai Zhang · Surendra P. Shah
Editors

Nanotechnology
in Construction for Circular
Economy
Proceedings of NICOM7, 31 October–02
November, 2022, Melbourne, Australia
Editors
Wenhui Duan Lihai Zhang
Department of Civil Engineering Department of Infrastructure Engineering
Monash University University of Melbourne
Clayton, VIC, Australia Parkville, VIC, Australia

Surendra P. Shah
Civil and Environmental Engineering
Northwestern University
Evanston, IL, USA

ISSN 2366-2557 ISSN 2366-2565 (electronic)

Lecture Notes in Civil Engineering
ISBN 978-981-99-3329-7 ISBN 978-981-99-3330-3 (eBook)
https://doi.org/10.1007/978-981-99-3330-3

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Organization

Advisory Committee

S. P. Shah
K. Wang
K. Sobolev
M. S. Konsta-Gdoutos
L. Ferrara

Scientific Advisory Committee

K. Wang (Chair)
S. Chen (Secretary)
L. Ferrara
G. A. Ferro
W. Young
A. Heidarpour
M. Yellishetty
P. Mendis
V. Sirivivatnanon
S. Setunge
C. Chen
B. Samali
D. Law
C. Li
Z. Tao
R. S. Nicolas
W. Li
L. Zhang

v
vi Organization

C. Caprani
H. Huang
D. Wu
B. Sainsbury
D. Robert
T. Yu
T. Ren
A. Remennikov
Y. Zhang
P. D. Silva
W. Gao
G. Li
X. Liu
K. Vessalas
R. Erkmen
R. Shrestha
J. Li
P. Thomas
N. Gowripalan
S. Nejadi
H. Wang
P. Mutton
R. Zou
Y. Huang
V. Tam
Y. Zhuge
M. Ghodrat
J. Zhao
Q. Zhang
M. M. Alam
K. Le

Local Organizing Committee

L. Zhang (Chair)
K. Sagoe-Crentsil (Co-Chair)
B. Chang (Secretary)
S. Miramini (Coordinator)
D. Chen (Coordinator)
S. Zhang
X. Yao
H. Sui
Y. Liu
Organization vii

F. Basquiroto
H. Nguyen
W. Wang

Award Committee

Y. Mai
D. Nethercot
J. Torero
S. T. Quek
R. Amal
M. Bradford
S. Kitipornchai
R. Kell
P. Phillip
A. Paradowska
R. Yeo
NICOM7 Preface

The Seventh International Symposium on Nanotechnology in Construction

(NICOM7), with the theme of “Nanotechnology in Construction for Circular Econ-
omy”, provided a bridge between research advances and industry/commercial oppor-
tunities within the construction sector underpinned by nanoscience. To this end,
NICOM7 showcased the latest developments across the entire supply chain involving
researchers, manufacturers, suppliers, and end users.
NICOM7 was delivered via a hybrid mode (onsite & online) for a total of 8
plenary/keynote speeches and 12 parallel sessions containing 61 presentations and
attracting nearly 100 attendees from Australia, China, Germany, Brazil, USA, India,
UK, and Singapore. It was jointly hosted by Monash University and the University
of Melbourne, Australia, and coordinated by an Advisory Committee (S. P. Shah,
USA; K. Wang, USA; K. Sobolev, USA; M. S. Konsta-Gdoutos, Greece; L. Ferrara,
Italy), Scientific Advisory Committee (49 members), Local Organising Committee
(12 members), and Awards Committee (11 members). The editors acknowledge
their excellent contributions that enabled successful delivery of NICOM7, as well as
support from all reviewers, helpers, and conference organizers.
This proceeding volume is based on contributions presented at NICOM7. The 56
high-quality papers/extended abstracts cover a wide range of topics across nanotech-
nologies in concrete structures, structural health monitoring, nanocomposite cement
replacements rails, pavements, AI and nanomodification of cementitious materials,
the alkali–silica reaction, concrete durability, geopolymer concretes, etc. The focus
of the proceedings is next-generation nanotechnologies for the broader construction
sector. The editors sincerely thank the authors for their outstanding contributions,
and hope the NICOM7 experience will promote innovation to enhance the rapid
promotion and integration of nanoscience into mainstream construction practices.
The editors also anticipate that the proceedings of NICOM7 will have great value

ix
x NICOM7 Preface

for both the research and engineering communities by advancing nanotechnology

initiatives towards a globally sustainable and innovative construction sector.

Clayton, Australia Surendra P. Shah

Honorary Chair
Parkville, Australia Wenhui Duan
Conference Chair
Evanston, USA Lihai Zhang
Conference Co-Chair
Contents

Nonlinear Wind-Induced Vibration Behaviors of Multi-tower

Suspension Bridges Under Strong Wind Conditions . . . . . . . . . . . . . . . . . . 1
R. Zhou, Y. J. Ge, Y. Yang, Y. D. Du, and L. H. Zhang
Thermal Transfer Effects of CRTS II Slab Track Under Various
Meteorological Conditions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11
R. Zhou, W. H. Yuan, Y. D. Du, H. L. Liu, and L. H. Zhang
Investigation on Superhydrophobicity and Piezoresistivity
of Self-sensing Cement-Based Sensors Using Silane Surface
Treatment . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 17
W. K. Dong, W. G. Li, X. Q. Lin, and S. P. Shah
Use of Brown Coal Ash as a Replacement of Cement in Concrete
Masonry Bricks . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 23
D. W. Law, C. Gunasekara, and S. Setunge
Composition of Alkali–Silica Reaction Products in Laboratory
and Field Concrete . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 27
M. J. Tapas, K. Vessalas, P. Thomas, N. Gowripalan,
and V. Sirivivatnanon
Behavior of Hybrid Engineered Cementitious Composites
Containing Nanocellulose . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 37
H. Withana and Y. X. Zhang
Investigation of ASR Effects on the Load-Carrying Capacity
of Reinforced Concrete Elements by Ultra-Accelerated Laboratory
Test . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 43
J. Cao, N. Gowripalan, V. Sirivivatnanon, and J. Nairn
3D printed Ultra-High Performance Concrete: Preparation,
Application, and Challenges . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53
G. Bai, G. Chen, R. Li, L. Wang, and G. Ma

xi
xii Contents

Nanosilica-Modified Hydrogels Encapsulating Bacterial Spores

for Self-healing Concrete . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 67
J. Feng and S. Qian
Reusing Alum Sludge as Cement Replacement to Develop
Eco-Friendly Concrete Products . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 75
Y. Liu, Y. Zhuge, and W. Duan
Role of Aggregate Reactivity, Binder Composition, and Curing
Temperature on the Delayed Ettringite Formation and Associated
Durability Loss in Concrete . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 83
L. Martin, P. Thomas, P. De Silva, and V. Sirivivatnanon
Effect of Blending Alum Sludge and Ground Granulated
Blast-Furnace Slag as Cement Replacement to Mitigate
Alkali-Silica Reaction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 93
W. Duan, Y. Zhuge, and Y. Liu
Optimisation of Limestone Calcined Clay Cement Based
on Response Surface Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 103
G. Huang, Y. Zhuge, T. Benn, and Y. Liu
Designing Waterborne Protective Coatings Through Manipulating
the Nanostructure of Acrylic-Based Nanocomposites . . . . . . . . . . . . . . . . . . 113
S. Ji, H. Gui, G. Guan, M. Zhou, Q. Guo, and M. Y. J. Tan
Analysis of Categories That Delay Global Construction Projects . . . . . . . 127
M. Abonassrya, M. Alam, and A. Saifullah
Chloride Penetration in Low-Carbon Concrete with High Volume
of SCM: A Review Study . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 141
C. Xue and V. Sirivivatnanon
A Compact Review on the Waste-Based Lightweight Concrete:
Advancement and Possibilities . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 151
M. M. U. Islam, J. Li, R. Roychand, and M. Saberian
Influence of Reinforcement on the Loading Capacity of Geopolymer
Concrete Pipe . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 165
S. Dangol, J. Li, V. Sirivivatnanon, and P. Kidd
Creep of Slag Blended Cement Concrete with and Without
Activator . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 177
H. T. Thanh, M. J. Tapas, J. Chandler, and V. Sirivivatnanon
Partially-Unzipped Carbon Nanotubes as Low-Concentration
Amendment for Cement Paste . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 187
S. Iffat, F. Matta, J. Gaillard, M. Elvington, M. Sikder, M. Baalousha,
S. Tinkey, and J. Meany
Contents xiii

Effect of Fine Aggregates and Test Settings on the Self-sensing

Response of Cement-Based Composites with Carbon Nanotubes
as Conductive Filler . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 197
T. C. dos Santos, P. A. Carísio, A. P. S. Martins, M. D. M. Paiva,
F. M. P. Gomes, O. A. M. Reales, and R. D. Toledo Filho
Effect of Carbonation on the Microstructure and Phase
Development of High-Slag Binders . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 213
M. J. Tapas, A. Yan, P. Thomas, C. Holt, and V. Sirivivatnanon
A New Dispersion Strategy to Achieve High Performance
Graphene-Based Cement Material . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 223
Z. Zhang, Y. Yao, H. Liu, Y. Zhuge, and D. Zhang
Accelerated Mortar Bar Test to Assess the Effect of Alkali
Concentration on the Alkali–Silica Reaction . . . . . . . . . . . . . . . . . . . . . . . . . 233
B. Boyd-Weetman, P. Thomas, P. DeSilva, and V. Sirivivatnanon
Development of High-Strength Light-Weight Cementitious
Composites with Hollow Glass Microspheres . . . . . . . . . . . . . . . . . . . . . . . . . 241
X. Li, Y. Yao, D. Zhang, Z. Zhang, and Y. Zhuge
Co-effects of Graphene Oxide and Silica Fume on the Rheological
Properties of Cement Paste . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 251
D. Lu, Z. Sheng, B. Yan, and Z. Jiang
Automated 3D-Printer Maintenance and Part Removal by Robotic
Arms . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 259
K. Andrews, K. Granland, Z. Chen, Y. Tang, and C. Chen
Machine Vision-Based Scanning Strategy for Defect Detection
in Post-Additive Manufacturing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 271
S. Zhang, Z. Chen, K. Granland, Y. Tang, and C. Chen
Electrical and Sulfate-Sensing Properties of Alkali-Activated
Nanocomposites . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 285
Maliheh Davoodabadi, Marco Liebscher, Massimo Sgarzi,
Leif Riemenschneider, Daniel Wolf, Silke Hampel,
Gianaurelio Cuniberti, and Viktor Mechtcherine
Advances in Characterization of Carbonation Behavior
in Slag-Based Concrete Using Nanotomography . . . . . . . . . . . . . . . . . . . . . . 297
B. Mehdizadeh, K. Vessalas, B. Ben, A. Castel, S. Deilami, and H. Asadi
Application of Surface-Modified Nanosilica for Performance
Enhancement of Asphalt Pavement . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 309
D. He, H. H. Chan, Z. H. Xiao, T. Wu, L. M. Leung, M. Sham,
B. Chen, S. F. S. Lee, and C. K. K. Kwan
xiv Contents

Effect of Different Additives on the Compressive Strength of Very

High-Volume Fly Ash Cement Composites . . . . . . . . . . . . . . . . . . . . . . . . . . . 313
R. Roychand, J. Li, M. Saberian, S. Kilmartin-Lynch, M. M. Ul Islam,
M. Maghfouri, and F. Chen
Spalling Resistance of Hybrid Polyethylene and Steel
Fiber-Reinforced High-Strength Engineered Cementitious
Composite . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 321
S. Rawat, Y. X. Zhang, and C. K. Lee
Roads Issues and the Social License to Operate . . . . . . . . . . . . . . . . . . . . . . . 327
W. Young and M. Shackleton
An Intelligent Multi-objective Design Optimization Method
for Nanographite-Based Electrically Conductive Cementitious
Composites . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 339
W. Dong, Y. Huang, B. Lehane, and G. Ma
Machine Learning-Aided Nonlinear Dynamic Analysis
of Engineering Structures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 347
Y. Feng, Q. Wang, D. Wu, and W. Gao
Insights into the Size Effect of the Dynamic Characteristics
of the Perovskite Solar Cell . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 353
Q. Li, D. Wu, and W. Gao
Non-probabilistic Informed Structural Health Assessment
with Virtual Modelling Technique . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 359
Q. Wang, Y. Feng, D. Wu, and W. Gao
Modeling the Alkali–Silica Reaction and Its Impact
on the Load-Carrying Capacity of Reinforced Concrete Beams . . . . . . . . 365
T. N. Nguyen, J. Li, V. Sirivivatnanon, and L. Sanchez
Guidelines for Enzymatic Soil Stabilization . . . . . . . . . . . . . . . . . . . . . . . . . . 373
B. O’Donnell, A. Swarup, A. Sidiq, D. Robert, and S. Setunge
Deterioration Modeling of Concrete Bridges and Potential
Nanotechnology Application . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 399
H. Tran and S. Setunge
Transfer and Substrate Effects on 2D Materials for Their Sensing
and Energy Applications in Civil Engineering . . . . . . . . . . . . . . . . . . . . . . . . 409
Q. Zhang, C. Zheng, K. Sagoe-Crentsil, and W. Duan
Experimental and Numerical Studies on the In-Plane Shear
Behavior of PVC-Encased Concrete Walls . . . . . . . . . . . . . . . . . . . . . . . . . . . 421
Kamyar Kildashti and Bijan Samali
Contents xv

Recycled Glass-Based Capping Layer for Foundations in Expansive

Soils . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 431
H. Karami, D. Robert, S. Costa, J. Li, S. Setunge, and S. Venkatesan
Submicroscopic Evaluation Studies to Minimize Delayed Ettringite
Formation in Concrete for a Sustainable Industry and Circular
Economy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 445
Yogesh Kumar Ramu, Paul Stephen Thomas, Kirk Vessalas,
and Vute Sirivivatnanon
A Novel Concrete Mix Design Methodology . . . . . . . . . . . . . . . . . . . . . . . . . . 457
D. Kumar, M. Alam, and J. Sanjayan
Characterization of the Nano- and Microscale Deterioration
Mechanism of the Alkali–Silica Reaction in Concrete Using
Neutron and X-ray Scattering Techniques: A Review . . . . . . . . . . . . . . . . . 469
E. Nsiah-Baafi, M. J. Tapas, K. Vessalas, P. Thomas,
and V. Sirivivatnanon
Life Cycle Assessment of the Environmental Impacts of Virgin
Concrete Replacement by CO2 Concrete in a Residential Building . . . . . 479
M. Ma, Y. Zhou, V. W. Y. Tam, and K. N. Le
Economic Impacts of Environmentally Friendly Blocks: The Case
of Nu-Rock Blocks . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 483
V. W. Y. Tam, K. N. Le, I. M. C. S. Illankoon, C. N. N. Tran,
D. Rahme, and L. Liu
Life Cycle Assessment (LCA) of Recycled Concrete Incorporating
Recycled Aggregate and Nanomaterials . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 491
W. Xing, V. Tam, K. Le, J. L. Hao, J. Wang, and P. Yang
Analysis of the Compressive Strength of CO2 Concrete While
Eliminating Overshadowing Concrete Variables . . . . . . . . . . . . . . . . . . . . . . 495
V. W. Y. Tam, A. Butera, K. N. Le, and L. Liu
Harmonic Vibration of Inclined Porous Nanocomposite Beams . . . . . . . . 497
D. Chen and L. Zhang
Influence of Carbon Nanotubes on the Fracture Surface
Characteristics of Cementitious Composites Under the Brazilian
Split Test . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 503
Y. Gao, J. Xiang, Z. Yu, G. Han, and H. Jing
Effect of Carbon Nanotubes on the Acoustic Emission
Characteristics of Cemented Rockfill . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 513
Z. Yu, H. Jing, Y. Gao, X. Wei, and A. Wang
xvi Contents

Effects of Graphene Oxide Content on the Reinforcing Efficiency

of C–S–H Composites: A Molecular Dynamics Study . . . . . . . . . . . . . . . . . 521
W. Chen, J. Xiang, Y. Gao, and Z. Zhang
Graphene-Induced Nano- and Microscale Modification of Polymer
Structures in Cement Composite Systems . . . . . . . . . . . . . . . . . . . . . . . . . . . . 527
Z. Naseem, K. Sagoe-Crentsil, and W. Duan
Nonlinear Wind-Induced Vibration
Behaviors of Multi-tower Suspension
Bridges Under Strong Wind Conditions

R. Zhou, Y. J. Ge, Y. Yang, Y. D. Du, and L. H. Zhang

Abstract The aerodynamic characteristics of a multispan suspension bridge differ

from those of a two-span suspension bridge. In this study we investigated the
nonlinear aerodynamic characteristics of the Maanshan Bridge under nonstationary
flow using combination quasi-3D finite element (FE) bridge models of 2D nonlinear
aerodynamic force models and 3D nonlinear FE bridge models. The developed model
predictions were validated by wind tunnel tests involving a 2D sectional stiffness
model and 3D full-bridge aeroelastic model. Results showed that the developed
model could potentially describe the nonlinear and unsteady aerodynamic effects on
the bridge. Furthermore, the flutter behavior of the Maanshan Bridge under uniform
flow changed from the stable limit cycle of soft flutter to unstable limit cycle with
the disconnection of two hangers at the 1/2L of the right main span, while the flutter
behavior of the bridge under turbulence flow could be defined as the fracture failure
of the hangers from the 1/2L of the left main span.

Keywords Failure mode · Multispan suspension bridges · Nonlinear flutter

behavior · Quasi-3D FE model

1 Introduction

Due to the advantages of not sharing anchorage, relatively shorter main span and
lower cost, more and more multi-tower suspension bridges are being built worldwide,
especially for the bridges with the main span >1000 m [1]. The Taizhou and Maanshan

R. Zhou (B) · Y. D. Du
College of Civil and Transportation Engineering, Shenzhen University, Shenzhen, China
e-mail: zhourui@szu.edu.cn
Y. J. Ge · Y. Yang
State Key Lab for Disaster Reduction in Civil Engineering, Tongji University, Shanghai, China
L. H. Zhang
Department of Infrastructure Engineering, University of Melbourne, Melbourne, VIC, Australia

© The Author(s) 2023 1

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_1
2 R. Zhou et al.

Yangtze River bridges with a main span of 1080 m are two excellent examples of super
long-span three-tower suspension bridges in China. The structural characteristics
and aerodynamic performance of a multi-tower suspension bridge are significantly
different from that of a two-tower suspension bridge [2], and they are susceptible to
flutter instability under wind loading [3]. Under strong wind loads, the wind-induced
vibrations and post-critical flutter behaviors of long-span multi-tower bridges remain
challenges for wind engineers.
Because the nonlinear flutter behavior of long-span multi-tower bridges is a
complex phenomenon involving aerodynamic nonlinearities and structural nonlin-
earities due to structural large deformations, which can result in damage of partial
components, and ultimately the collapse of the whole bridge structure [4]. In the
past decades, most research focused on modeling the self-excited forces using
rational or indicial functions in the time domain (e.g., Chen and Kareem [5]; Diana
et al. [6]). However, the unsteady and nonlinear effects of the aerodynamic forces
produced by the wind–structure interaction were been simultaneously considered
in those studies. Recently, Wu and Kareem [7] presented a nonlinear convolution
scheme based on Volterra-Wiener theory, and Arena et al. [8] used a nonlinear
quasi-steady aerodynamic model for time-periodic oscillations of suspension bridges
and conducted global bifurcation analysis of post-critical behaviors. Gao et al.
proposed a nonlinear self-excited force model in terms of nonlinear flutter deriva-
tives [9], and Xu et al. studied the flutter performance and hysteresis phenomena
of a streamlined bridge deck sectional model using a large-amplitude free vibra-
tion test [10]. More importantly, Liu proposed a nonlinear aerodynamic force model
(NAFM) based on nonlinear differential equations that could produce aerodynamic
hysteresis phenomena [11], and Zhou et al. developed the NAFM by considering the
vortex-induced force and then analyzed the nonlinear wind-induced behavior of long-
span bridges [12, 13]. Through a series of wind-tunnel tests and three-dimensional
nonlinear finite-element (FE) analyses, we further compared the comprehensive
wind-resistance performance of a suspension bridge with various slot ratios [12],
vertical stabilizers [14, 15], grid porosities [16], guide plates [17, 18], and combina-
tion of aerodynamic measures [19–21]. Previous studies have shown that nonlinear
FE models of a 3D bridge incorporated with the NAFM is a commendable approach
to simulating the nonlinear behavior of wind-induced vibration of bridges under
strong wind.
This study aimed to understand the nonlinear dynamic behaviors in flutter and
post-flutter of multi-tower suspension bridges under strong wind excitation. Firstly,
the 2D displacement responses and the aerodynamic forces in the NAFM of a closed-
box girder were calculated based on the computational fluid dynamics (CFD) simula-
tion. Subsequently, an integrated numerical approach for a 3D three-tower suspension
bridge under different wind excitations using a combination of a NAFM and nonsta-
tionary flows was developed. Finally, the nonlinear displacement responses and
flutter collapse of the 3D bridge under uniform and turbulent flow, respectively, were
analyzed. The present study could potentially contribute to further understanding of
the flutter mechanism of multi-tower suspension bridges.
Nonlinear Wind-Induced Vibration Behaviors of Multi-tower … 3

2 NAFM of Bridges

2.1 Three-Tower Suspension Bridge

A typical three-tower suspension bridge with a span arrangement of 360 + 2 × 1080

+ 360 m and a sag-to-span ratio of the main cable of 1/9 was studied. All three
towers are 176 m high with the middle tower being 128 m above the deck, and each
of the side towers was 143 m above the deck. The distance between the two cables
was 35 m, and the spacing between two adjacent hangers was 16 m. As shown in
Fig. 1, the deck cross-section of the bridge was a closed streamlined box steel girder
of 38.5 m wide × 3.5 m deep, in which the vertical and torsional frequencies were
0.08 ad 0.26 Hz, respectively.

2.2 2D Displacement Responses of the Closed-Box Girder

Based on the fluid–structure interaction in the CFD simulation, an unstructured grid

system incorporating large eddy simulation modeling with a Smagorinsky subgrid-
scale model was used to discretize the governing Navier–Stokes equations. A steady
uniform flow velocity was applied at the inlet boundary, and an opening pressure
condition was given at the outlet boundary. The overall mesh division around two
closed-box girders in the CFD simulations with the computational domain of a 60B
× 40B rectangle is described in Fig. 2a. As illustrated in Fig. 2b, the vertical displace-
ment of the bridge deck remained stable when the wind velocity (U) increased from
0 to ≈78 m/s, and then rapidly increased when the U was >80 m/s. Finally, the
maximum vertical and torsional displacement responses were as high as ≈2.5 h/H
and 80°, respectively. It can be seen that the torsional displacement reached ≈5°
under U = 75.5 m/s as the soft flutter phenomenon. Therefore, the critical flutter
wind velocity of the three-tower suspension bridge from the CFD simulation was
147.5 m/s. Moreover, the hysteresis loops of FD , FL , and FM predicted by the NAFM
were close to those from the CFD simulation shown in Fig. 2c, d. The parameters of
static force, self-excited force, and buffeting force in the NAFM of the closed-box
girder were identified.

Fig. 1 Cross-section of the closed-box girder (unit: m)

4 R. Zhou et al.

Displacement responses of a closed-box girder under free vibration

(a) (b) 3 Vertical
100

Torsion
2 75

1 50

0 25
← Vertical

α(deg)
Y/H
-1 0
→ Torison
-2 -25

-3 -50

-4 -75

-5 -100
73 74 75 76 77 78 79 80 81 82 83
U(m/s)

Hysteresis curves of the calculated aerodynamic forces by CFD Hysteresis curves of the fitted aerodynamic forces by NAFM
(c) 1.5 Ur=1 Ur=2 (d) 1.5 Ur=1 Ur=2
Ur=4 Ur=6 Ur=4 Ur=6

Dimensionless Lift Force FL

1.0 1.0
Dimensionless Lift Force FL

Ur=8 Ur=10 Ur=8 Ur=10

0.5 0.5

0.0 0.0

-0.5 -0.5

-1.0 -1.0
-1.5 -1.5
-2.0 -2.0
-20 -15 -10 -5 0 5 10 15 20 -20 -15 -10 -5 0 5 10 15 20
Torsional angles (deg) Torsional angles (deg)

Fig. 2 Parameters in the NAFM: a overview of the CFD simulation of the closed-box deck;
b vertical and torsional displacement responses; c, d hysteresis curves of lifting force and torsional
angle of the CFD and NAFM

3 Nonlinear Flutter and Post-Flutter Behaviors

of a Multi-tower Bridge

3.1 Turbulence Flow at Bridge Site

The reference height and the corresponding average wind velocity at the bridge site
were Zref = 57.83 m and Uref = 39.3 m/s, respectively, and z0 = 0.01 and α = 0.12
because the bridge site belongs to the B-type terrain. Based on the combination of
weighted amplitude wave superposition and Fast Fourier Transform (FFT) technique,
the time histories of the vertical and horizontal (along-bridge) wind velocities at the
height of z = 54.8 m at the middle point of the two main spans, are shown in Fig. 3.
Nonlinear 3D FE models of the three-tower suspension bridge were established
with a total of 1228 elements. The nonlinear governing coupled equations in the
integrated FE model were numerically solved using the Newton–Raphson method
in combination with the Newmark-β method. Accordingly, the turbulent flow at the
bridge site was firstly simulated to reflect the influence of turbulent flow on aero-
dynamic performance. Then the nonlinear behavior of the displacement responses,
structural frequencies, oscillation configurations, and failure modes of the 3D bridge
under uniform and turbulent flow, respectively, were obtained.
Nonlinear Wind-Induced Vibration Behaviors of Multi-tower … 5
y p p p g y p p p g
20 120
Vertical Horizontal
15 (a) 100
10 (b)
80
5
Uz (m/s)

Uy (m/s)
0 60

-5
40
-10
20
-15

-20 0
0 500 1000 1500 2000 2500 0 500 1000 1500 2000 2500
Time(s) Time(s)

20 120
Vertical Horizontal
15
(c) 100 (d)
10
80
5
Uz (m/s)

Uy (m/s)
0 60

-5
40
-10
20
-15

-20 0
0 500 1000 1500 2000 2500 0 500 1000 1500 2000 2500
Time (s) Time (s)

Fig. 3 Time histories of turbulent flow of a multi-tower suspension bridge: a, b vertical and hori-
zontal wind velocities at the left main span; c, d vertical and horizontal wind velocities at the right
main span

3.2 Flutter and Post-Flutter Behaviors Under Uniform Flow

As presented in Fig. 4, the relative vertical displacement (Y/H) and torsional displace-
ment responses (α) of the right main span rapidly decreased and approached a balance
location at wind velocity of U = 70 m/s. However, there was an obvious soft flutter
phenomenon of the bridge when the wind velocity increased to U = 75 m/s. In partic-
ular, the vertical and torsional displacements gradually become larger after 100 s,
and then maintained a sinusoidal oscillation after 250 s with the relative value of
Y/H ∼ = 0.5 and α = 3.5°. Subsequently, the displacement responses presented an
ever-increasing trend with increasing wind velocity. Both the vertical and torsional
displacements rapidly increased after 60 s under U = 82.5 m/s and divergence finally
occurred with the extreme values of Y/H ∼ = 5 and α = 30°. As a result, the soft flutter
phenomenon of the bridge under U = 75 m/s fell into a stable limit cycle. In addition,
Fig. 4 shows that the spatially dependent lateral amplitudes along the bridge span
were generally very small in comparison with the vertical or torsional components
of oscillation, and the motion configuration of the two main spans was an antisym-
metrical vertical and torsional coupled oscillation. The failure of the whole bridge
occurred after two hangers were finally damaged at the middle ½ L of the right main
span under U = 82.5 m/s. Therefore, the whole flutter collapse of the three-tower
suspension bridge under uniform flow can be defined as the change from the stable
limit cycle of soft flutter to the unstable limit cycle of bending-torsional coupled
divergence.
6 R. Zhou et al.

-0.05 0.2
U=70 m/s U=70 m/s
0.1
-0.15 (a) (b)

α(deg)
Y/H

0
-0.25
-0.1

-0.35 -0.2
0 100 200 300 400 500 600 0 100 200 300 400 500 600
Time (s) Time (s)
p p g p g g p p g p g g
1 5
U=75 m/s U=75 m/s
0.5 2.5

α(deg)
Y/H

0 0

-0.5 -2.5

-1 -5
0 50 100 150 200 250 300 350 0 50 100 150 200 250 300 350
Time (s) Time (s)

5 20
U=82.5 m/s U=82.5 m/s
10
2.5

α(deg)
0
Y/H

0
-10
-2.5
-20

-5 -30
0 30 60 90 120 150 0 30 60 90 120 150
Time (s) Time (s)

Fig. 4 Time histories of displacement responses: a vertical; b torsional

3.3 Displacement Responses and Failure Modes Under

Turbulent Flow

As shown in Fig. 5, the time-dependent torsional displacement response showed a

nonlinear growth rate with increasing wind velocity, in which the maximum α was
close to 8° under U = 72 m/s and reached 20° under U = 75 m/s. Additionally, the
amplitude of torsional oscillation was greatest in three direction vibrations, and the
motion configuration of the two main spans was antisymmetrical coupled oscillation.
As for the failure mode of the bridge, the failure of the hangers started from those
located in the middle ½ L of the left main span of the bridge. Therefore, the whole
flutter collapse of the three-tower suspension bridge under turbulent flow directly
shifted from the paroxysmal bifurcation to the chaos of bending-torsional coupled
divergence.

3.4 Comparison of Displacement Responses

Finally, the calculated relationship between the displacement responses and wind
velocity using the integrated approach were compared with the experimental results
of full-bridge aeroelastic model wind-tunnel tests [22]. It can be seen that all the calcu-
lated Ucr are generally higher than the checked flutter wind velocity. The minimum
experimental Ucr was 74.2 m/s at the wind attack angle of +3° under uniform flow,
which was lower than the calculated value of Ucr = 82.5 m/s. However, the experi-
mental measurement (i.e., Ucr = 85.8 m/s) was higher than the calculated value (i.e.,
Ucr = 75 m/s) under turbulent flow. Moreover, Fig. 6 shows the relationship between
Nonlinear Wind-Induced Vibration Behaviors of Multi-tower … 7

(a)
(b)
Anti-clockwise

Clockwise

(c) (d)
4
UY

3
UZ
b*RotX
Hangers’ fracture
from the 1/2L deck
2
Vibration amplitude

-1

-2 U=82.5m/s
-3

-4
-1080-945 -810 -675 -540 -405 -270 -135 0 135 270 405 540 675 810 945 1080
Main Span

Fig. 5 Flutter collapse of the bridge: a clockwise rotation of the left main span; b clockwise rotation
of the right main span; c oscillation configuration; d failure mode

the maximum displacement responses and wind velocity at the midspan, the quarter
point (1/4 L near the side tower) and the three-quarter point (3/4 L near the middle
tower) of the bridge. It demonstrates that all three calculated maximum displacement
responses under uniform flow showed a stable upward trend with increasing wind
velocity, and the trend became more dramatic when U was over certain threshold
(i.e., U = 75 m/s, soft flutter). In addition, the maximum value of torsional displace-
ment (i.e. Ucr = 82.5 m/s) was the largest among the three dispacements, followed
by the vertical displacement. It should also be mentioned that there was a sudden
increase in all three experimental maximum displacement responses when the value
of U approached 90 m/s under 0° wind attack. Further, all three maximum displace-
ment responses at the midspan of the bridge were the largest compared with other
bridge locations, while the values of the displacement responses at ¼ L near the side
tower were the smallest. Although the values of all three maximum displacement
responses under turbulent flow also increased with increasing the wind velocity, the
growth rates under turbulent flow were much higher than those under uniform flow
(Fig. 7).
8 R. Zhou et al.

(a) (b)
8 20
U=72 m/s under turbulence flow U=75 m/s
6

4 10

α(deg)
α(deg)

0
0

-2
-10
-4

-6
-20
-8
0 50 100 150 200 250 300 0 20 40 60 80 100 120 140
Time (s) Time (s)

5
(c) (d)
UY
UZ
4
b*RotX
Hangers’ fracture
3
from the 1/2L deck
2
Vibration amplitude

-1

-2
U=75m/s
-3

-4
-1080-945 -810 -675 -540 -405 -270 -135 0 135 270 405 540 675 810 945 1080
Main Span

Fig. 6 Flutter collapse of the bridge under turbulent flow: a, b torsional displacement responses at
U = 72 m/s and U = 75 m/s; c oscillation configuration; d failure mode

25 25

20
Test-Lateral
Calculation-Lateral
(a) Test-Lateral
Calculation-Lateral (b)
Test-Vertical 20
Test-Vertical
Calculation-Vertical Calculation-Vertical
15 15
Test-Torsional Test-Torsional
Calculation-Torsional Calculation-Torsional
Maximum values values

10 10
Maximum values

5 5

0 0

-5 -5

-10 -10

-15 -15

-20 -20
0 10 20 30 40 50 60 70 80 90 100 0 10 20 30 40 50 60 70 80 90 100
U (m/s) U (m/s)

Fig. 7 Comparison of displacement responses: a, b 1/2 point under uniform flow and turbulent
flow

4 Concluding Remarks

In this study, a time-dependent numerical approach was developed to investigate the

nonlinear flutter and post-flutter behaviors of a three-tower suspension bridge under
different wind excitations. The major findings were:
(1) The oscillation configuration of the bridge was very dependent on the coupled
antisymmetrical bending-torsional oscillation of the two main bridge spans. The
flutter performance under uniform flow (Ucr = 82.5 m/s) was better than that
under turbulent flow (Ucr = 75 m/s) due to the increase in the vertical degree
participation in the coupled motion.
Nonlinear Wind-Induced Vibration Behaviors of Multi-tower … 9

(2) Under uniform flow, the flutter collapse process of the bridge can be described
as the change from the stable limit cycle of soft flutter to the unstable limit cycle
with disconnection failure of two hangers at the middle of the right main span.
(3) Under turbulent flow, the failure mode of the bridge can be described as the
sequential fracture failure of multiple hangers at the middle of the left main
span of the bridge, and the flutter collapse process of the bridge can be defined
as the direct shift from the paroxysmal bifurcation to chaos.
The present study presents some new sights into nonlinear dynamic behaviors in
the flutter collapse process of a three-tower suspension bridge. It should be mentioned
that the spatial spanwise effects along the bridge of the aerodynamic forces were not
taken into account in this study and should be investigated in future research.

Acknowledgements The authors gratefully acknowledge the support for the research work
jointly provided by the Guangdong Province Natural Science Foundation (No. 2023A1515030148,
2019B111106002, 2019A1515012050), National Science Foundations of China (Nos. U2005216,
51908374, and 52178503), and the Shenzhen Science and Technology Program under grant (Nos.
JCYJ20220531101609020, KQTD20180412181337494, and ZDSYS20201020162400001).

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stability of a long-span triple-tower suspension bridge. J Wind Eng Ind Aerodyn 128:12–21
2. Zhou R, Ge YJ, Yang YX et al (2020) Aerodynamic performance evaluation of different
cable-stayed bridges with composite decks. Steel Compos Struct 34(5):699–713
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suspension bridge based on structural nonlinearity[J]. Structures 34:787–803
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non-Gaussian and nonlinear features. J Wind Eng Ind Aerodyn 122:21–37
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aerodynamic nonlinearities. J Eng Mech 129(8):885–895
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nonlinearities in time domain. J Wind Eng Ind Aerodyn 96(10–11):1871–1884
7. Wu T, Kareem A (2013) A nonlinear convolution scheme to simulate bridge aerodynamics.
Comput Struct 128:259–271
8. Arena A, Lacarbonara W, Marzocca P (2016) Post-critical behavior of suspension bridges
under nonlinear aerodynamic loading. J. Comput. Nonlinear Dyn 11(1):011005:1–11
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bridges with various aerodynamic shapes. Nonlinear Dyn 94:1095–1115
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effects on twin-box girder suspension bridges. Eng Struct 183:1072–1090
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Thermal Transfer Effects of CRTS II
Slab Track Under Various
Meteorological Conditions

R. Zhou, W. H. Yuan, Y. D. Du, H. L. Liu, and L. H. Zhang

1 Introduction

With the evolution of climate change, the thermal transfer effects of ballastless
track in high-speed railways under complicated environmental conditions becomes
increasingly important, governed by a number of meteorological factors, including
solar radiation, ambient temperature, wind speed and direction, humidity, and many
others [1]. Because these meteorological factors are highly site-specific, the huge
area traversed by high-speed railway in China is affected by varying meteorolog-
ical conditions that could have a significant effect on the mechanical behavior of
track–bridge systems. The China Railway Track System type II (CRTS II) is a typical
ballastless slab track for high-speed railway systems. In order to guarantee good struc-
tural performance for long-term operation, it is necessary to study the influence of
meteorological conditions on the thermal transfer effects of CRTS II track.
In recent years, more and more researchers have studied the temperature field
and the CRTS II track based on monitoring of field data. Dai et al. [2] and Huang
et al. [3] investigated the temperature distribution characteristics of the CRTS II
track using conventional statistical methods, while Yang et al. [4] and Song et al. [5]
revealed the relationship between meteorological factors and the internal temperature
of the CRTS II track through finite element models analysis. Using temperature tests
of scaled models, Cai et al. [6] and Zhou et al. [7] investigated the influence of
cyclic and overall temperature on the displacement, strain, and temperature field
of a scaled CRTS II track–bridge structure. Furthermore, Zhu et al. [8] and Zhang
et al. [9] explored interfacial damage development of CRTS II track under complex

R. Zhou (B) · W. H. Yuan · Y. D. Du · H. L. Liu

College of Civil and Transportation Engineering, Shenzhen University, Shenzhen, China
e-mail: zhourui@szu.edu.cn
L. H. Zhang
Department of Infrastructure Engineering, University of Melbourne, Melbourne, VIC, Australia

© The Author(s) 2023 11

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_2
12 R. Zhou et al.

temperature conditions in a cohesive zone model (CZM) and concrete damaged

plasticity model. In addition, Zhou et al. [10–12] studied the mechanical behavior of
CRTS II track under the coupling effect of train and environment loads.

2 Meteorological Parameter Collection

Using a typical zone of CRTS II track on a bridge in the Beijing–Shanghai high-speed

railway system as a case study, the meteorological data and internal temperature of
the track structure were collected for 6 months. The relationships between ambient
temperature and the temperature at the midspan and end of the bridge are depicted
in Fig. 1, and the relationships among four meteorological parameters (tempera-
ture, solar radiation, wind speed, humidity) were also investigated. The change in
the temperature within the track structure obviously lagged behind the change in
the ambient temperature, and the temperature variation was smaller than that of the
ambient temperature. Furthermore, Pearson correlation analysis showed a strong
positive correlation between the ambient temperature and solar radiation, with a
Pearson correlation coefficient of 0.85, while the correlation coefficient of 0.22
between the ambient temperature and air humidity was the smallest.

3 Heart Transfer Model

Based on the finite element (FE) software of Comsol, the heat transfer numerical
models of CRTS II track on a simply-supported box bridge were established (Fig. 2a).
The total dimensions of the five slab tracks were 32 m length × 13.4 m width ×
3.35 m height, and the track slab, the CAM layer, base plate and box girder were
simulated by the solid elements. Two ends of the slab tracks were constrained, and
the thermal responses of the track structure under the four meteorological parame-
ters were compared. As shown in Fig. 2b, c, the internal temperature and vertical
displacement at the mid-span of the slab tracks become larger with increasing wind
speed or solar radiation, especially for wind speeds >6 m/s or solar radiation >750 W/
m2 . According to the 3D temperature and stress field shown in Fig. 2d–f, increasing
solar radiation or ambient temperature could lead to rapid heat transfer from the slab
track to the base plate. The role of wind speed on the heat transfer effect in the track
structure was limited.
 (b) (c)
N
20 NNW NNE
(a) 18
16 ——Winter
NW NE
14
12
10 WNW ENE
8 >= 8
6 7-8
4 6-7
2
0 W E 5-6
2 4-5
4 3-4
6 2-3
8
10 WSW ESE 1-2
12 0-1
14
16 SW SE
18
20 SSW SSE
S
Thermal Transfer Effects of CRTS II Slab Track Under Various …

Fig. 1 Four meteorological parameters evaluated

13
14 R. Zhou et al.

(a) (b) 40 3.5

Vertical displacement of track slab (mm)

3.0

Temperature of track slab (℃)

38
2.5
36
2.0
34
1.5
32
1.0
30 Temperature 0.5
Displacement
28 0.0

0 3 6 9
Wind speed (m/s)

56 7
(c) (d) t(73)=18h Temperature (degC)

Vertical displacement of track slab (mm)

54 Temperature
Temperature of track slab (℃)

Displacement 6
52
50
5
48
46
4
44
42 3
40
38 2
0 600 750 1100
solar radiation (W/m2)

(e) t(73)=18h Temperature (degC)

(f) t(73)=18h Stress x (MPa)

Fig. 2 Thermal responses comparison under meteorological parameters: a FE model; b, c temper-

ature and displacement responses under various wind speeds and solar radiation levels; d,
e temperature field under 1100 W/m2 solar radiation; f stress field under 1100 W/m2 solar radiation

4 Conclusions

Based on the combination of field measurement data and FE analysis, the thermal
transfer effects in a CRTS II slab track–bridge system under various meteorological
conditions were studied. The major findings were:
(1) Three meteorological conditions—ambient temperature, solar radiation, and
wind speed—had large correlation coefficients, showing they had the greatest
influence on thermal transfer in the track structure.
(2) Increasing solar radiation or ambient temperature could lead to increasing defor-
mation and longitudinal stress of the slab track structure, but only wind speeds
>6 m/s affected thermal transfer in the track structure.
Thermal Transfer Effects of CRTS II Slab Track Under Various … 15

Acknowledgements The authors gratefully acknowledge support for this research provided
by National Natural Science Foundation of China (No.52278311), the National Key Tech-
nologies Research and Development Program (No.2022YFB2603300), Guangdong Province
Natural Science Foundation (No. 2022A1515010665), the Shenzhen Science and Technology
Program under grant (Nos. KQTD20180412181337494, and GJHZ20200731095802007), the
Project of Science and technology research and development of China Railway Co., Ltd. (No.
K2022G038) and the Open Project of the State Key Laboratory of High-speed Railway Track Tech-
nology (No. 2021YJ143) and State Key Laboratory of Mountain Bridge and Tunnel Engineering
(No. SKLBT-ZD2101).

References

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mountainous areas. Struct Eng Int 28(3):288–295
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16 R. Zhou et al.

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Investigation on Superhydrophobicity
and Piezoresistivity of Self-sensing
Cement-Based Sensors Using Silane
Surface Treatment

W. K. Dong, W. G. Li, X. Q. Lin, and S. P. Shah

Abstract Cement-based sensors are highly susceptible to the effects of watery envi-
ronments due to the hydrophilic properties of the cement matrix. In this paper, we
applied a surface treatment using a silane/isopropanol solution to graphene/cement-
based sensors to achieve superhydrophobicity and mitigate piezoresistive instability
in watery environments. After treatment, impressive water contact angles of 163.4°
and 142.0° were achieved for the surface and inner cement-based sensors, respec-
tively. Moreover, the piezoresistivity of the coated cement-based sensors exhibited
greater stability compared to their untreated counterparts. These results provide
valuable insights into the piezoresistivity of hydrophobic cement-based sensors
in moist environments, offering promising prospects for future structural health
monitoring applications.

Keywords Cement-based sensors · Graphene nanoplates · Piezoresistivity ·

Silicone hydrophobic powder

1 Introduction

Piezoresistivity-based self-sensing cementitious composites have attracted

increasing attention recently, for their ability to achieve structural and pave-
ment health monitoring automatically. Cement-based sensors have been widely

W. K. Dong · W. G. Li (B) · X. Q. Lin

School of Civil and Environmental Engineering, University of Technology Sydney, Sydney,
NSW, Australia
e-mail: wengui.li@uts.edu.au
W. K. Dong
Institute of Construction Materials, Technische Universität Dresden, Dresden, Germany
S. P. Shah
Center for Advanced Construction Materials, The University of Texas at Arlington, Arlington,
TX, USA

© The Author(s) 2023 17

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_3
18 W. K. Dong et al.

investigated, ranging from their types and content of conductive filler, matrix, addi-
tives, curing, and drying methods to field application conditions [1, 2]. However,
due to the hydrophilic and porous structure of the cement matrix, the electrical
conductivity and piezoresistivity of cement-based sensors can be easily affected
by the working environment, especially watery, and humid conditions [3, 4].
Previous studies have attempted to remove the influence of penetrated water on the
piezoresistivity of cement-based sensors. The water absorption of cement-based
sensors was significantly reduced in the early age, but the efficiency was relatively
low, with unstable piezoresistivity in the long term [5].
Basically, using waterproofing materials to treat the surface of cement-based
sensors can prevent water penetrating the cement matrix [6, 7], which can reduce
the interference of water molecules on electrical resistivity and piezoresistivity. In
this study, we propose a special surface treatment of graphene/cement-based sensors
by immersing the sensors in a silane/isopropanol solution. The hydrophobic silane
is expected to penetrate the cement-based sensors and improve the waterproofing
properties.

2 Methods

In this paper, graphene nanoplate (GNP) is used to achieve intrinsic self-sensing

ability and piezoresistivity of cement-based sensor. The GNP is commercially avail-
able, and its specific properties are documented [8]. The raw materials consisted
of general-purpose cement, silica fume, superplasticizer, GNP, silane hydrophobic
powder (SHP), and tap water. The addition of SHP can enhance the waterproofing
and hydrophobic behaviors of cement-based sensors. The silane used was aqueous
trichlorosilane for surface modification, and the solvent isopropanol was chosen
to disperse the silane and control the concentration. The physical and chemical
properties of silane and isopropanol are listed in Table 1.
To remove surface impurities and smooth the surface, the cement-based sensors
were firstly polished and cleaned using sandpaper before coating. The surface modi-
fication followed the steps shown in Fig. 1. Isopropanol was prepared in a measuring
beaker, and 4% silane by volume of isopropanol was added, followed by 5 min of

Table 1 Physical and chemical properties of silane and isopropanol

Product Appearance Color Relative Formula Molecular
density weight
Trichloro (1H, 1H, Clear liquid Colorless 1.3 g/cm3 C8 H4 Cl3 F13 Si 481.54 g/
2H, mol
2H-perfluorooctyl)
silane
2-Propanol Clear liquid Colorless 0.785 g/ C3 H8 O 60.10 g/mol
mL
Investigation on Superhydrophobicity and Piezoresistivity … 19

Fig. 1 a–e Treatment procedure of cement-based sensors

mechanical mixing to dissolve and disperse the silane. The cement-based sensors
were placed above a copper mesh in a plastic container at a distance of 5.0 mm,
so all surfaces had continual contact with the silane/isopropanol or silane solution.
The mixed solution was gently poured into the container until the top surfaces of
cement-based sensors were just covered. It should be noted that the electrodes of
the cement-based sensors were not immersed in the silane/isopropanol solution, to
ensure excellent conductivity of the electrodes. The container was sealed with plastic
film to avoid volatilization of isopropanol. The cement-based sensors were immersed
for 2 h to ensure the thorough entrance of the solution into the cracks and pores of
the cementitious material. Finally, the cement-based sensors were dried in an oven
at 50 °C for 4 h to volatilize the isopropanol.
The water contact angle (CA) measurements of the cement-based sensors before
and after surface modification were performed with an optical tensiometer (Attension
Theta). The test liquid was deionized water with a volume of 0.2 µL for each water
drop. The water CA of the intact surface of the cement-based sensors represents
the hydrophobic coating efficiency. To obtain the hydrophobic behavior of the inner
sensor, the CA tests were also performed on cross-sections of the cement-based
sensors.

3 Results and Discussion

3.1 Hydrophobic Behavior

Figure 2 shows the surface water CA of the cement-based sensors before and after
surface modification at the time of 0, 1, 5, and 9 s from water dropping to stabilization.
The cement-based sensors without a coating shown in Fig. 3a exhibited hydrophilic
behavior with an initial CA of 79.2°. Subsequently, the water CA gradually decreased
over time until the smallest value of 70.1°, which implied that the water molecules
penetrated the cement-based sensors due to the hydrophilic behavior and porous
structure of the cement matrix. Consequently, the altered water content would be
able to permanently affect the electrical and piezoresistive properties of the cement-
based sensors, which indicates the necessity to coat them. For the coated cement-
based sensors, it was observed that they exhibited hydrophobic behavior, with a
final CA of 163.4°. In addition, the cement-based sensors without a coating showed
20 W. K. Dong et al.

Fig. 2 Water contact angles (CAs) of cement-based sensors a without and b with surface treatment

Fig. 3 Water contact angles

(CAs) of the interior of
cement-based sensors:
a without coating and b with
coating

decreasing CA, whereas their coated counterpart only showed a slight fluctuation
rather than continual decline.
Figure 3a, b shows the water CAs the cross-sectional surface of the cement-
based sensors before and after treatment, to display the hydrophobic or hydrophilic
behavior of the interior of the sensors. The cement-based sensor without a coating
displayed hydrophilic behavior and a similar CA of 71.2° to that of surface. In
contrast, the interior of the sensor became hydrophobic with a CA value of 142.0°,
which demonstrated that the silane/isopropanol solution could penetrate into the core
of cement-based sensors through micropores and cracks, resulting in hydrophobicity
of the cut cross-section.

3.2 Piezoresistivity

The stress-sensing performance of the cement-based sensors before and after surface
modification is shown in Fig. 4. Fractional changes of resistivity (FCR) of the cement-
based sensors exhibited an excellent relationship to compressive stress, with first a
Investigation on Superhydrophobicity and Piezoresistivity … 21

8 15 0
Compressive stress
Compressive stress (MPa)
6
10 -2
4
5 -4
2

FCR (%)
FCR (%)
0 0 -6

-2
-5 -8
-4
-10 -10 Without coating
-6 Without coating After coating
After coating
-8 -15 -12
0 5 10 15 20 25 30 35 40 45 50 55 60 65 0 -100 -200 -300 -400 -500 -600
Time (s) Strain (10-6)
(a) Stress self-sensing (b) Strain self-sensing

Fig. 4 a Stress- and b strain-sensing capacities of cement-based sensors before and after surface
modification. FCR, fractional change of resistivity

decrease and then returned resistivity in the loading and unloading processes. This
finding demonstrated that the silane-based surface modification did not eliminate
the piezoresistivity of the cement-based sensors. The graphene-filled cement-based
sensor without a coating showed the highest FCR value of 12.6%, followed by an
average FCR of 7.2% for the cement-based sensors after surface treatment. These data
implied that the stress-sensing efficiency might be weakened by the silane modifica-
tion. In addition, small fluctuations can be seen for the silane-coated cement-based
sensors, mainly due to the brittleness and heterogeneity of the cementitious materials,
which led to sudden changes of electrical resistivity. Secondly, the intruded silane
aggravated the fluctuation because of its poor electrical conductivity. For the strain-
sensing performance, the FCR showed an excellent relationship with compressive
strain and showed a similar changing mode to compressive stress.

4 Conclusions

Cement-based sensors can easily absorb water molecules because of their porous
structure and hydrophilic behavior. In this study, a silane-based surface modification
was applied to improve the waterproofing and superhydrophobic behavior, while
maintaining excellent piezoresistivity of the cement-based sensors. The final water
CA significantly increased to 163.4°, and the piezoresistivity was relatively well
maintained. The piezoresistivity of the cement-based sensors seemed to decrease
after the surface modification, with slightly poorer linearity and repeatability, lower
gauge factor, and higher hysteresis. Despite this, the cement-based sensors exhibited
acceptable linearity and repeatability.
22 W. K. Dong et al.

References

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and perspective. Constr Build Mater 203:146–163
2. Han B, Yu X, Kwon E (2009) A self-sensing carbon nanotube/cement composite for traffic
monitoring. Nanotechnology 20(44):445501
3. Dong W, Li W, Lu N, Qu F, Vessalas K, Sheng D (2019) Piezoresistive behaviours of cement-
based sensor with carbon black subjected to various temperature and water content. Compos
Part B 178:107488
4. del Moral B, Baeza FJ, Navarro R et al (2021) Temperature and humidity influence on the strain
sensing performance of hybrid carbon nanotubes and graphite cement composites. Constr Build
Mater 284:122786
5. Dong W, Li W, Zhu X, Sheng D, Shah SP (2021) Multifunctional cementitious composites with
integrated self-sensing and hydrophobic capacities toward smart structural health monitoring.
Cem Concr Compos 118:103962
6. Zhou Z, Li S, Cao J, Chen X et al (2021) The waterproofing effect and mechanism of graphene
oxide/silane composite emulsion on cement-based materials under compressive stress. Constr
Build Mater 308:124945
7. Li F, Yang Y, Tao M et al (2019) A cement paste–tail sealant interface modified with a silane
coupling agent for enhancing waterproofing performance in a concrete lining system. RSC Adv
9(13):7165–7175
8. Dong W, Li W, Wang K, Shah SP (2021) Physicochemical and Piezoresistive properties of
smart cementitious composites with graphene nanoplates and graphite plates. Constr Build
Mater 286:122943

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Use of Brown Coal Ash as a Replacement
of Cement in Concrete Masonry Bricks

D. W. Law, C. Gunasekara, and S. Setunge

Abstract Portland cement production is not regarded as environmentally friendly,

because of its associated high carbon emissions, which are responsible for 5% of
global emissions. An alternative is to substitute fly ash for Portland cement. Australia
has an abundance of brown coal fly ash, as it is the main source of primary energy
in the State of Victoria. Currently, the majority of this material is stored in landfills
and currently there is no commercial use for it in the cement industry because brown
coal fly ash cannot be used as a direct replacement material for Portland cement due
to the high sulfur and calcium content and low aluminosilicate content. However,
the potential exists to use brown coal fly ash as a geopolymeric material, but there
remains a significant amount of research needed to be conducted. One possible
application is the production of geopolymer concrete bricks. A research project was
undertaken to investigate the use of brown coal fly ash from Latrobe Valley power
stations in the manufacture of geopolymer masonry bricks. The research developed
a detailed understanding of the fundamental chemistry behind the activation of the
brown coal fly ash and the reaction mechanisms involved to enable the development
of brown coal fly ash geopolymer concrete bricks. The research identified suitable
manufacturing techniques to investigate relationships between compressive strength
and processing parameters and to understand the reaction kinetics and microstruc-
tural developments. The first phase of the research determined the physical, chem-
ical, and mineralogical properties of the Loy Yang and Yallourn fly ash samples to
produce a 100% fly ash-based geopolymer mortar. Optimization of the Loy Yang
and Yallourn geopolymer mortars was conducted to identify the chemical properties
that were influential in the production of satisfactory geopolymer strength. The Loy
Yang mortars were able to produce characteristic compressive strengths acceptable
in load-bearing bricks (15 MPa), whereas the Yallourn mortars produced character-
istic compressive strengths only acceptable as non-load-bearing bricks (5 MPa). The
second phase of the research transposed the optimal geopolymer mortar mix designs
into optimal geopolymer concrete mix designs while merging the mix design with
the optimal Adbri Masonry (commercial partner) concrete brick mix design. The

D. W. Law (B) · C. Gunasekara · S. Setunge

School of Engineering, RMIT University, Melbourne, VIC, Australia
e-mail: david.law@rmit.edu.au

© The Author(s) 2023 23

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_4
24 D. W. Law et al.

reference mix designs allowed for optimization of both the Loy Yang and Yallourn
geopolymer concrete mix designs, with the Loy Yang mix requiring increased water
content because the original mix design was deemed to be too dry. The key factors
that influenced the compressive strength of the geopolymer mortars and concrete
were identified. The amorphous content was considered a vital aspect during the
initial reaction process of the fly ash geopolymers. The amount of unburnt carbon
content contained in the fly ash can hinder the reactive process, and ultimately, the
compressive strength because unburnt carbon can absorb the activating solution, thus
reducing the particle to liquid interaction ratio in conjunction with lowering worka-
bility. Also, fly ash with a higher surface area showed lower flowability than fly ash
with a smaller surface area. It was identified that higher quantity of fly ash particles
<45 microns increased reactivity whereas primarily angular-shaped fly ash suffered
from reduced workability. The optimal range of workability lay between the 110–
150 mm slump, which corresponded with higher strength displayed for each respec-
tive precursor fly ash. Higher quantities of aluminum incorporated into the silicate
matrix during the reaction process led to improved compressive strengths, illustrated
by the formation of reactive aluminosilicate bonds in the range of 800–1000 cm–1
after geopolymerization, which is evidence of a high degree of reaction. In addition,
a more negative fly ash zeta potential of the ash was identified as improving the
initial deprotonation and overall reactivity of the geopolymer, whereas a less nega-
tive zeta potential of the mortar led to increased agglomeration and improved gel
development. Following geopolymerization, increases in the quantity of quartz and
decreases in moganite correlated with improved compressive strength of the geopoly-
mers. Overall, Loy Yang geopolymers performed better, primarily due to the higher
aluminosilicate content than its Yallourn counterpart. The final step was to transition
the optimal geopolymer concrete mix designs to producing commercially accept-
able bricks. The results showed that the structural integrity of the specimens was
reduced in larger batches, indicating that reactivity was reduced, as was compressive
strength. It was identified that there was a relationship between heat transfer, curing
regimen and structural integrity in a large-volume geopolymer brick application.
Geopolymer bricks were successfully produced from the Loy Yang fly ash, which
achieved 15 MPa, suitable for application as a structural brick. Further research is
required to understand the relationship between the properties of the fly ash, mixing
parameters, curing procedures and the overall process of brown coal geopolymer
concrete brick application. In particular, optimizing the production process with
regard to reducing the curing temperature to ≤80 °C from the current 120 °C and the
use of a one-part solid activator to replace the current liquid activator combination
of sodium hydroxide and sodium silicate.

Keywords Bricks · Brown coal fly ash · Concrete · Geopolymer

.
Use of Brown Coal Ash as a Replacement of Cement in Concrete … 25

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Composition of Alkali–Silica Reaction
Products in Laboratory and Field
Concrete

M. J. Tapas, K. Vessalas, P. Thomas, N. Gowripalan, and V. Sirivivatnanon

Abstract This study investigated the composition of alkali–silica reaction (ASR)

products formed in mortar and concrete that underwent accelerated ASR testing using
two test methods: the accelerated mortar bar test (AMBT) and the simulated pore
solution immersion test (SPSM). The composition of the ASR products formed in the
accelerated tests was compared with those in a 25-year old bridge in New South Wales
demolished due to ASR. Results showed that the ASR products inside an aggregate
contained calcium (≈20%), silicon (≈60%), and alkalis (≈20%) regardless of the
ASR test method used. The ASR products in the AMBT sample only contained
sodium, whereas the ASR products in the SPSM test and the demolished bridge both
contained significant amounts of sodium and potassium, which indicated that the
type of alkali in the ASR product is largely affected by the dominant alkali in the
pore solution. However, considering that the total alkali content (Na + K) in the ASR
products was similar regardless of the ASR test method used, this suggests that the
total alkali content has more influence on the rate of expansion than the type of alkali.
The composition of the ASR products also notably varied depending on the location
in the concrete. ASR products closer to the cement paste had a higher calcium and
lower alkali content than those inside an aggregate, which suggests that the calcium
as well as the alkali content of the ASR products plays a significant role in the degree
of ASR expansion.

Keywords Accelerated mortar bar test · Alkali–silica reaction (ASR) · ASR

products

M. J. Tapas (B) · K. Vessalas · P. Thomas · N. Gowripalan · V. Sirivivatnanon

School of Civil and Environmental Engineering, University of Technology Sydney, Sydney,
NSW, Australia
e-mail: mariejoshua.tapas@uts.edu.au

© The Author(s) 2023 27

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_5
28 M. J. Tapas et al.

1 Introduction

The dissolution of the reactive silica components of an aggregate can result in a

concrete durability issue known as the alkali–silica reaction (ASR). The reactive
silica components of an aggregate, which dissolve when the concrete pore solution’s
alkali concentration is sufficiently high, can react with calcium and alkali ions in
the pore solution to form the ASR products that lead to expansion and deleterious
cracking of the concrete [1]. Reactive silica minerals are those that are amorphous,
have poor crystallinity, highly strained, or contain high amount of defects making
them susceptible to alkali dissolution.
Two test methods are widely used to assess ASR potential: the accelerated mortar
bar test (AMBT) and the concrete prism test (CPT). AMBT AS1141.60.1 makes
use of 1 M NaOH at 80 °C to accelerate the reaction and an expansion of <0.1%
at 21 days indicates that the aggregate is nonreactive [2]. CPT AS1141.60.2, on the
other hand, involves boosting the cement alkali content to 1.25% Na2 Oeq and storing
the concrete prisms in a sealed, high humidity environment at 38 °C. The CPT is a
longer test method than the AMBT and has an expansion limit of 0.03% at 1 year
for an aggregate to be considered nonreactive [3]. The Australian AMBT and CPT
methods are based on ASTM C1260 and ASTM C1293 respectively, with slightly
modified test limits.
Both the AMBT and CPT remain widely used although heavily criticized owing to
their limitations. The use of high temperature as well as excessive alkali supply in the
AMBT can result in false positives (i.e., identifying an aggregate as reactive even if it
is not). Moreover, the influence of cement alkalinity, which significantly contributes
to the ASR, is hard to determine [4]. CPT on the other hand is prone to alkali leaching,
which can lead to underestimation of expansion and is reportedly ≈25–35% in 1 year
[5, 6]. Because of the limitations of these test methods, an alternative ASR test, the
simulated pore solution method (SPSM), has been developed at the Laboratory of
Construction Materials, EPFL Switzerland. This test method involves immersing the
concrete prisms in a solution based on the pore solution’s alkali concentration for the
concrete being assessed for ASR potential and so far has shown promising potential
as an alternative ASR test method [7, 8].
Due to the availability of various test methods for assessing the risk of an aggregate
for ASR, it is therefore important to assess the effect of these test methods on the
type of ASR products. Moreover, the comparison of the composition of the ASR
products formed using these methods to those in actual structures affected by ASR is
also important to determine whether ASR products formed during accelerated testing
resemble those in the field. Therefore, we investigated the composition of the ASR
products in both the AMBT, and SPSM and compared them to ASR products in a
25-year-old bridge in Australia, decommissioned because of ASR.
Composition of Alkali–Silica Reaction Products in Laboratory … 29

2 Methods

2.1 ASR Testing as Per AMBT and SPSM

Expansion tests using AMBT and SPSM were carried out using Australian reactive
aggregates and Australian cement. The cement complied with the 0.6% Na2 Oeq alkali
limit.
For the AMBT, mortar bars composed of 1 part cement to 2.25 parts graded
aggregate by mass (440 g cement per 990 g of aggregate) and a water to cementitious
materials ratio equal to 0.47 were prepared in accordance with AS1141.60.1. The
mortar specimens were prepared in 25 × 25 × 285 mm molds with a gauge length of
250 mm, then cured in a high humidity environment at room temperature (23 ± 2 °C)
for 24 h. Next, the specimens were demolded and placed in a water-filled container,
before being placed in an oven at 80 °C for another 24 h to allow the specimens
to slowly equilibrate to 80 °C. Horizontal comparator was used to obtain zero-hour
length measurements before immersing the specimens in 1 M NaOH solution at
80 °C for 28 days. Succeeding expansion measurements were obtained at 1, 3, 7, 10,
14, 21, and 28 days. Three readings were taken per mortar specimen at each age.
Total expansion incurred by the aggregate after 10 and 21 days of NaOH immersion
was used to classify its ASR potential when used in the field in accordance with
AS1141.60.1.
For the SPSM, concrete prisms (70 × 70 × 280 mm) were prepared with a cement
content of 410 kg/m3 and water-to-cement ratio of 0.46. The concretes were cured
for 28 days in a high humidity chamber (>90% relative humidity) before storage in
simulated pore solution at 60 °C. Expansion measurements were taken before storage
and every month thereafter using a vertical comparator. The simulated pore solution
was prepared based on the extracted pore solution of an equivalent binder system at
28 days.

2.2 Analysis of the Composition of the ASR Products

Polished sections of mortar/concrete that underwent expansion tests using the AMBT,
and SPSM, as well as the concrete from the demolished bridge were prepared and
subjected to scanning electron microscopy–energy-dispersive spectroscopy (SEM–
EDS) analysis. The AMBT samples were sectioned after 28 days in the AMBT
bath while the concrete prisms were sectioned after 6 months in the simulated pore
solution at 60 °C.
The mortars and concretes were cut to fit a 25-mm diameter mold, then vacuum
impregnated with epoxy resin and polished, first with silicon carbide paper until
the sample surface had been fully uncovered from the resin, followed by automated
polishing using MD Largo Struers discs lubricated with petrol and diamond spray
as the polishing agent (9 µm, 3 µm and 1 µm particle sizes). After polishing, the
30 M. J. Tapas et al.

samples were cleaned in an ultrasonic bath for 2 min and then stored in a vacuum
desiccator for at least 2 days to dry. The samples were coated with carbon to prevent
charging during SEM imaging.
Imaging and elemental analysis of the carbon-coated polished sections were
carried out using an FEI Quanta 200 with Bruker XFlash 4030 EDS detector. The
microscope was operated in backscattered electron (BSE) mode, 15 kV accelerating
voltage and 12.5 mm working distance in a high vacuum.

3 Results and Discussion

The AMBT and SPSM expansion plots are shown in Fig. 1, confirming the high
reactivity of the aggregates as indicated by the significant degree of expansion. The
AMBT mortars (dacite and greywacke) both exceeded the 0.1% limit at 10 days and
0.3% at 21 days, making them reactive as per AS1141.60.1. There is currently no
established limit for the SPSM but the high degree of expansion of the dacite concrete
confirmed the reactivity of the aggregate.
Figure 2 presents images of the 25-year-old bridge before it was decommissioned,
showing extensive damage due to ASR. The cracks observed have a map crack
appearance, which is characteristic of ASR [9].
Figure 3 shows the ASR damage observed in the mortar and concrete that under-
went AMBT and SPSM respectively. In both cases, the cracks were concentrated
within the aggregate and extend towards the paste, which indicated that deleterious
ASR damage originated within the aggregate and explains the characteristic map
crack appearance of ASR damage in affected structures. Figure 4 shows the ASR
products observed in the AMBT sample, and Fig. 5 shows the ASR products observed
in the concrete sample subjected to SPSM with their corresponding EDS maps.
For both cases, the presence of calcium, silicon and alkalis are notable confirming
the composition of the ASR product (alkali-calcium silicate hydrate). It is however

0.60%
0.20% b
Greywacke mortar a Dacite Concrete
0.50%
Dacite mortar 0.15%
0.40%
Expansion

Expansion

0.30% 0.10%
0.20%
0.05%
0.10%

0.00% 0.00%
0 5 10 15 20 25 30 0 1 2 3 4 5 6
Days -0.05% Months

Fig. 1 Expansion plots of the aggregates subjected to a accelerated mortar bar test and b simulated
pore solution immersion test
Composition of Alkali–Silica Reaction Products in Laboratory … 31

crack

Fig. 2 A 25-year-old bridge in New South Wales, Australia, suffering from alkali–silica reaction

a paste
aggregate b
paste
aggregate
aggregate

paste
aggregate
paste

paste
paste aggregate
aggregate

Fig. 3 Scanning electron microscopy images showing extensive damage in the mortar and concrete
after a accelerated mortar bar test and b simulated pore solution immersion test

notable that whereas the alkali present in the AMBT is only sodium, the ASR product
in the concrete prism has both sodium and potassium. This indicates that the type
of alkali in the ASR product is strongly affected by the dominant alkalis in the pore
solution.
Figure 6 shows the ASR products forming around the aggregate and in the
paste, which has a notably darker color than the ASR products inside the aggre-
gate, suggesting a difference in composition. Table 1 tabulates the EDS results of the
ASR products inside the aggregate (SPSM sample) and the ASR products around
the aggregate and near the paste (SPSM sample). As can be seen, the ASR products
outside the aggregate have a composition closer to C-S–H and a much higher Ca/Si
ratio and lower Na + K/Si than the ASR products inside the aggregate. In general,
the silicon content of the ASR product decreased and calcium content increased
as the product came in closer contact with the cement paste [10–14]. The role of
calcium, however, remains controversial. Although higher calcium content in the
ASR product results in higher stiffness [15], as the ASR product becomes more
rigid, it also has decreased swelling potential [1, 16]. The substitution of alkalis with
calcium suggests there is a competitive reaction between calcium and alkalis and
32 M. J. Tapas et al.

ASR product calcium silicon

aggregate

sodium potassium
No potassium

Fig. 4 Alkali–silica reaction (ASR) product observed in the accelerated mortar bar tested mortar
with EDS maps showing strong presence of calcium (Ca), silicon (Si) and sodium (Na). BSE,
backscattered electron mode; EDS, energy-dispersive spectroscopy

ASR product calcium silicon

aggregate
paste

sodium potassium

Fig. 5 Alkali–silica reaction (ASR) product observed in the concrete subjected to the simulated
pore solution method (SPSM) sample with EDS maps showing strong presence of calcium (Ca),
silicon (Si), sodium (Na) and potassium (K). BSE, backscattered electron mode; EDS, energy-
dispersive spectroscopy

that calcium is always preferentially absorbed, which supports the alkali recycling
theory [1].
Table 2 tabulates the EDS results of the AMBT sample (inside the aggregate).
As can be observed, consistent with the EDS mapping results, almost no potassium
can be detected in the ASR products. The total alkali content (Na + K) is, however,
comparable to the SPSM samples (≈20%), as well as the Ca/Si and (Na + K)/Si.
This indicates that the composition of the ASR products inside an aggregate has a
Composition of Alkali–Silica Reaction Products in Laboratory … 33

ASR product
paste

ASR product
paste

aggregate aggregate

Fig. 6 Alkali–silica reaction (ASR) products observed in the simulated pore solution method
sample surrounding the aggregate and located in the cement paste

Table 1 Energy-dispersive spectroscopy results for the Alkali–silica reaction (ASR) product
observed in the simulated pore solution method (SPSM)sample (normalized without oxygen)
SPSM Location Ca Si Al Na K Na + K Ca/Si (Na + K)/Si
sample
ASR Inside 18.03 59.04 1.14 8.50 13.29 21.79 0.31 0.37
product 1 aggregate
ASR Inside 20.00 58.31 0.81 6.54 14.34 20.88 0.34 0.36
product 2 aggregate
ASR Near/in the 46.51 40.96 6.02 3.37 3.13 6.51 1.14 0.16
product 3 paste
ASR Near/in the 49.88 39.76 6.27 2.89 1.20 4.10 1.25 0.10
product 4 paste

Table 2 Energy-dispersive spectroscopy results for alkali–silica reaction (ASR) products observed
in the accelerated mortar bar test (AMBT) sample (normalized without oxygen)
AMBT Location Ca Si Al Na K Na + K Ca/Si (Na + K)/Si
sample
ASR Inside 19.29 61.46 0.70 17.41 1.13 18.55 0.31 0.30
product 1 aggregate
ASR Inside 18.88 62.86 0.79 16.19 1.28 17.47 0.30 0.28
product 2 aggregate

similar stoichiometric ratio of calcium, silicon and alkali regardless of the ASR test
method. The type of alkali, however, varies depending on the dominant alkali/s in
the pore solution.
Figure 7 shows the ASR product in the demolished concrete bridge with corre-
sponding EDS maps. As can be observed, the ASR product was also concentrated
inside the aggregate and also contained calcium, silicon and alkali similar to the ASR
products in the AMBT and SPSM samples. The ASR product in the bridge, however,
34 M. J. Tapas et al.

calcium silicon

ASR product

sodium potassium

Fig. 7 Alkali–silica reaction (ASR) product in the demolished bridge (NSW, Australia) with
energy-dispersive spectroscopy mapping

showed the presence of both sodium and potassium and hence its composition was
closer to the SPSM sample than the AMBT sample.

4 Conclusions

This study showed a difference in composition between ASR products found in a

mortar subjected to AMBT, concrete subjected to SPSM and a 25-year-old bridge
demolished because of ASR. Our results showed the following.
1. The ASR products in the AMBT, SPSM and demolished bridge samples all
contained calcium, silicon and alkali. The ASR products in the AMBT sample,
however, only contained sodium whereas those in the SPSM and bridge samples
contained both sodium and potassium, which indicated that the type of alkali
in the ASR product is significantly affected by the dominant alkali in the pore
solution.
2. Although the type of alkali in the ASR product was affected by the test method,
the total Na + K in the ASR products found inside the aggregate remained the
same regardless of the test method, which suggests that the total alkali plays a
more significant role in ASR expansion than the type of alkali.
3. The composition of the ASR products varied according to the location in the
concrete. ASR products within the aggregate had a lower Si/Ca ratio and higher
(Na + K)/Si ratio than ASR products in the cement paste and those surrounding
the aggregates. Thus, the closer the ASR product is to the cement paste, the higher
the calcium content and the lower the alkali content, which confirms the theory
of alkali recycling where Ca2+ substitutes for alkali. This finding also suggests
Composition of Alkali–Silica Reaction Products in Laboratory … 35

that the composition of the ASR products affects the rate of expansion and that
ASR products with lower Si/Ca ratio and higher alkali content (observed inside
the aggregate) may be more deleterious.

Acknowledgements This study was part of University of Technology Sydney research funded
through the Australian Research Council Research Hub for Nanoscience-Based Construction
Materials Manufacturing (NANOCOMM) with the support of Cement Concrete and Aggregates
Australia (CCAA) and the Australian Government Research Training Program Scholarship. This
work would also not have been possible without laboratory equipment provided by the Laboratory
of Construction Materials at EPFL Switzerland, courtesy of Professor Karen Scrivener.

References

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current understanding of the reaction mechanisms and the knowledge gaps. Cem Concr Res
76:130–146
2. Standards Australia, AS 1141.60.1 Methods for sampling and testing aggregates method 60.1:
potential alkali-silica reactivity-accelerated mortar bar method, standards Australia Limited,
Sydney, Australia, 2014
3. Standards Australia, AS 1141.60.2 Methods for sampling and testing aggregates method 60.2:
potential alkali-silica reactivity-concrete prism method, standards Australia Limited, Sydney,
Australia, 2014
4. Islam MS, Alam MS, Ghafoori N, Sadiq R (2016) Role of Solution Concentration, cement alkali
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concrete tested for alkali-silica reaction. Mater Struct 40:909–921
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preventive measures for controlling expansion due to alkali–silica reaction in concrete. Cem
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7. Chappex T, Sofia L, Dunant C, Scrivener K (2016) A robust testing protocol for the assessment
of ASR reactivity of concrete. 15th international conference on alkali aggregate reaction in
concrete (ICAAR) São Paulo Brazil
8. Tapas MJ, Sofia L, Vessalas K, Thomas P, Sirivivatnanon V, Scrivener K (2021) Efficacy of
SCMs to mitigate ASR in systems with higher alkali contents assessed by pore solution method.
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identification handbook, federal highway administration Washington, DC
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concrete structures. Mater Charact 60:655–668
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uptake property of alkali-silica reaction products. Cem Concr Res 121:58–71
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concrete railroad ties: SEM-EDX and X-Ray diffraction analyses. Cem Concr Res 26:309–318
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14. Scrivener KL, Monteiro PJM (1994) The alkali-silica reaction in a monolithic opal. J Am
Ceram Soc 77:2849–2856
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Cem Concr Res 34:1389–1402

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
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Commons license, unless indicated otherwise in a credit line to the material. If material is not
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the copyright holder.
Behavior of Hybrid Engineered
Cementitious Composites Containing
Nanocellulose

H. Withana and Y. X. Zhang

Abstract Nanocellulose (NC) is a promising reinforcing material for cementitious

composites, but its effect on the mechanical properties of hybrid engineered cemen-
titious composites (ECCs) has not been studied. In this paper, we investigated a
hybrid polyethylene (PE)-steel fibre ECC reinforced with NC and the effect of NC
dosages ranging from 0.1% to 0.4% on the compressive strength of the hybrid ECC.
The optimum quantity of NC for the best mechanical property of ECC was deter-
mined. Enhancement of compressive strength was observed for all the mixes with
NC compared with the reference mix, and the mix containing 0.2% NC showed the
maximum improvement.

Keywords Engineered cementitious composites · Nanocellulose · Polyethylene

fibers · Steel fibers

1 Introduction

Engineered cementitious composites (ECCs) micromechanically designed by adding

microfibers show significant improvements in the ductility, tensile strength, and
toughness of cement-based materials. Nevertheless, they have no noticeable influence
on compressive and flexural strength [1]. Microfibers such as polymer fibers, due to
their relatively small surface areas, can have limited interfacial strength. Thus, they
can pose problems in reinforced cementitious composites by entrapping air voids and
degrading the strength [2]. In this regard, reinforcement of the ECC at the nanoscale
might be an alternative option.
Research of cementitious composites reinforced with nanomaterials is rapidly
expanding and recently the growing push for greener construction materials has
shifted the attention to plant-based nanomaterials such as Nano cellulose (NC). NC

H. Withana · Y. X. Zhang (B)

School of Engineering, Design and Built Environment, Western Sydney University, Kingwood,
NSW 2747, Australia
e-mail: Sarah.Zhang@westernsydney.edu.au

© The Author(s) 2023 37

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_6
38 H. Withana and Y. X. Zhang

is a nanofiber derived from cellulose, which is the primary component of plant

cell walls. Being an abundant natural ingredient, it is one of the most sustainable
raw materials [3]. In addition to its environmental benefits, including renewability,
biodegradability, low environmental impact, and low health risks associated with
production, it also has a low production cost. NC also has outstanding mechanical
and physical benefits such as high modulus and strength, high specific and reactive
surface, high aspect ratio and hydrophilic and hygroscopic properties, all of which
are favorable for use as a reinforcement for cementitious composites [4].
NC-based composites have been successfully used in areas such as the medical,
food, paper, and electrochemical industries [5]. However, their application in the
construction industry is limited. Although improvement of the mechanical charac-
teristics of NC-based composites has been established in the literature, research into
their use in cementitious composites, notably ECC, has been minimal and none for
hybrid ECC.
In this study, a novel NC-reinforced hybrid polyethylene (PE)-steel ECC was
developed and the effect of NC on its compressive strength was investigated. A series
of uniaxial compression tests were carried out, incorporating different dosages of NC
to study the effect on the compressive properties. Compressive strength of the mixes
was compared with a reference mix without NC, and the optimum quantity of NC
that modifies the performance of ECC was determined.

2 Specimen Preparation

The raw materials used included cement, sand, water, fly ash, silica fume, super-
plasticizer, steel fibers, PE fibers, and NC. General Purpose Portland cement that
conformed to AS3972, fly ash of ASTM class F (SG 2–2.5) and high-grade silica
fume were used as the binding materials, while sand with a mean grain size of 225 µm
was used as fine aggregate. PE fibers (1.5% by volume) and steel fibers (0.5% by
volume) were used as high modulus and low modulus fibers, respectively, and cellu-
lose nanofibrils (CNF) were used as the NC. Additionally, polycarboxylate-based
high-range water-reducing admixture Rheobuild 10000N7 was used to attain good
workability with consistent rheological properties for uniform fiber dispersion.
The NC was purchased from Cellulose Lab in Canada. It was derived from
bleached softwood pulp and prepared by subjecting the pulp to intensive mechanical
treatment by a high-pressure homogenizer. NC was provided as a slurry in aqueous
gel form with 3.0 wt%.
Six NC-reinforced ECC mixes were prepared by adding NC dosages of 0.1%,
0.2%, 0.25%, 0.3% and 0.4% to the hybrid ECC. ECC with 0% NC was used as the
reference mix. The hybrid ECC used in this study was a 0.5% steel and 1.5% PE
ECC that was developed by the authors in a previous study [6]. Compressive tests
were carried out for all seven mixes including the reference mix.
To make the samples, the dry ingredients including cement, fly ash, silica fume and
sand) were mixed for about 2 min. The fibers were then slowly added, and the mixing
Behavior of Hybrid Engineered Cementitious Composites Containing … 39

was continued for a few more minutes until all the fibers were evenly distributed.
Water and superplasticizer were combined and gradually added to the dry mix while
the mixing continued. CNF was diluted in water using a hand blender and added to
the mix and the mixing was continued for 3 min. Once the mix was liquefied and
in a consistent and uniform state, it was poured into molds and vibrated for 30 s.
Following this, the molds were cling film-wrapped and kept at room temperature for
24 h until demolding. The specimens were then wrapped in plastic sheets and placed
in an oven at 23 ± 1 °C and relative humidity of 50% until the age of testing.

3 Results of Uniaxial Compression Test

Uniaxial compression tests were carried out in accordance with AS 1012. The tests
were conducted at 28 days after casting the 50-mm cubic specimens. An INSTRON
5500R machine with 1000 KN capacity was used at a loading rate of 20 MPa/min.
Three specimens from each mix were tested.
The results of the compression tests are presented in Fig. 1. All the mixes with NC
showed an improvement in compressive strength compared with the reference mix
without any NC. NC concentrations of 0.1%, 0.2%, 0.25%, 0.3% and 0.4% improved
the compressive strength by 29.1% 45.3%, 29.4%. 15.8% and 5.8%, respectively. A
threshold was reached at 0.2%, where a maximum compressive strength of 68.4 MPa
was achieved. Beyond this threshold concentration, the compressive strength started
to decrease.

Compressive Stress
80.0
68.4
70.0
Compressive Strength (MPa)

60.0 52.8 53.0

50.0 44.3
37.4 39.7
40.0

30.0

20.0

10.0

0.0
0% NC 0.1% NC 0.2% NC 0.25% NC 0.3% NC 0.4% NC

Fig. 1 Compressive strength at 28 days with different dosages of nanocellulose (NC)

40 H. Withana and Y. X. Zhang

Increased interaction between the nanofibers due to a densified matrix can be

attributed to the improvement in strength, while the reduction in strength at concen-
trations above 0.2% could be attributed to fiber agglomeration arising from difficulty
of dispersion.
An internal curing effect of NC that stems from the hydrophilic and hygroscopic
nature of NC enhances the degree of hydration (DOH) of the matrix [7–9]. When the
DOH is enhanced, porosity is reduced, and strength is increased. Moreover, the high
aspect ratio and the high specific surface area increase the cellulose hydroxyl groups
available for hydrogen bonding in the cementitious matrix [10]. Consequently, strong
bonding is promoted, resulting in enhanced fiber–matrix interaction and a densified
matrix.
On the other hand, the -OH groups at high density on the surface of cellulose fibers
try to bond with adjacent –OH groups by hydrogen bonds causing agglomeration or
entanglement of the fibers [11]. Due to the larger surface area, this effect is more
significant in nanofibers. Therefore, when NC is applied in excessive quantities,
dispersion becomes difficult and the fibers agglomerate, causing high porosity and
the fibers acting as stress concentrators, resulting in a reduction in strength.

4 Conclusions

Compressive strength increased as the NC concentration increased until a threshold

was reached at 0.2%, where the maximum compressive strength of 68.4 MPa was
achieved. Beyond this threshold concentration, the compressive strength started to
decrease. All the mixes with NC achieved enhancements of compressive strength
compared with the reference mix, with the mix containing 0.2% NC achieving the
highest improvement of 45.3%. The decline in strength at high doses of NC was
attributed to fiber agglomeration. Microscale studies will be conducted in the future
for a more insightful investigation.

References

1. Li VC (2008) Engineered cementitious composites (ECC) material, structural, and durability

performance
2. Wang S, Li VC (2004) Tailoring of pre-existing flaws in ECC matrix for saturated strain
hardening
3. Eichhorn SJ, Dufresne A, Aranguren M, Marcovich NE, Capadona JR, Rowan SJ, Weder C,
Thielemans W, Roman M, Renneckar S, Gindl W, Veigel S, Keckes J, Yano H, Abe K, Nogi
M, Nakagaito AN, Mangalam A, Simonsen J, Benight AS, Bismarck A, Berglund LA, Peijs T
(2010) Review: current international research into cellulose nanofibres and nanocomposites. J
Mater Sci 45(1):1–33. https://doi.org/10.1007/s10853-009-3874-0
4. Claramunt J, Ardanuy M, Fernandez-Carrasco LJ (2015) Wet/dry cycling durability of
cement mortar composites reinforced with micro-and nanoscale cellulose pulps. BioResources
10(2):3045–3055
Behavior of Hybrid Engineered Cementitious Composites Containing … 41

5. Abitbol T, Rivkin A, Cao Y, Nevo Y, Abraham E, Ben-Shalom T, Lapidot S, Shoseyov O (2016)

Nanocellulose, a tiny fiber with huge applications. Curr Opin Biotechnol 39:76–88. https://doi.
org/10.1016/j.copbio.2016.01.002
6. Withana H, Zhang YX, Zhong T (2022) Material properties of a new hybrid fibre-reinforced
engineered cementitious composite with high volume fly ash. Manuscript in preparation
7. Haque MI, Ashraf W, Khan RI, Shah S (2022) A comparative investigation on the effects
of nanocellulose from bacteria and plant-based sources for cementitious composites. Cement
Concr Compos 125:104316
8. Hisseine OA, Omran AF, Tagnit-Hamou A (2018) Influence of cellulose filaments on cement
paste and concrete. J Mater Civil Eng 30(6). https://doi.org/10.1061/(asce)mt.1943-5533.000
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9. Onuaguluchi O, Panesar DK, Sain M (2014) Properties of nanofibre reinforced cement
composites. Constr Build Mater 63:119–124. https://doi.org/10.1016/j.conbuildmat.2014.
04.072
10. Cao Y, Zavaterri P, Youngblood J, Moon R, Weiss J (2015) The influence of cellulose
nanocrystal additions on the performance of cement paste. Cement Concr Compos 56:73–83
11. Wang B, Sain M (2007) Isolation of nanofibers from soybean source and their reinforcing
capability on synthetic polymers. Compos Sci Technol 67(11–12):2521–2527

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Investigation of ASR Effects
on the Load-Carrying Capacity
of Reinforced Concrete Elements
by Ultra-Accelerated Laboratory Test

J. Cao, N. Gowripalan, V. Sirivivatnanon, and J. Nairn

Abstract The alkali–silica reaction (ASR) can cause expansion, cracking, and
degradation of the mechanical properties of affected concrete. Concerns about the
safety of ASR-damaged reinforced concrete structures have driven the demand for
studying the effects of ASR on residual load capacity of the deteriorated struc-
ture. Conventionally, field load testing methods are used to assess the residual load
capacity of ASR-affected structures. In this study, a novel accelerated laboratory
test using the LVSA 50/70 autoclave to accelerate ASR was applied to investigate
the flexural and shear behavior of small-scale reinforced concrete beams affected by
ASR. The specimens were subjected to three cycles of 80 °C steam curing at atmo-
spheric pressure in the autoclave, with 60 h/cycle. Significant expansion and ASR
damage were observed. Load carrying capacity tests on the small-scale reinforced
concrete beams showed that, at the expansion levels achieved, the flexural capacity
of the reinforced concrete beams was not significantly affected. Shear resistance of
the reinforced concrete beams, however, was found to increase compared with their
28-day counterparts, which could be attributed to the prestressing effect due to ASR
expansion. It appears that the multicycle 80 °C steam-curing autoclave test is suitable
for investigating ASR deterioration of actual concrete mixes within a short period of
time. ASR effects on the load carrying capacity of reinforced concrete elements at
higher expansion levels, however, need further investigation.

Keywords Alkali–silica reaction (ASR) · Accelerated ASR test · Expansion ·

Load-carrying capacity

J. Cao (B) · N. Gowripalan · V. Sirivivatnanon

School of Civil and Environmental Engineering, University of Technology Sydney, Ultimo,
NSW 2007, Australia
e-mail: Jinsong.cao@uts.edu.au
J. Nairn
Cement Concrete and Aggregates Australia, Sydeny, NSW 2020, Australia

© The Author(s) 2023 43

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_7
44 J. Cao et al.

1 Introduction

The alkali–silica reaction (ASR) is one of the major durability problems for concrete
structures and has been observed and studied for decades worldwide. Concerns about
the safety of ASR-damaged reinforced concrete structures have driven the demand
for studying the effects of ASR on the performance of the structure and the effect of
ASR on residual load capacity of the deteriorated structure [1].
During the past decades of extensive research on ASR, field load testing on real
structures was used to assess the residual load capacity of ASR-affected structures
[2]. In addition, under controlled laboratory conditions, efforts were made to inves-
tigate the flexural and shear behavior of small-scale to full-scale reinforced concrete
specimens [3–5]. Large-scale in-situ field exposure testing has also been conducted
by different researchers [6]. Some researchers tested specimens in long-term tests
with up to 10 years of field exposure to accelerate ASR [7]. These tests provided
valuable results and knowledge on evaluating the residual load capacity of ASR-
affected members. The long test duration of these field tests is required due to the
reality that ASR damage takes a long time to develop in structures, but research
needs call for rapidly and reliably producing ASR expansion in the laboratory with
appropriate accelerated test conditions [1]. Hence, a reliable and rapid accelerated
laboratory test to determine the risk of ASR expansion is needed.
In this study, we applied a novel accelerated test using an autoclave with 80 °C
steam curing to study the flexural and shear behavior of small-scale reinforced
concrete beams affected by ASR. The beams were longitudinally reinforced with two
levels of reinforcement ratios. For simplicity, no shear reinforcement was used for
the beams. Load carrying capacity tests on the small-scale reinforced concrete beams
were conducted. Moreover, the mechanical properties of ASR-affected concrete
under accelerated tests were investigated.

2 Methods

2.1 Materials and Mix Proportions

A general-purpose (Type GP) cement with equivalent alkali content (Na2 Oeq ) of
0.50%, a nonreactive sand (Sydney sand), and a highly reactive dacite aggregate with
a maximum nominal size of 20 mm as coarse aggregate, were used in the concrete
mixes. As for the reinforcement, deformed bars with either 5-mm diameter (N5) or
8-mm diameter (N8) were used. In addition, to promote ASR in the accelerated test,
technical-grade sodium hydroxide (NaOH) pellets with purity of 98% were used to
raise the alkali content to 2.5% Na2 Oeq by mass of cement in the concrete. The NaOH
pellets were pre-dissolved in a fraction of the mixing water 24 h prior to concrete
mixing. The mix proportions for all of the small-scale reinforced concrete beams,
cylinders and prisms were: cement: 520 kg/m3 ; nonreactive sand: 620 kg/m3 ; 20-mm
Investigation of ASR Effects on the Load-Carrying Capacity … 45

highly reactive dacite aggregate: 1160 kg/m3 ; water: 192.5 kg/m3 ; and NaOH pellets:
13.69 kg/m3 .

2.2 Specimen Fabrication and Steam-Curing Procedure

We fabricated 12 small-scale reinforced concrete beams with a size of 100 × 100

× 340 mm; 6 beams were reinforced with two N5 deformed bars and the other 6
beams were reinforced with two N8 deformed bars as the main reinforcement. The
reinforcement ratios for the two series of small-scale beams were 0.39% and 1.0%,
respectively. No transverse reinforcement was provided for the beams. Concrete
cover of the beams was maintained at 20 mm. Details of reinforcement of the beams
are shown in Fig. 1.
Plain ∅ 100 mm (diameter) × 200 mm cylinders and 75 × 75 × 285 mm prisms
were also cast. The cylinders were used for the mechanical properties test and
the prisms were for free expansion measurement. Specimens were demolded after
24 h from casting. Thereafter, all specimens were stored in a humidity cabinet with
temperature of 23 °C and relative humidity (RH) of 90%.
At the age of 28 days, a testing regime of 80 °C steam curing using an autoclave
(Zirbus LVSA 50/70 with a chamber volume of 153 L) was used to accelerate ASR.

Fig. 1 Reinforcement details of small-scale reinforced concrete beams: a with two N5 deformed
bars and b with two N8 deformed bars (all dimensions are in mm)
46 J. Cao et al.

Fig. 2 Acceleration of the alkali–silica reaction (ASR) using the Zirbus LVSA 50/70 autoclave:
a specimens in the autoclave and b time–temperature cycles for accelerating ASR

The inside the autoclave chamber was kept at atmospheric pressure. In total, three
cycles of steam curing were applied, with 60 h of steam curing for each cycle.
Figure 2a shows test specimens as placed in the autoclave, and Fig. 2b illustrates the
temperature–time relationship of steam curing for the three cycles.
At the end of each cycle, the free expansion of the prisms, due to the accelerated
ASR, was recorded and the next cycle was applied.

2.3 Expansion Measurements

Initial lengths of the prisms were measured and recorded using a digital comparator
after demolding. After each cycle of steam curing in the autoclave, the prisms were
taken out and stored in sealed plastic bags for 6 h to cool down to room temperature
at 23 ± 2 °C, and then the length measurements were taken. Changes in length were
used to calculate the expansion of the specimens after 1, 2, and 3 cycles of autoclave
steam curing.

2.4 Mechanical Property Testing

At the age of 28 days, the modulus of elasticity and compressive strength were
measured in accordance with AS1012.17 [8] and AS1012.9 [9] on ∅ 100 × 200 mm
cylinders; at the end of each cycle, three cylinders were taken out of the autoclave
and mechanical property tests were carried out.
Investigation of ASR Effects on the Load-Carrying Capacity … 47

2.5 Load Carrying Capacity Testing Under Four-Point

Loading

For each batch, load carrying capacity tests were conducted on two reinforced beams
at the age of 28 days under four-point loading; at the end of each cycle, one rein-
forced beam was taken out, cooled to room temperature, and tested for load carrying
capacity. The remaining beam was kept for investigating long-term ASR effects on
load carrying capacity. During the load capacity test, a real-time digital image corre-
lation system (Mercury RT®) was used to carry out strain and in-plane displacement
measurements.

3 Results and Discussion

3.1 Cracking of Specimens

Figure 3 shows the cracking pattern of the concrete cylinders and prisms after accel-
erated ASR. Typical external map cracking was observed on the surface of the spec-
imens due to accelerated ASR expansion after three cycles of steam curing in the
autoclave.

Fig. 3 External map cracking on cylinders and prisms after three cycles of steam curing in an
autoclave
48 J. Cao et al.

3.2 Expansion of Concrete Prisms

Figure 4 shows the length change of the concrete prisms from the time of demolding
and after three cycles of steam curing in the autoclave. A slight shrinkage of approx-
imately 0.019% was recorded during storage in the humidity cabinet up to the age
of 28 days. Afterwards, due to accelerated ASR in the autoclave, the length of the
prisms increased with each steam-curing cycle. The average expansion of the prisms
was recorded as 0.05% after one cycle, 0.13% after two cycles and it reached about
0.18% after three cycles. According to ASTM C1778-20, aggregate having 1-year
CPT expansion ≥ 0.12% and < 0.24% can be classified as highly reactive. For
dacite aggregate, the 1-year CPT expansion result from Cement Concrete & Aggre-
gates Australia is ≈0.23%. In the current study, using three cycles of steam curing,
the expansion reached ≈0.18%. This result shortens the testing period for classifying
aggregate reactivity. However, more cycles are needed for slowly reactive aggregate.
Further study is suggested to test more aggregates ranging from nonreactive to very
highly reactive to establish a standard testing procedure with fine-tuned parameters
including temperature, duration, heating and cooling rates and number of cycles,
to determine aggregate reactivity, following the ASTM C1778-20 expansion limit
criterion.

Fig. 4 Expansion of concrete prisms under three cycles of accelerated ASR in the autoclave
Investigation of ASR Effects on the Load-Carrying Capacity … 49

Fig. 5 Modulus of elasticity a and compressive strength b and before and after three cycles of
steam curing in an autoclave

3.3 Mechanical Properties of Concrete Under Accelerated

ASR Test Conditions

Figure 5a shows the modulus of elasticity test results at 28 days and after 1, 2,
and 3 cycles of steam curing in the autoclave. It is generally acknowledged that the
modulus of elasticity is the most sensitive mechanical property influenced by ASR.
As can be seen, it decreased as expected with the ASR expansion achieved after
each cycle of 80 °C steam curing using the autoclave. After the third cycle when
average expansion reached 0.18%, a reduction of 39% was recorded in comparison
with the initial 28-day value. The reduction was attributed to the microcracking of
the concrete caused by accelerated ASR.
Figure 5b shows the compressive strength test results, which demonstrated that
the compressive strength initially increased with increasing of number of cycles
until the end of the second cycle, at a relatively low expansion level, and thereafter,
the compressive strength showed a decreasing trend. With increasing expansion,
compressive strength is expected to continue to decrease. This trend had been already
observed by Gautam et al. [10]. They boosted the alkali content of concrete speci-
mens to 1.25% Na2 Oeq . Samples were stored in hermetically sealed plastic pails and
conditioned at 38 °C with RH > 95%. They reported that at age 365 days when ASR
expansion reached 0.24–0.35%, the maximum reduction in compressive strength was
4–6%, in comparison with the 28-day compressive strength [10].

3.4 Load Carrying Capacity of Reinforced Concrete Beams

To investigate the reduction in load capacity of the reinforced concrete beams after
accelerated ASR, load carrying capacity tests were conducted under four-point
50 J. Cao et al.

loading at the age of 28 days and after 1, 2, and 3 cycles of accelerated ASR. All the
beams with 2 × N5 reinforcing steel bars failed in flexure and all the beams with 2 ×
N8 bars failed in shear. Figure 6 shows the initial load capacity of a typical reinforced
concrete beam with 2 × N5 tested at the age of 28 days failing in flexure. Figure 7
demonstrates a reinforced concrete beam with 2 × N8 reinforcing bars tested after
two cycles of accelerated ASR showing typical shear failure.
The load carrying capacity test results are shown in Fig. 8. It can be seen that, for
the reinforced beams with 2 ×N5 bars, the flexural capacity was not significantly

Fig. 6 Reinforced concrete beam with 2 × N5 bars tested at 28 days: a test set-up and b load–
displacement curve

Fig. 7 Reinforced beam with 2 × N8 bars tested after two cycles of accelerated alkali–silica
reaction: a failure mode and b load–displacement curve
Investigation of ASR Effects on the Load-Carrying Capacity … 51

Fig. 8 Reinforced concrete beams with a 2 × N5 bars failed in flexure and b with 2 × N8 bars
failed in shear

influenced by ASR expansion achieved under three cycles of accelerated test. Failure
load of the reinforced beam with 2 × N8 bars was found increased after 1, 2, and
3 cycles of accelerated ASR in comparison with the 28-day value. This could be
attributed to the prestressing effect of ASR expansion. Meanwhile, some reduction
in the failure load after each cycle was recorded due to only one sample being tested
for each cycle. Load carrying capacity at higher ASR expansion levels, however,
needs further investigation.

4 Conclusions

(1) Using three cycles of 80 °C steam curing in an autoclave, the ultra-accelerated

test produced ASR damage and expansion of alkali boosted concrete elements
within 9 days.
(2) Modulus of elasticity systematically decreased with increasing number of 80 °C
steam-curing cycles. Compressive strength increased until the end of the second
cycle, and thereafter showed a decreasing trend.
(3) Shear resistance of reinforced beams with 2 × N8 bars was increased after 1,
2, and 3 cycles of accelerated ASR in comparison with the 28-day value. This
could be attributed to the prestressing effect of ASR expansion.
(4) Flexural capacity of the beams reinforced with 2 × N5 bars was not significantly
influenced by the extent of ASR expansion achieved in the current accelerated
autoclave test.

Acknowledgements This research was funded through an Australian Research Council Research
Hub for Nanoscience Based Construction Materials Manufacturing (IH150100006) with the support
of Cement Concrete and Aggregates Australia.
52 J. Cao et al.

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the performance of a reinforced concrete structure over a six-year period. Mag Concr Res
41(147):69–77
3. Fan S, Hanson JM (1998) Effect of alkali silica reaction expansion and cracking on structural
behaviour of reinforced concrete beams. ACI Struct J 95:498–505
4. Swamy RN, Al-Asali M (1989) Effect of alkali-silica reaction on the structural behavior of
reinforced concrete beams. ACI Struct J 86(4):451–459
5. Bilodeau S, Allard A, Bastien J, Pissot F, Fourinier B, Mitchell D, Bissonnette B (2016)
Performance evaluation of thick concrete slabs affected by alkali-silica reaction (ASR)–part
II: structural aspects
6. Deschenes D, Bayrak O, Folliard K (2009) Shear capacity of large-scale bridge bent specimens
subject to alkali-silica reaction and delayed ettringite formation, Structures Congress. In: Don’t
mess with structural engineers: expanding our role, pp 1–9
7. Hamada H, Otsuki N, Fukute T (1989) Properties of concrete specimens damaged by alkali-
aggregate reaction, laumontite related reaction and chloride attack under marine environments.
In: Proceedings of the 8th international conference on AAR. Kyoto, Japan, pp 603–608
8. AS 1012.17 (2014) Methods of testing concrete, method 17: determination of the static shord
modulus of elasticity and poisson’s ratio of concrete specimens, Standards Australia Ltd,
Sydney, Australia
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mortar and grout specimens, Standards Australia Ltd, Sydney, Australia
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3D printed Ultra-High Performance
Concrete: Preparation, Application,
and Challenges

G. Bai, G. Chen, R. Li, L. Wang, and G. Ma

Abstract 3D printed ultra-high performance concrete (3DP-UHPC) plays an impor-

tant role in the realization of ultra-high compressive and tensile strengths. Consid-
ering the particular characteristics of UHPC, the conversion of UHPC to 3DP-UHPC
is a complex phenomenon and has been the subject of numerous studies. It is very
important to be able to design a thixotropic structure in the early hydration stage
for bridging the gap between the slow setting of UHPC and the rapid setting of the
3D printing procedure. In the design and application of 3DP-UHPC, requirements
such as the ratio of coagulant and flocculant, fiber alignment, reinforced 3D printed
no-rebar reinforced concrete, safety, cost etc. need to be taken into account. We
present a comprehensive review of 3DP-UHPC in concrete construction from prepa-
ration to application, including design method, raw materials, mechanical, reinforced
methods, and applications. Finally, recommendations are provided to promote the
application of 3DP-UHPC in engineering practice.

Keywords 3D printing · Fiber alignment · Printability · Ultra-high performance

concrete

1 Introduction

The rapid development of 3D concrete printing (3DCP) has the potential to greatly
reduce labor demand, improve sustainability, reduce construction costs, and effec-
tively overcome the dilemma faced by traditional construction methods [1–4]. In

G. Bai · L. Wang (B) · G. Ma

School of Civil and Transportation Engineering, Hebei University of Technology, Tianjin, China
e-mail: wangl1@hebut.edu.cn
G. Chen
Zhong Dian Jian Ji Jiao Expressway Investment Development Co., Ltd., Shijiazhuang, China
R. Li
Yaobai Special Cement Technology R&D Co., Ltd., Xian, China

© The Author(s) 2023 53

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_8
54 G. Bai et al.

recent years, substantial achievements have been made by 3DCP in the field of archi-
tecture and civil engineering. One of the challenges is that conventional steel bar rein-
forcement cannot be directly integrated into the printed concrete. Researchers and
engineers have tried different reinforcement methods by applying continuous fiber,
shut fiber, microcable, mesh and U-nails etc., to improve the brittleness of 3D printed
concrete [5–8]. Fiber reinforcement is widely applied for printed concrete due to its
effectiveness and ease of operation. Small-sized fibers show less interference with the
flexible extrusion characteristics of 3D printing. More importantly, the mechanical
properties of fiber-reinforced concrete can meet the structural requirements, such as
compressive strength and tensile strain exceeding 100 MPa and 4%, respectively [9,
10]. Ultra-high performance concrete (UHPC), as a type of fiber-reinforced concrete,
can meet these structural requirements [11, 12] and is currently mainly being used
for new structures, reinforcement, and repair of existing infrastructure [13, 14].
The development of 3D printed UHPC (3DP-UHPC) will greatly drive the appli-
cation of 3D printing technology to structural engineering. UHPC is considered
a combination of self-compacting concrete, high-performance concrete, and fiber-
reinforced concrete [15]. The good construction performance of UHPC is strong
related to the casting procedure, but this advantage is difficult to match with the
3D printing construction of layer by layer stacking. Current research has eliminated
the gap between cast and 3D printed UHPC in construction by adding viscosity-
modifying admixtures (VMA), such as nano-clay, hydroxypropyl methylcellulose
(HPMC), etc. [16–20]. Similar to traditional UHPC, 3DP-UHPC will be widely
used in new and existing structures. For example, the construction of a curvilinear
bench with free form and light structure [16]. In addition, more attention is being
paid to the dynamic performance. Zhou et al. [20] discussed the performance of
3DP-UHPC based on projectile and explosive impacts tests. Yang et al. [21] carried
out split-Hopkinson pressure bar (SHPB) tests and analyzed the strain rate effect of
3DP-UHPC
3DP-UHPC has attracted much attention because of its excellent mechanical prop-
erties and ongoing research mainly focuses on the preparation of materials, static,
and dynamic mechanical properties, and so on. Given the complexity of the prepa-
ration method of 3DP-UHPC and the unknowns of the problems that may be faced
in its applications, we systematically reviewed the latest research progress on 3DP-
UHPC. Finally, some suggestions are put forward to promote the development of
3DP-UHPC based on the current challenges.
3D printed Ultra-High Performance Concrete: Preparation, Application … 55

2 Preparation of 3DP-UHPC

2.1 Design of UHPC for 3DCP

The gap between the printing and casting procedures of UHPC mainly depend on
rheological properties. 3D printed concrete is a typical yield stress material; that
is, its yield stress first increases and then decreases with increasing shear rate, and
finally maintains at a certain yield stress platform, as shown in Fig. 1a. The cast
UHPC is self-leveling in the static state because its yield stress makes it difficult
to maintain its shape. Moreover, cast UHPC is a shear thickening fluid; that is, its
yield stress increases rapidly with increasing shear rate, as shown in Fig. 1a. On the
other hand, shear thickening fluid means that the shear rate needs to be continuously
increased to overcome the yield stress of UHPC. UHPC strips will be narrowed or
even interrupted when the yield stress exceeds the maximum shear stress provided
by the 3D printer.

Fig. 1 a Fluid types of ultra-high performance concrete (UHPC) and 3D printed concrete (3DPC);
b printability principle of 3DP-UHPC [22]; c schematic of shear resistance of fresh 3DPC without
and with coarse aggregate [4]
56 G. Bai et al.

The methods for modifying UHPC to achieve printability are the chemical hydra-
tion accelerated hardening method [20] (CM) and the physical flocculation method
(PM) [16, 17]. The CM matches the hydration rate of UHPC with the printing
rate by adding materials that change the cement hydration rate. For example, three
levels of fast, medium, and slow hydration rates of UHPC can be represented by
the curves ➀–➂ in Fig. 1b. These different rates can be achieved by reducing the
accelerator dosage, such as sulfoaluminate cement. The curves ➀–➂ can match the
printing rate of slow, medium, and fast, respectively. However, the matching gap
between the hydration rate and printing rate of UHPC usually leads to a short open
time of CM. It is easy to cause large deformation and rough surface dry cracking
or even collapse and fracture due to insufficient hydration and too rapid printing.
PM matches the printing rate by adding VMAs, such as silica fume (SF), nano-clay,
and HPMC, to make thixotropic structures in the UHPC before hydration structure
formation. For example, based on curves ➀–➂ in Fig. 1b, the addition of the same
VMA gets curves ➃–➅. The addition of VMA makes the material printable earlier.
The thixotropic structure will produce obvious deformation when the accumulated
weight of the UHPC exceeds its yield stress, affecting the forming accuracy and even
causing collapse. Therefore, PM is limited to a certain range of printing rates to allow
the hydration rate to follow up.
The potential third printability control method, namely the framing effect provided
by raw materials, needs attention because of the current development trend of mixing
large-size aggregate into UHPC and the existence of steel fibers. In our previous study
on the printability of large-size aggregate, we found that construction deformation
and strength were derived from the bonding force of cementitious materials and the
aggregate biting force [4], as shown in Fig. 1c. Therefore, the contribution of the
framing effect to the improvement of yield stress cannot be ignored.
Designing UHPC for 3DCP will be based on the above principles. Specifically, the
purpose of the preparations is to improve the early yield strength of UHPC, reduce its
fluidity and improve its shape retention ability by adding a regulator or VMA alone
or adding both. Limited by the lack of quantitative preparation theory, the specific
dosage can only be obtained by testing. It should be noted that determination of
these doses is significantly related to the printing equipment and the selected printing
process, which reduces the repeatability of the material mix to a certain extent.

2.2 Manufacturing of 3DP-UHPC

Table 1 lists the raw materials and mix proportions of 3DP-UHPC in the existing
literature. Zhou et al. [20] also adopted CM to promote the hydration rate to achieve
matching with the printing process, specifically by adding slag. Arunothayan et al.
[16, 17] used PM to achieve printability of UHPC, by adding HPMC and nano-clay. It
cannot be ignored that the proportion of SF is relatively large, accounting for ≈30% of
the cementitious materials. Excessive SF also contributed to the printability of UHPC.
To avoid the defects of the short open time of CM and the low unit construction rate
3D printed Ultra-High Performance Concrete: Preparation, Application … 57

of PM, our group proposed a method of combining CM and PM to achieve printable

UHPC[19]. Specifically, sulfoaluminate cement was used to replace the 10% mass
fraction of Portland cement to accelerate the hydration rate to preliminarily match
the printing rate. Then, the thixotropic material Nano clay and HPMC were added
to supplement the insufficient yield stress of UHPC to realize the continuous and
stable printing. Here, the mass ratio of the accelerator composed of sulfoaluminate
to the flocculant composed of nano-clay and HPMC was ≈11:1. The mix proportions
of Yang et al. [18] are the most popularized, because the raw material composition
does not different from the cast UHPC base except for the addition of an appropriate
amount of nano-calcium carbonate, and its mechanical properties are also the highest,
which may be related to the use of a water reducing agent.
The microstructure of UHPC mixed with sulfoaluminate cement, sulfoaluminate
cement and VMA was observed by scanning electron microscopy. The preparation
time of the two samples was between the initial and final setting. Figure 2 shows the
obvious differences in the microstructure of UHPC before and after adding VMA.
The microstructure of UHPC without VMA is mainly C-S–H gel with 3D network
structures (Fig. 2a); that is, 3DP-UHPC prepared by CM. The microstructure of
UHPC doped with sulfoaluminate cement and VMA is mainly flocculent, formed by
the adsorption of hydration products by VMA. A similar flocculent microstructure
after adsorption was reported by Zhang et al. [23].

Table 1 Mix proportions of 3DP-UHPC (kg/m3 )

Raw materials References
Arunothayan et al. Yang et al. Bai et al. Zhou et al.
[16] [17] [18] [19] [20]
Cement 700 273 750 612 500
Fly ash – 410 165 112 –
Sand 1000 1268 1080 1200 1000
HRWRA 15 18 10 10 15.5
Steel fibera 2% (13 mm) 2% (6 mm) 1% (6 mm) 2.5% (12 mm) 4.5%b (6 mm)
Water 160 156 154 180 200
Silica fume 300 293 165 200 200
CSA – – – 68 –
HPMC 3 – – 1 –
Slag – – – – 300
NCC – – 24 – –
Nano clay – 5.5 – 5 –
a Fiber volume fraction; b fiber composed of 4% steel fiber and 0.5% basalt fiber CSA, sulfoaluminate

cement; HPMC, hydroxypropyl methylcellulose; HRWRA, high-range water-reducing admixture;

NCC, nano-calcium carbonate
58 G. Bai et al.

Fig. 2 Scanning electron microscopy images of a hydrating structure: C–S–H gel; b flocculation
structure formed by nano-clay and hydroxypropyl methylcellulose (HPMC)

2.3 Mechanical Anisotropy of 3DP-UHPC

The compression and flexural properties of 3DP-UHPC in the existing literature

were statistically analyzed. The average compressive strength of all 3DP-UHPC
samples at 28 days was in the range of 120–160 MPa. The flexural strength of
specimens was between 30 and 50 MPa. The printing procedure had\s a favorable
effect on the orientation of fibers; that is, the orientation of steel fibers was parallel
to the printing direction. Figure 3a, b show computed tomography scans of single or
interwoven printed 3DP-UHPC samples. The consistency between fiber orientation
and printing direction is obvious. The fiber orientation and stress direction should
be designed to be consistent to reach optimal application of UHPC materials [24].
For example, the flexural strength of 3DP-UHPC specimens nearly doubled that of
cast UHPC specimens [22], as shown in Fig. 3c. The flexural and tensile properties
of 3DP-UHPC will be either ductile or brittle in different directions due to fiber
orientation [24]. 3DP-UHPC has obvious mechanical anisotropy and generally has
weaker mechanical properties along the interstrip or interlayer directions, as shown
in Fig. 3d. Therefore, it is necessary to optimize the printing path to fully realize the
toughening effect of fiber orientation and achieve the excellent bearing performance
of 3DP-UHPC structures [18].

3 Applications of 3DP-UHPC

3.1 Large-Scale Special Components

3DP-UHPC with its ultra-high mechanical properties is expected to alleviate the

contradiction that 3D printed concrete structures cannot be directly used in engi-
neering under steel-free reinforcement conditions. Some specially shaped structures
with independent bearing capacity have been constructed, although no large-scale
3D printed Ultra-High Performance Concrete: Preparation, Application … 59

(a) (b)

20
PS0 PS5 PS10
Flexural strength/MPa

CS0 CS5 CS10 CS15

15

10

0
0 2 4 6 8
Displacement/mm
(c) (d)

Fig. 3 Computed tomography scans of a single printed and b interwoven printed 3D printed ultra-
high performance concrete (3DP-UHPC) sample; c flexural strength [22] and d the modulus of
rupture (MOR) [24] of 3DP-UHPC. CS, cast sample; PS, printed sample

application of 3DP-UHPC has been found. There are two construction paths for
3DP-UHPC in new structures; one is the construction using 3DP-UHPC alone, such
as a curvilinear bench [16] or hollow column [17], as shown in Fig. 4a, b; another
is to use 3DP-UHPC to construct permanent templates, such as the specially shaped
columns constructed by our group, shown in Fig. 4c. Attention should be paid to
the construction of steel-free reinforcement concrete structures with 3DP-UHPC,
namely UHPC-reinforced concrete (URC). This has obvious advantages in main-
taining the flexibility of 3D printing and reducing the time and labor consumption
of rebar implantation.

3.2 As the Reinforcement Materials

The in-process reinforcing method (IRM) was inspired by reinforced concrete mate-
rials under the technical constraints of steel bar implantation. As a type of IRM, the
dual 3D printing procedure for URC has proved to be a potential and effective rein-
forcement method [19]. This is because UHPC and concrete are both cement-based
60 G. Bai et al.

Fig. 4 a Curvilinear bench [16]; b hollow column [17]; c variable diameter hollow column

materials, which means that there are advantages in the printing path following and
interface bonding performance [25]. The principle of an extrusion system for dual
3D concrete printing is that UHPC and 3DPC are fused at the nozzle to complete
synchronous printing, as shown in Fig. 5a. Figure 5b shows the three-direction profile
of the sample printed by this technique. It can be seen that the 3DP-UHPC is similar
to the rebar arrangement in concrete. Another advantage of this technique is that the
concrete wrapping UHPC blocks air and water to protect steel fibers from corrosion.
In addition, it can be predicted that this method will maximize the utilization rate of
UHPC to achieve structural strengthening and toughening, such as printing along the
stress line of the bridge or in the required position based on the results of topological
optimization.

3.3 As Protection Against Explosion and Impact

3DP-UHPC application has been extended to anti-explosion and anti-impact struc-

tures due to its good energy absorption capacity. The effectiveness of 3DP-UHPC
has been demonstrated by experimental studies on explosion resistance and impact
resistance. Yang et al. [21] tested the dynamic compression performance of 3DP-
UHPC. Zhou et al. [20] explored the unique behavior of 3DP-UHPC subjected to
projectile and explosive impacts and revised an empirical formula for calculating the
crater depth of 3DPC. The preparation process is shown in Fig. 6(a). Our recent work
3D printed Ultra-High Performance Concrete: Preparation, Application … 61

Fig. 5 a Schematic of extrusion system for dual 3D concrete printing; b three orthogonal cross-
sections of 3DP-UHPC reinforced concrete specimen [19]

explored the contact explosion performance of 3DP-UHPC. In that work, 3DP-UHPC

with different thicknesses was used to strengthen ordinary concrete. The preparation
process is shown in Fig. 6(b). It is worth noting that there is no rebar inside the slabs of
the test group. We found that PURN8, the thickness of 3DP-UHPC replacing ordinary
concrete was 53.3%, was marginally improved to that of reinforced concrete.

4 Challenges of 3DP-UHPC

Despite the increasing amount of research published thus far on 3DP-UHPC, many
challenges and research barriers require further innovative exploration.

4.1 Standardization of Materials

The most fundamental challenge is standardization of 3DP-UHPC. In the prepara-

tion process of 3DP-UHPC, our proposed method that combines the CM and PM
had qualitative mixing ratio designs based on experience. Following a certain rule
of mix ratio design will be more conducive to standardization of 3DP-UHPC and
improvement of material properties. The mixing ratio design of 3DP-UHPC has been
completed, but regulating the fluidity of UHPC brings a series of challenges to the
ultra-high performance of the material, such as mechanical and durability weakening
problems due to the density of the material and the bonding performance between
the fiber and the matrix. In addition, steel fibers cannot solve the strain-softening
62 G. Bai et al.

Fig. 6 a Technology for preparing 3D printed target [20]; b construction process of test slabs.
UHPC, ultra-high performance concrete

behavior after peak stress, and it is difficult to achieve large tensile strains of mate-
rials. Exploratory studies have found that mixing steel fibers and polyethylene fibers
can effectively improve this problem [10]. Therefore, an integrated 3DP-UHPC mix
ratio design method combining comprehensive fluid performance regulation, fiber–
matrix interface, and synergistic toughening of multiple fibers can contribute to
optimization of material properties, and further research on this material is required,
such as durability and mechanical properties under different strain rate conditions,
etc.
3D printed Ultra-High Performance Concrete: Preparation, Application … 63

4.2 Flexible Design of Structures

It is very important to manufacture 3DPC structures considering both safety and

economy in structural applications. 3DP-UHPC has good mechanical toughness
under different strain rate conditions and will be widely used in the construction
of new structures and the reinforcement of existing military and civil buildings.
Reinforcing plain concrete where it is needed provides a specific way to achieve the
construction of 3DPC structures. 3DP-UHPC is regarded as a disruptive method for
construction due to the advantages of 3D printing and its economics. Initial explo-
ration has demonstrated the feasibility of 3DP-UHPC to partially replace or even
completely replace steel reinforcement for concrete structures, but its application
in larger-scale structures requires refined design, such as combining prestressing
methods and topological optimization methods, design, etc. The reinforcement of
existing buildings with 3DP-UHPC is also a very important structural application
scenario. The coordinated deformation of interfaces of 3DP-UHPC and other mate-
rials under different working conditions will require special attention, such as the
shrinkage deformation performance and load-deformation performance of the two
materials. The structural response of 3DP-UHPC under strain rate loading will also
be a very interesting research area.

5 Summary and Outlook

Combined with the reviewed research, we have provided a specific understanding of

the research progress of 3DP-UHPC, which will be of significance to promote the
application of 3DCP.
• Integrated material design combining fluid modification, fiber–matrix interface
enhancement, and hybrid fiber toughening can be used to improve the mix
design of 3DP-UHPC at the theoretical level, such as quantifying the impact
of VMAs on the printability of UHPC, improving fiber–matrix cohesion through
nanomaterials, and finding critical fiber lengths, etc.
• Further exploration of the mechanical properties and durability of materials under
different strain rates and establish corresponding constitutive models to provide
a theoretical basis for structural design.
• Further exploration of the application of 3DP-UHPC in newly constructed 3DPC
structures or the reinforcement of existing buildings and explore the structural
response laws under different strain rate loads to realize the safety and economy
of 3D printed structures.

Acknowledgements We acknowledge the financial support from the National Natural Science
Foundation of China (Nos 51878241, 52078181 and 52178198), the Natural Science Foundation
of Hebei (Nos. E2021202039 and E2022202041), and the Natural Science Foundation of Tianjin
(No. 20JCYBJC00710).
64 G. Bai et al.

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Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Nanosilica-Modified Hydrogels
Encapsulating Bacterial Spores
for Self-healing Concrete

J. Feng and S. Qian

Abstract Microbially induced calcium carbonate precipitation is effective in

achieving self-healing of concrete cracks when the bacteria are well protected in
concrete with a high pH and dense microstructure. Calcium alginate hydrogels are
appropriate for encapsulating bacteria in concrete due to the mild environment with
rich moisture in the hydrogels. Nevertheless, the low alkaline tolerance and breakage
ratios of the hydrogels after concrete cracking restrict their applications in concrete.
To address these problems, nanosilica was doped into calcium alginate hydrogels
with encapsulated bacterial spores to react with the Ca(OH)2 surrounding hydrogels
in concrete. Due to the modification by nanosilica, the bond of the hydrogels with
cement matrix was enhanced as needle-like C–S–H was generated at the interface
after hydration for 7 days. Moreover, the urease activity of the encapsulated spores
in the modified hydrogels was higher than that in plain hydrogels after submersion in
saturated Ca(OH)2 solution or simulated concrete solution for 7 days. Therefore, it
can be concluded that nanosilica holds promise for modifying hydrogels to improve
the effectiveness of encapsulated bacterial spores for self-healing of concrete.

Keywords Bacterial spores · Hydrogels · Nanosilica · Self-healing concrete

1 Introduction

Concrete is widely used in construction engineering due to its good durability,

adjustable strength, and low cost. However, cracking can occur in concrete due
to overloading and volume instability [1], which will reduce both the mechanical
performance and durability of concrete structures. Hence, to prolong the service life
of concrete structures, timely repair of concrete is critical.

J. Feng · S. Qian (B)

School of Civil and Environmental Engineering, Nanyang Technological University, Nanyang,
Singapore
e-mail: szqian@ntu.edu.sg

© The Author(s) 2023 67

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_9
68 J. Feng and S. Qian

Currently, concrete repair is achieved manually using epoxy resin [2] or some
cementitious material [3]. Nevertheless, manual repairing cannot be timely because
the detection of cracks takes time and is not feasible for inaccessible cracks. There-
fore, developing concrete with a self-healing capacity that can effectively heal cracks
without human intervention is desired.
Concrete itself has a certain self-healing ability resulting from further hydration
of cementitious materials and carbonation [4, 5], but only for cracks of limited sizes.
The self-healing capacity can be further improved by incorporating some compo-
nents specific for self-healing into the concrete. For example, minerals [6, 7], super-
absorbent polymers [8, 9], and shape memory alloy [10, 11] have been added to
concrete to enhance self-healing. In addition to these materials, bacteria-based self-
healing agents for self-healing concrete has been investigated in recent years and
promising results have been obtained [12]. For bacterial self-healing agents, carriers
to encapsulate or immobilize the bacteria are required as the concrete environment
with high pH [13] and dense microstructure is incompatible [14].
Among the carriers to encapsulate bacteria in concrete, hydrogels with moderate
pH environment and rich moisture content have high potential for protecting bacteria
[14]. Specifically, calcium alginate, which has good biocompatibility [15, 16], can
be used to encapsulate bacteria, but its susceptibility to environmental factors [17]
and poor bonding with concrete matrix could lead to the ingress of alkalis and low
efficiency of releasing bacteria after cracking.
To address these issues, nanosilica was doped into calcium alginate hydrogels
to react with the surrounding calcium hydroxide, simultaneously lowering the local
pH and generating C–S–H at the interface between the hydrogel and cement matrix,
which can enhance the bonding of hydrogels with concrete. A previous report [18]
revealed that hydration product could be generated around or within hydrogels that
contain silica, indicating the feasibility of the approach in this research. Herein, the
microstructure of cement paste with hydrogels was observed and the alkali tolerance
of hydrogels encapsulating bacterial spores was evaluated to analyze the effects of
nanosilica modification on calcium alginate hydrogels.

2 Methods

2.1 Preparation of Bacterial Spores

One ureolytic bacteria Lysinibacillus sphaericus LMG 22,257 from Belgian Co-
ordinated Collection of Micro-organisms were used to prepare bacterial spores. The
preparation methods were in accordance with the steps in [19].
Nanosilica-Modified Hydrogels Encapsulating Bacterial Spores … 69

2.2 Preparation of Hydrogels Encapsulating Bacterial Spores

7.5 g/L nanosilica powder (10–20 nm) was dispersed in sodium alginate solution
(15 g/L) by sonication for 15–20 min before 2.0wt% bacterial spores were added to
the mixture. Next, the mixture was dropped into calcium nitration solution (0.1 M)
using a peristaltic pump with a rotary speed of 3 rpm. The prepared hydrogels were
collected by gravity sedimentation then hardened in a fresh 0.1 M calcium nitrate
solution for 24 h at 5 °C. Finally, the hydrogels were separated and washed three
times with distilled water before being freeze-dried and stored at 4 °C. The plain
hydrogels were synthesized using the same procedures except for the incorporation
of nanosilica powder.

2.3 Observation of Interface Between Hydrogel and Cement

Matrix

Cement paste with hydrogels was prepared by using Portland Cement I 52.5, tap
water and hydrogels at a mass ratio of 1:0.5:0.0055. After moist curing (23±1°C,
75±5%°C relative humidity) for 7 and 28 days, the specimens were broken into
pieces before being immersed in isopropanol for 24 h to stop hydration. The samples
were impregnated in epoxy, then cut with a precision saw to expose the cement paste
with hydrogels; the exposed surface was further ground, polished and washed with
ethanol. Afterwards, the samples were vacuum dried and coated with gold before
being observed under a scanning electron microscope (FESEM, JEOL JSM-7600F)
at backscattered electron (BSE) mode at an accelerating voltage of 15 keV.

2.4 Revival of Bacterial Spores Encapsulated in Hydrogels

After Exposure to Alkali Environments

The hydrogels with encapsulated endospores were immersed in saturated Ca(OH)2

solution and simulated concrete pore solution containing 0.001 M Ca(OH)2 , 0.2 M
NaOH and 0.6 M KOH [20] for 7 days respectively. After that, 1 g hydrogels were
removed from the solution and washed three times with distilled water before being
incubated in 100 mL sterile UYE medium (20 g/L urea and 20 g/L yeast extract). The
optical density at a wavelength of 600 nm and the urease activities of the medium were
measured by using a UV/Vis spectrometer (UV mini-1240, Shimadzu) and conduc-
tivity meter (DDS-11A, Lightning), respectively, after the inoculated mediums were
cultivated for 1, 2, and 3 days.
70 J. Feng and S. Qian

3 Results and Discussion

The morphology of cement paste with modified hydrogels is shown in Fig. 1. After
hydration for 7 days, nanosized calcium silicate hydrate (C–S–H) with needle-like
morphology was observed on the outer surface of modified hydrogel, as shown
in Fig. 1a, b. The generations of C–S–H were due to the reaction of the incor-
porated nanosilica on the hydrogel surface with calcium hydroxide in the cement
paste, suggesting the effect of nanosilica modification on interface enhancement
between the hydrogels and cement matrix. After hydration for 28 days, the BSE
image shown in Fig. 1c revealed no obvious gaps or cracking at the interface, indi-
cating the modified hydrogels bonded well with the cement matrix. The improved
bonding of the hydrogels with cement matrix could facilitate the release of encapsu-
lated bacterial spores after concrete cracking, as the cracks might propagate through
the hydrogels rather than the interfaces, leading to breakage of the hydrogels and
exposure of bacterial spores to the cracks. After the release of bacterial spores, the
endospores can conduct germination and outgrowth with leaching of nutrients from
the concrete matrix, contributing to the generation of calcium carbonate within cracks
by producing urease to catalyze hydrolysis of urea.
The alkaline tolerance of the hydrogels encapsulating bacterial spores is shown
in Fig. 2. After immersion in the alkaline solution for 7 days, the growth of the
encapsulated bacteria in the modified hydrogels during incubation was more rapid
than in the plain hydrogels, as illustrated in Fig. 2a. The optical densities of the
medium with the modified hydrogels reached approximately 2 and 2.5 approximately
after cultivation for 2 days respectively, while the optical densities of the medium
with bacteria in plain hydrogels were close to 2 after cultivating for around 3 days.
With the growth of bacteria, the medium with bacteria encapsulated bacteria in
modified hydrogels presented higher urease activities than that with bacteria in plain
hydrogels, as shown in Fig. 2b. After incubation for 1 day, the urease activities of
medium were 3.7 U/mL and 6.6 U/mL approximately for hydrogels undergoing 7-
day submersion in simulated concrete or saturated Ca(OH)2 solution respectively,
while those in medium with plain hydrogels was < 1 U/mL. Although the urease
activities of medium with plain hydrogels further increased with incubation duration,

(a) (b) (c)

Fig. 1 Backscattered electron images of cement paste with nanosilica modified hydrogels. a After
hydration for 7 days; b close-up view of the red box in (a); c after hydration for 28 days
Nanosilica-Modified Hydrogels Encapsulating Bacterial Spores … 71

Fig. 2 a Optical densities of mediums with modified and plain hydrogels after being immersed in
simulated concrete or saturated Ca(OH)2 solution for 7 days; b urease activities of mediums with
modified and plain hydrogels after being immersed in simulated concrete or saturated Ca(OH)2
solution for 7 days. SMHB denotes the nanosilica modified hydrogels with encapsulated bacterial
spores; HB denotes plain hydrogels with encapsulated bacterial spores

they were still lower than those with modified hydrogels after incubation for 2 and
3 days. Collectively, the bacteria in the modified hydrogels grew more rapidly and
showed higher urease activities than those in plain hydrogels after exposure to an
alkaline environment, which suggested the effectiveness of nanosilica modification
on alkaline tolerance improvement of hydrogels for bacterial encapsulation.

4 Conclusions

This research investigated the feasibility of nanosilica modification of calcium

alginate hydrogels for bacterial encapsulation in self-healing concrete. The results
revealed that the nanosilica modification enhanced the bonding of the hydrogel with
the cement matrix due to C–S–H generation at the interface. Moreover, the modi-
fied hydrogels with encapsulated endospores showed improved alkaline tolerance,
because as the 7-day immersion in simulated concrete and saturated Ca(OH)2 solution
caused less inhibition on bacterial growth and urease activity in medium inoculated
with modified hydrogels encapsulating endospores. Hence, it is feasible to modify
calcium alginate hydrogels with nanosilica to increase the effectiveness of bacterial
encapsulation in concrete. In the future, the effects of nanosilica-modified hydrogels
encapsulating bacterial spores on self-healing performance (i.e. strength recovery,
etc.) of concrete can be investigated.
72 J. Feng and S. Qian

References

1. Mindess S, Young JF (1981) Concrete. Prentice-Hall

2. Yoo DY, Oh T, Shin W, Kim S, Banthia N (2021) Tensile behavior of crack-repaired ultra-high-
performance fiber-reinforced concrete under corrosive environment. J Mater Res Technol
3. Chindaprasirt P, Sriopas B, Phosri P, Yoddumrong P, Anantakarn K, Kroehong W (2021)
Hybrid high calcium fly ash alkali-activated repair material for concrete exposed to sulfate
environment. J Build Eng:103590
4. Huang HL, Ye G, Qian CX, Schlangen E (2016) Self-healing in cementitious materials:
materials, methods and service conditions. Mater Des 92:499–511
5. Reinhardt HW, Jooss M (2003) Permeability and self-healing of cracked concrete as a function
of temperature and crack width. Cem Concr Res 33(7):981–985
6. Feng J, Dong H, Wang R, Su Y (2020) A novel capsule by poly (ethylene glycol) granulation
for self-healing concrete. Cem Concr Res 133:106053
7. Wu X, Huang H, Liu H, Zeng Z, Li H, Hu J, Wei J, Yu Q (2020) Artificial aggregates for
self-healing of cement paste and chemical binding of aggressive ions from sea water. Compos
B Eng 182:107605
8. Lee HXD, Wong HS, Buenfeld NR (2016) Self-sealing of cracks in concrete using superab-
sorbent polymers. Cem Concr Res 79:194–208
9. Snoeck D, Van den Heede P, Van Mullem T, De Belie N (2018) Water penetration through
cracks in self-healing cementitious materials with superabsorbent polymers studied by neutron
radiography. Cem Concr Res 113:86–98
10. Teall O, Pilegis M, Davies R, Sweeney J, Jefferson T, Lark R, Gardner D (2018) A shape
memory polymer concrete crack closure system activated by electrical current. Smart Mater
Struct 27(7)
11. Chen W, Lin B, Feng K, Cui S, Zhang D (2022) Effect of shape memory alloy fiber content and
preloading level on the self-healing properties of smart cementitious composite (SMA-ECC).
Constr Build Mater 341:127797
12. Van Tittelboom K, De Belie N, De Muynck W, Verstraete W (2010) Use of bacteria to repair
cracks in concrete. Cem Concr Res 40(1):157–166
13. Wang J, Van Tittelboom K, De Belie N, Verstraete W (2012) Use of silica gel or polyurethane
immobilized bacteria for self-healing concrete. Constr Build Mater 26(1):532–540
14. Wang JY, Mignon A, Snoeck D, Wiktor V, Van Vlierberghe S, Boon N, De Belie N (2015)
Application of modified-alginate encapsulated carbonate producing bacteria in concrete: a
promising strategy for crack self-healing, Front Microbiol:6
15. Dong Y, Zhang Y, Tu B, Miao J (2014) Immobilization of ammonia-oxidizing bacteria by
calcium alginate. Ecol Eng 73:809–814
16. Seifan M, Samani AK, Hewitt S, Berenjian A (2017) The effect of cell immobilization by
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Nanosilica-Modified Hydrogels Encapsulating Bacterial Spores … 73

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Reusing Alum Sludge as Cement
Replacement to Develop Eco-Friendly
Concrete Products

Y. Liu, Y. Zhuge, and W. Duan

Abstract Alum sludge is a typical by-product of the water industry. The traditional
sludge management method, disposing of sludge in landfill sites, poses a critical envi-
ronmental and economic concern due to a significant increase in sludge amount and
disposal cost. In this paper, the feasibility of reusing sludge as cement replacement
is investigated, and the physical performance and microstructure modification of
concrete products made with sludge is discussed. The obtained results indicated that
a satisfying pozzolanic reactivity of sludge after calcination at high temperatures and
grinding to the appropriate size was identified. When 10% cement was replaced with
sludge, the reaction degree of sludge was up to 39%, and the obtained concrete blocks
exhibited superior mechanical performance. Based on the microstructural analysis,
e.g., x-ray diffraction, thermogravimetric analysis, and advanced nanoindentation
method, the high aluminum content in sludge was incorporated into C–(A)–S–H gel;
the original “Al-minor” C–(A)–S–H gel in pure cement paste was converted to ‘Al-
rich’ C–(A)–S–H gel. Also, sludge promoted the formation of aluminum-bearing
hydrates, such as ettringite and calcium aluminate hydrates (C–A–H). Although the
Al incorporation had no significant effect on the hardness and modulus of C–(A)–
S–H gel, the homogeneous mechanical properties (hardness and modulus measured
with nanoindentation) of binder paste degraded with increasing sludge ash content
above 10%, attributing to the lower hardness of unreacted sludge than cement clinker
and the relatively lower reaction degree. Using sludge in concrete products offers
an economical and environmentally friendly way to dispose of sludge and preserve
diminishing natural resources. Also, the reduction of cement usage may contribute
to achieving carbon neutrality.

Keywords Advanced nanoindentation technique · Cement replacement · Alum

sludge · Green concrete · Supplementary cementitious materials · Value-added
recycling

Y. Liu · Y. Zhuge (B) · W. Duan

UniSA STEM, University of South Australia, Mawson Lakes, SA 5095, Australia
e-mail: Yan.Zhuge@unisa.edu.au

© The Author(s) 2023 75

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_10
76 Y. Liu et al.

1 Introduction

Alum sludge is a typical by-product of the drinking water industry. The hetero-
geneous sludge waste is formed when the aluminum-based coagulant is combined
with suspended solids, dissolved colloids, organic matter, and microorganisms in
raw water. It is estimated that global sludge production has exceeded 10,000 tonnes
per day, and the rapid population growth and economic development may result in
a significant increase in its amount in future decades [1]. In Australia, most sludge
is disposed of at landfill sites (see Fig. 1), which may cause severe environmental
issues because of land wastage and secondary pollution. In view of the transition
toward a circular economy, vast-available sludge should be considered as a resource
with the potential to be valorized instead of a waste.
Most alum sludge has 20–63 wt% Al2 O3 and 17–41 wt% SiO2 [2]. Its alumi-
nosilicate nature makes sludge can be recycled as cement replacement, proposing a
possible solution to reuse sludge in large quantities. Also, reducing cement usage may
contribute to achieving the target of carbon neutrality. Some previous studies have
already investigated the feasibility of alum sludge as cement replacement in concrete
products [3]. In general, raw alum sludge exhibits no pozzolanic reaction, and the high
organic matter in sludge may hinder the cement hydration, resulting in deteriorated
mechanical and durability performance of concrete products [4]. Treating sludge
with high temperatures, ranging from 600 °C to 800 °C, can efficiently improve
the pozzolanic reactivity of sludge due to the fact that crystal phases of silicon
and aluminum were dehydroxylated to form disordered phases with high reactivity
[5]. However, the optimum temperature to activate sludge activity is still controver-
sial. For the performance of sludge-derived concrete products, a moderate cement
replacement (e.g., 10%) with calcined sludge is feasible without compromising the
mechanical and durability properties [6–8].
Based on the above literature, the ideal temperature (between 600 and 800 °C) to
activate the pozzolanic reactivity of alum sludge needs to be clarified. The reaction

Fig. 1 Sludge landfill sites in South Australia

Reusing Alum Sludge as Cement Replacement to Develop Eco-Friendly … 77

degree of sludge, composition of hydration products, and chemo-mechanical proper-

ties of sludge-cement binders need to be in-depth discussed. Therefore, in this study,
the pozzolanic reactivity of alum sludge under different temperatures was assessed
by the strength index test. The reaction degree of sludge was determined by the
selective dissolution method. The chemical composition of sludge-cement binders
was assessed by x-ray diffraction (XRD) and thermogravimetric analysis (TGA).
The chemo-mechanical properties of blended binders, e.g., indentation hardness
and modulus, were investigated by an advanced nanoindentation method. Final, the
strength of concrete blocks with cement replaced with sludge at weight percentages
of 0%, 10%, 20%, and 30% was studied.

2 Methodology

2.1 Materials

Alum sludge is collected from a local water treatment plant in South Australia. First,
raw sludge was crushed and milled to less than 75 µm. Then, milled sludge was
calcined at 600°C, 700°C, and 800°C, respectively. General-purpose cement is used
according to AS 3972. Fine aggregates and coarse aggregates used to manufacture
concrete blocks (CB) were concrete sand and crushed limestone. The particle size
distribution of the sludge under different temperatures was determined by Mastersizer
3000. The chemical composition of sludge and cement was investigated by X-ray
fluorescence.

2.2 Sample Preparation

CB were manufactured with the dry mix method, and the detailed cast procedures
are described in a previous study [9]. The pastes were also cast in plastic molds with
dimensions of 10 × 10 × 10 mm3 , which contained the same water to cement ratio
and sludge content as CB, assisting in the study of hydration products by eliminating
aggregate interference. Table 1 shows the mix design of the CB.

2.3 Experimental Methods

The pozzolanic reactivity of sludge under different temperatures was determined

by the strength index (SAI) test according to ASTM C311. The selective dissolu-
tion method was used to assess the reaction degree of sludge based on a mixture
of ethylenediaminetetraacetic acid–triethanolamine–NaOH [10]. The compressive
78 Y. Liu et al.

Table 1 Mixture design of blocks (kg/m3 )

Samples Sludge Cement Fine aggregate Coarse aggregate SP
CB-0 0 335 962 985 2.0
CB-10 34 301 962 985 2.6
CB-20 68 267 962 985 3.0
CB-30 100 235 962 985 3.3
CB, concrete blocks; SP, superplasticizer

strength of the CB was measured according to AS 4456.4. The load was applied at
a constant rate of 2 kN/s without any shock until failure.
The hydration products in the cement-sludge binder matrix were character-
ized by XRD and TGA, and the hydration reaction of samples was stopped by
ethanol immersion. An advanced nanoindentation technology (coupling conventional
statistic nanoindentation and chemical mapping) was used to characterize the in-situ
chemo-mechanical properties. Detailed sample preparation procedures and the data
analysis method for the nanoindentation test are described in our previous study [10].

3 Results and Discussion

3.1 Material Characterization

Table 2 shows the chemical composition and the particle size distribution of cement
and sludge under different calcination temperatures. The main components in raw
sludge were Al2 O3 , SiO2 , and organic matter, and minor contents of Fe2 O3 , CaO,
and K2 O were also observed. After calcination, most organic matter content was
eliminated, resulting in an increase in the proportions of Al2 O3 and SiO2 . Because the
chemical composition of sludge calcined between 600°C and 800°C was similar, only
the 800°C-treated one is shown in Table 2. It is worth noting that the sum of the Al2 O3 ,
SiO2 , and Fe2 O3 content in calcined sludge was higher than 70%, which satisfied
the composition requirement of natural pozzolan material based on ASTM C618. In
Table 3, the particle size distribution of calcined sludge and cement is shown. The
Blaine fineness of sludge decreased with increasing calcination temperature, which
could be explained by the fact that dehydroxylated particles agglomerate under high
temperatures to produce new porous grains, especially fine particles [11]. The cement
exhibited finer particle size but comparable Blaine fineness to that of calcined sludge.
The results of the SAI test are shown in Table 4. After calcination at 700 °C and 800
°C, sludge exhibited a satisfactory pozzolanic reactivity (SAI ≥ 75%). In contrast,
the 600 °C-treated sludge could not be used as a pozzolan material. Compared with
700 °C, 800 °C is a better temperature to activate the pozzolanic reactivity of sludge,
Reusing Alum Sludge as Cement Replacement to Develop Eco-Friendly … 79

Table 2 Chemical composition (LOI, loss on ignition)

Type Al2 O3 SiO2 Fe2 O3 Na2 O CaO K2 O TiO2 LOI
Raw sludge 28.3 26.4 7.7 – 5.4 1.2 – 29.5
Calcined sludge 47.7 31.1 5.9 – 4.3 – – 6.2
Cement 4.6 20.0 3.1 0.2 63.4 0.4 0.3 3.3

Table 3 Particle size distribution

Cement 600°C-sludge 700°C-sludge 800°C-sludge
D10 (µm) 3.4 5.4 5.8 6.1
D50 (µm) 14.5 29.6 29.2 30.4
D90 (µm) 33.1 61.3 60.3 60.9
Blaine fineness (m2 /kg) 400.0 445.4 429.8 423.4

Table 4 SAI test of sludge

600°C-sludge 700°C-sludge 800°C-sludge
under different calcination
temperatures (SAI, strength 7-day SAI (%) 15.07 61.69 87.00
index) 28-day SAI (%) 14.59 91.38 113.60

in which the SAI value is up to 113.60%. Therefore, 800 °C-treated sludge was used
to cast the paste and CB samples.

3.2 Reaction Degree and Hydration Mechanism of Sludge

The selective dissolution method determined the reaction degree of sludge in the
blended binders. The paste samples containing 10% sludge by weight exhibited
the highest reaction degree, up to 39.0%. Further increasing the sludge content to
20% and 30%, decreased the reaction degree to 25.2% and 24.8%, respectively.
The decrease in reaction degree could be attributed to a lack of sufficient port-
landite. Figure 2 shows the TGA and XRD. In Fig. 2a, c, a significant endothermic
peak occurred ≈ 120º, which was associated with the decomposition of the AFt
phase, (e.g., ettringite). The ettringite peak intensity increased with increasing sludge
content; thus, sludge might promote the formation of AFt phases. Also, adding sludge
resulted in the formation of calcium aluminate hydrate (C–A–H). These results could
be related to the high reactive Al content in sludge, leading to the formation of addi-
tional Al-bearing phases [12]. In Fig. 2b, d, the XRD patterns of pastes at 28 days and
90 days are shown. The reflection peaks related to C–A–H and stratlingite were only
observed in the blended pastes, not the reference ones. The ettringite peak inten-
sity was enhanced with increasing sludge content, which was consistent with the
TGA analysis. In both the TGA and XRD analyses, the content of portlandite (CH)
80 Y. Liu et al.

Fig. 2 a, c TGA and b, d XRD analysis of blended pastes (modified from Liu et al. [13])

decreased with increasing sludge content, confirming the pozzolanic reactivity of

sludge. It is interesting to note that the CH content in the paste with 30% sludge was
significantly low compared with the reference or 10% sludge pastes. Such a lower
concentration of available CH content might result in a decrease in the reaction degree
of sludge.

3.3 Chemo-Mechanical Properties of Pastes and Strength

of Blocks

Table 5 shows the chemo-mechanical properties of the pastes and the strength of CB.
The homogenized modulus of binders significantly decreased when >10% of cement
was replaced with sludge. Such a reduction could be attributed to decreased High-
density (HD) C–S–H gel in samples with 30% sludge. However, the total volume
of C–S–H gel in the different pastes was almost the same, indicating that the filler
effect of sludge might compensate for the cement dilution effect, although the high
amount of sludge hindered the transformation of Low-denisty (LD) C–S–H gel to
Reusing Alum Sludge as Cement Replacement to Develop Eco-Friendly … 81

Table 5 Chemo-mechanical properties of C–A–S–H gel in pastes and the strength of CB

Cementitious pastes Ref 10% sludge 30% sludge
Homogenized modulus (GPa) 35.3 36.7 30.7
HD C–S–H modulus (GPa) 37.4 36.1 34.7
LD C–S–H modulus (GPa) 21.7 22.7 21.0
Relative Al intensity in HD C-S–H 0.14 0.11 0.27
Relative Al intensity in LD C-S–H 0.14 0.13 0.33
HD C–S–H gel volume (%) 0.52 0.55 0.36
LD C–S–H gel volume (%) 0.28 0.24 0.42
Total C–S–H gel volume (%) 0.80 0.79 0.78
Concrete blocks
7-day compressive strength (MPa) 18.2 13.4 9.1
28-day compressive strength (MPa) 20.4 20.3 11.0
90-day compressive strength (MPa) 22.3 25.8 12.7

HD C–(A)–S–H gel. Based on the results for Al intensity in the C–S–H gel, with the
addition of sludge to 30%, the original “Al-minor” C–(A)–S–H gel in pure cement
paste was converted to “Al-rich” C–(A)–S–H gel. There was no significant difference
in indentation modulus for C–(A)–S–H gel in the different pastes, indicating that Al
incorporation had negligible effect on the mechanical properties of C–(A)–S–H gel.
At a curing age of 7 days, the compressive strength of CB containing sludge
was significantly lower than that of the reference samples. However, after curing
for 28 days, the samples with 10% sludge exhibited a comparable or even higher
compressive strength. The optimum sludge content in blocks was 10%, which was
in agreement with the results of the nanoindentation analysis.

4 Conclusions

In general, the feasibility of reusing sludge as cement replacement was confirmed

in this study. 800 °C was the best temperature to activate the pozzolanic reactivity
of sludge. 10% of cement could be replaced with sludge in concrete blocks without
compromising mechanical performance. The reaction degree of sludge in blended
pastes could be up to 39%, and the Al content in C–(A)–S–H gel increased with
adding sludge. However, the C–(A)–S–H gel modification had no significant effect
on the indentation modulus or hardness of the gel. In addition to incorporation into
C–(A)–S–H gel, the high content of reactive Al promoted the formation of other
Al-bearing phases (e.g., ettringite and stratlingite).
82 Y. Liu et al.

References

1. Liu Y et al (2020) Recycling drinking water treatment sludge into eco-concrete blocks with
CO2 curing: Durability and leachability. Sci Total Environ 746
2. Zhuge Y, Liu Y, Pham PN (2022) Sustainable utilization of drinking water sludge. Low carbon
stabilization and solidification of hazardous wastes. Elsevier, pp 303–320
3. Gomes SDC et al (2020) Recycling of raw water treatment sludge in cementitious composites:
effects on heat evolution, compressive strength and microstructure. Resour Conserv Recycl:161
4. Wang L et al (2018) A novel type of controlled low strength material derived from alum sludge
and green materials. Constr Build Mater 165:792–800
5. Tantawy MA (2015) Characterization and pozzolanic properties of calcined alum sludge. Mater
Res Bull 61:415–421
6. Frías M et al (2014) Influence of activated drinking-water treatment waste on binary cement-
based composite behavior: Characterization and properties. Compos B Eng 60:14–20
7. Bohórquez González K et al (2020) Use of sludge ash from drinking water treatment plant in
hydraulic mortars. Mater Today Commun:23
8. Owaid HM, Hamid R, Taha MR (2014) Influence of thermally activated alum sludge ash on the
engineering properties of multiple-blended binders concretes. Constr Build Mater 61:216–229
9. Liu Y et al (2020b) Properties and microstructure of concrete blocks incorporating drinking
water treatment sludge exposed to early-age carbonation curing. J Clean Prod:261
10. Liu Y et al (2021) Cementitious composites containing alum sludge ash: An investigation of
microstructural features by an advanced nanoindentation technology. Constr Build Mater:299
11. Fabbri B, Gualtieri S, Leonardi C (2013) Modifications induced by the thermal treatment of
kaolin and determination of reactivity of metakaolin. Appl Clay Sci 73:2–10
12. Tironi A et al (2013) Assessment of pozzolanic activity of different calcined clays. Cement
Concr Compos 37:319–327
13. Liu Y et al (2021) The potential use of drinking water sludge ash as supplementary cementitious
material in the manufacture of concrete blocks. Resour Conserv Recycl:168

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
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the copyright holder.
Role of Aggregate Reactivity, Binder
Composition, and Curing Temperature
on the Delayed Ettringite Formation
and Associated Durability Loss
in Concrete

L. Martin, P. Thomas, P. De Silva, and V. Sirivivatnanon

Abstract The durability of concrete is critical to its worldwide use as a structural

material for buildings and infrastructure, with the lifetime service of concrete greatly
affecting its economic, environmental, and social costs. Causes of durability loss in
some concrete structures can be attributed to the alkali–silica reaction (ASR) and
delayed ettringite formation (DEF). Both are chemical reactions that have the poten-
tial to cause expansion and strength loss in affected elements. Significant overlap
exists in the factors contributing to ASR and DEF in concrete structures, with widely
reported evidence of deleterious DEF frequently occurring in conjunction with mild
or moderate ASR. For precast concrete, experiments in mortars have provided limits
in the alkali and sulfate content of the binder and maximum curing temperatures used
to minimize DEF risk. The role of other constituents in concrete specimens, notably
the aggregate, has been overlooked. We investigated the role of reactive aggregates
and ASR in the susceptibility of concrete to deleterious DEF.

Keywords Alkali-silica reaction · Delayed ettringite formation · Durability loss ·

Heat cure · Precast concrete

L. Martin (B)
School of Mathematical and Physical Sciences, University of Technology Sydney, Sydney,
NSW, Australia
e-mail: liam.martin@uts.edu.au
P. Thomas · V. Sirivivatnanon
School of Civil and Environmental Engineering, University of Technology Sydney, Sydney,
NSW, Australia
P. De Silva
School of Behavioural and Health Sciences, Australian Catholic University, Sydney, NSW,
Australia

© The Author(s) 2023 83

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_11
84 L. Martin et al.

1 Introduction

The durability of concrete as a structural material is vitally important to its modern

use in buildings, homes, and infrastructure around the world. Concrete and cement-
based materials have many desirable properties for use in a variety of structures,
allowing in-form construction, low labor cost, low technical threshold, low cost
and environmental concerns (excluding high carbon emissions), high strength, good
durability, and long-term efficacy [1, 2]. However, the extended service life, multi-
faceted application, and complex nature of concrete as a composite material creates
many potential causes of durability loss across the structural, mechanical, and chem-
ical fields. These include acid, chloride, or sulfate attack, steel corrosion, leeching,
freeze–thaw events, salt and de-icing ingress, humidity cycles, and erosion [3, 4].
Two potential causes of durability loss in concrete structures are the alkali–silica
reaction (ASR) and delayed ettringite formation (DEF), both of which are chem-
ical reactions that can lead to deleterious expansion, microcracking, and strength
loss in affected elements. DEF is a form of internal sulfate attack in concrete, with
the primary mechanism being the dissolution–precipitation reaction of the sulfate
mineral ettringite, which can cause expansion and cracking [5]. DEF is of most
concern in the precast concrete industry, where the conditions used to accelerate the
strength gain of concrete have the potential to increase the risk of deleterious DEF
[6]. The ASR is a chemical reaction in concrete involving the alkaline pore solution
and aggregate material containing non-crystalline silica, leading to the formation of
an expansive ASR–gel phase and the potential for expansion and cracking [7]. In
many real-world instances of DEF, ASR was also found in [8–10]. In an Australian
context, all reported cases of DEF in concrete have been in conjunction with mild
or moderate ASR, and have been linked to cracking in steam-cured precast concrete
and massive concrete elements [6]. Thus, manufacturing specifications for precast
and/or large concrete have a conservative view to DEF risk, leading to restrictive
processing methods and substantial industry costs [6, 11].
The DEF mechanism is based on the solubility of sulfate in the pore solution of
concrete. Ettringite normally forms during early cement hydration, with no negative
effects on the final concrete material. Elevated temperatures and alkalinity (high pH)
result in the decomposition of ettringite, and when the concrete system returns to
normal conditions of temperature and alkalinity, ettringite will recrystallize [12]. The
slow precipitation of ettringite as an expansive phase in hardened concrete at later
ages induces cracking and strength loss. To minimize the risk of deleterious DEF,
temperature limits are used in the manufacturing of precast concrete, as based on
laboratory mortar tests and chemical composition of the cement [6, 13]. The role of
other components of concrete, specifically the aggregate, has been downplayed.
Role of Aggregate Reactivity, Binder Composition, and Curing … 85

1.1 Project Outline

Available literature and research results concerning DEF have focused on labora-
tory experiments in mortars with modified cement systems. Guidelines and speci-
fications for cement production and the design of concrete structures are based on
these findings by convention. The consideration of additional factors found in true
concrete elements, including the potential for ASR, has been overlooked. Evidence-
based contributions to this topic will assist in the development of targeted guide-
lines and standards for the precast concrete industry and mitigation of DEF risk in
modern structures. By increasing the confidence in the best use of construction mate-
rials, our research will help reduce the environmental, economic, and social costs of
steam-cured concrete elements and the risk of catastrophic failure.
The aim of this research work was to investigate the process of DEF and the role
of aggregate reactivity, binder composition, and curing temperature on the suscep-
tibility of concrete systems to deleterious DEF and ASR–DEF. The primary objec-
tives of this study were to identify the conditions in concrete containing nonreactive
aggregates that can induce deleterious DEF, specifically related to cement composi-
tion and curing conditions, and to investigate the role of reactive aggregates on the
susceptibility of concrete to deleterious DEF.

2 Methods

The manufacture of concrete specimens for this study was carried out using local
sourced materials, complying with relevant Australian standards and industry guide-
lines. The binder used was an Australian produced, commercial grade general-
purpose (GP) cement. The elemental oxide composition of the cement was
determined by X-ray fluorescence (XRF) analysis, with results presented in Table 1.
Potable tap water was used for mixing of concrete and preparation of saturated
limewater. To retain sufficient workability of freshly mixed concrete, an alkali-free
superplasticizer was used (125 g/100 L).
Previous literature has reported a pessimum (worst case) condition for binder
compositions of alkali and sulfate content with regards to deleterious DEF [13, 14].
For pessimum concrete specimens, the alkali and sulfate binder content was increased
to 1% Na2 Oeq (4.5 kg/m3 ) and 4% SO3 (18.0 kg/m3 ) by the addition of dissolved
sodium hydroxide and powdered calcium sulfate dihydrate respectively. Aggregate
materials used were selected according to ASR-reactivity, as classified by Australian

Table 1 Chemical composition of GP cement via XRF

Component L.O.I CaO SiO2 Al2 O3 Fe2 O3 MgO SO3 Na2 O K2 O Cl Total Na2 Oeq
Content 4.0 64.1 19.4 4.9 3.0 1.2 2.6 0.17 0.45 0.037 100.1 0.47
(w/w %)
86 L. Martin et al.

Table 2 Concrete mix design

Property Cement content Coarse Fine aggregate w/c ratio Batch size Slump
aggregate
Detail 450 kg/m3 1190 kg/ 640 kg/m3 0.40 100 kg 120 mm
m3

standards AS1141.60.2 for coarse aggregates and AS1141.60.1 for fine aggregates.
Coarse material was either a nonreactive basalt (nRe) or reactive dacite (cRe), with a
grading of 20 and 10 mm in a 3:1 ratio. The fine material was a washed nonreactive
river sand. Concrete prisms (285 × 70 × 70 mm) were manufactured using AS1012.2
as the guide for preparation and mixing, and AS1141.60.2 for the design of prisms,
with three prisms for each set of results. Concrete cylinders (100 × 200 mm) were also
manufactured, with two cylinders for each set. The concrete mix design, presented
in Table 2, was based on a typical large structural element manufactured by the
Australian precast industry, utilizing a high cement content and low water content.
Concrete specimens were cured under one of two conditions: curing at ambient
temperature (23 ± 2 °C) or heat-cured. The heat-curing process was designed to
follow the internal temperature profile of a large precast element [11]; preset of
30 °C for 4 h, heating of 30 °C/h up to 90 °C, soak at 90 °C for 12 h, then cooling
to ambient temperature and demolding. During curing, all specimens were stored
in sealed plastic bags, with a damp cloth as the moisture source. After demolding,
specimens were stored in limewater tanks at ambient temperature. The compressive
strength of the concrete specimens was measured at 1 day and 28 days with cylinders,
and at 1 year with cubes cut from the prisms (75 mm, 2 cubes), as per AS1012.9.
Concrete prisms were measured for linear length and mass, using AS1141.60.2 as
the guide, at day 1, day 7 as reference, day 28 and then at monthly intervals up to
1 year.

3 Results and Discussion

Concrete specimens were manufactured as detailed above, with varying binder

(pessimum/1N 4$ or control/ctrl), aggregate (reactive/cRe or nonreactive/nRe), and
curing (ambient/amb or heat-cured/heat) conditions linked to the occurrence of ASR,
DEF, and ASR–DEF. Linear expansion and mass of prisms was monitored over 1 year,
and compressive strength was measured at 1 day, 28 days, and 1 year.
“DEF-only” systems had only nRe aggregates present and thus minimal potential
for ASR; “ASR–DEF” systems had cRe aggregates present and thus ASR could
occur; “pessimum” systems had an altered binder of 1.00% Na2 Oeq and 4.0% SO3 ,
which increased the risk of DEF; “Control” systems used the binder as received,
complying with the standards and chemical content limits of Australia.
The threshold for deleterious expansion was arbitrarily set at 0.03% of total linear
length, based on the Australian aggregate reactivity test method AS1141.60.2. Note
Role of Aggregate Reactivity, Binder Composition, and Curing … 87

that the specimens were treated differently in this study compared with AS1141.60.2,
specifically being immersed in limewater tanks at normal temperature (23 °C) and
lower alkali binder content (1% Na2 Oeq ).

3.1 DEF in Concrete

Concrete specimens were manufactured with selected binder and curing conditions
to promote deleterious DEF in the absence of potential ASR. Linear expansion, mass
gain, and compressive strength were measured over 1 year, as shown in Fig. 1.
Deleterious expansion and strength loss were observed only in specimens subject
to pessimum binder content and sustained heat-curing, and they were attributed to
DEF. All other concrete systems, with either curing at ambient temperature or as-
received cement, did not show deleterious effects or durability loss. This supports the

Fig. 1 Physical characteristics of concrete prisms with nonreactive (nRe) aggregates, showing
a linear expansion; b mass gain and c compressive strength, over 1 year. Binder used was at
pessimum condition (1% alkali, 4% sulfate/1N 4$) or as received (control/ctrl), curing was at
ambient temperature (23 °C/amb) or heat-cured (90 °C for 12 h/heat)
88 L. Martin et al.

current understanding that high temperature, elevated alkali, and elevated sulfate are
all essential factors for the occurrence of deleterious DEF in concrete. For structures
without any of these factors present, the risk of durability loss due to DEF is predicted
to be minimal.
Observed expansion in the heat-cured pessimum prisms was significant, reaching
2% total of original length at 1 year, with expansion slowed but ongoing in the
final months of measurement. It is expected that ongoing change would plateau in a
further 6 months. This overall expansion was severe but comparable to other reported
results of deleterious DEF, ranging from 1.2 to 1.8% [13]. For all other systems, the
expansion was below the 0.03% threshold at 1 year, with decreasing expansion of
pessimum/ambient, control/heat-cured, and control/ambient as the least. These minor
variations were attributed to increased porosity of heat-cured cement paste [6, 14],
and the sulfate addition inducing limited early expansion [4].
Changes in the mass of the concrete specimens closely followed that of expansion,
with only the pessimum/heat-cured system showing a large expansion of 2.5% at
1 year. All other concrete systems only had minor increases to the total mass after
1 year. This large change was attributed to two causes: the inflection of the expansion
curve indicating rapid development of new phases, including ettringite, which retains
significant amounts of water in bound hydrates, and the formation of microcracks
across the bulk material from internal expansive stress leading to ingress and uptake
of water.
Similarly with the strength of the concrete specimens, significant strength loss was
observed only in pessimum/heat-cured systems and attributed to DEF, with measured
strength at day 28 to 1 year decreasing from 45 MPa to 20.5 MPa, respectively. Again,
all other concrete systems showed strength gain as expected. For day-1 strength, heat-
cured specimens were equivalent to or better than the corresponding ambient-cured
specimens, as expected with precast manufacturing of concrete [15]

3.2 ASR–DEF in Concrete

Concrete specimens containing cRe aggregates were manufactured with selected

binder and curing conditions to promote deleterious DEF, with the potential of ASR.
Linear expansion, mass gain, and compressive strength were measured over 1 year,
shown in Fig. 2.
Measured linear expansion, mass gain, and compressive strength of the ASR–DEF
systems followed the same pattern as that of the DEF-only systems, with concrete
systems with pessimum conditions and heat-curing exhibiting deleterious results
over 1 year, while all other systems did not. For the pessimum/heat system, observed
results at 1 year were expansion of 1.4%, mass gain of 2.0%, and compressive
strength of 43.0 MPa at 28 days to 32.0 MPa at 1 year. The exception to this was
minor expansion in the control/heat-cured system, reaching 0.03% at 1 year, with no
corresponding change in mass or strength. This was attributed to increased porosity
and alkali release of the cement paste due to the heat-curing process, leading to ASR.
Role of Aggregate Reactivity, Binder Composition, and Curing … 89

Fig. 2 Physical characteristics of concrete prisms with coarse reactive (cRe) aggregates, showing
a linear expansion; b mass gain; and c compressive strength, over 1 year. Binder used was at
pessimum condition (1% alkali, 4% sulfate/1N 4$) or as received (control/ctrl), curing was at
ambient temperature (23 °C/amb) or heat-cured (12 h at 90 °C/heat)

Comparing the deleterious DEF-only and ASR–DEF systems, the most signifi-
cant difference was the reduced magnitude of durability loss in the concrete speci-
mens. The pessimum/heat-cured DEF-only system had more expansion, mass gain,
and strength loss than the equivalent system with cRe aggregates present at 1 year.
One possible cause is the development of ASR and related phases slowing the DEF
process and reducing the total amount of microcracking and expansive phases. In this
scenario, reactive silica and small amounts of ASR–gel could act as a weak supple-
mentary cementitious material (SCM), blocking transport pores, dispersing ettringite
away from the aggregate–paste boundary, and creating micro-voids for ettringite to
form in without restraint. Supporting evidence can be found in the efficacy of SCMs
such as fly ash to mitigate DEF [11], and the use of finely ground reactive aggre-
gates as replacement SCMs [16]. Otherwise, the trajectory of change in expansion
remained the same, with high early expansion (75–150 days), divergence (200 days),
and reduced rate in the later months (300–350 days).
90 L. Martin et al.

4 Conclusions

This research investigated the role of the ASR in the susceptibility of concrete to
DEF via an experimental study with concrete specimens. The factors of alkali,
sulfate, temperature, and aggregate reactivity were assessed for their contribution
to deleterious DEF and ASR–DEF in concrete.
Deleterious DEF was not observed in concrete specimens prepared with locally
produced Australian cement, which is linked to the low alkali and low sulfate binder
content. Binder systems with elevated alkali and sulfate content increase the risk of
deleterious DEF. The combination of binder composition with high alkali (1.00%
Na2 Oeq ) and high sulfate (SO3 %) content, and curing conditions of high temperature
(90 °C) and sustained heat (12 h) are necessary for deleterious DEF to occur in
concrete. Specimens not subjected to these conditions did not exhibit deleterious
expansion during 1 year of measurement.
Reactivity of aggregates influences expansion attributed to DEF, but only in the
presence of pessimum conditions and heat-curing. Observed durability loss with
expansion, mass gain, and strength loss was greater in DEF-only systems compared
with ASR–DEF systems.

Acknowledgements This research was conducted at the University of Technology Sydney and
supported by an Australian Government Research Training Program Scholarship. It was funded
through an Australian Research Council Research Hub for Nanoscience Based Construction Mate-
rials Manufacturing (NANOCOMM) (IH150100006) with the support of Cement Concrete and
Aggregates Australia (CCAA).

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Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
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the copyright holder.
Effect of Blending Alum Sludge
and Ground Granulated Blast-Furnace
Slag as Cement Replacement to Mitigate
Alkali-Silica Reaction

W. Duan, Y. Zhuge, and Y. Liu

Abstract The alkali–silica reaction (ASR) is a severe durability issue in cement-

based materials. Although using calcium-rich supplementary cementitious mate-
rials (SCMs) such as ground granulated blast-furnace slag (GGBS) is beneficial
for improving mechanical performance, it can lead to critical ASR-induced damage,
primarily when high-reactive aggregates are used. We used alum sludge, a byproduct
of drinking water treatment processes, and found it to have high efficiency in miti-
gating ASR in mortars containing GGBS as cement replacement and waste glass as
high-reactive aggregate. The raw alum sludge was calcined for 2 h at 800 °C and
ground to pass a 75-μm sieve. Ternary blended binders were made by replacing
10, 20 and 30% of cement with the mixture of alum sludge and GGBS (ratio 1:1).
The mortar samples exhibited a considerable compressive strength and significant
ASR resistance when 30% of cement was replaced with the mixture of alum sludge
and GGBS compared with the reference samples. Microstructural characterization
using X-ray diffraction, backscattered electron images and energy-dispersive X-ray
spectroscopy indicated that increasing the aluminum content of the alum sludge
could prevent the formation of detrimental Ca-rich and low-flowable ASR gels. The
hindering effect was attributed to the alkaline binding ability and the extra precip-
itation of calcium aluminum silicate hydrate phases due to the abundant Al in the
binder.

Keywords Alkali–silica reaction · Alum sludge · Ground granulated blast furnace

slag · Supplementary cementitious materials · Waste glass aggregates ·
Value-added recycling

W. Duan · Y. Zhuge (B) · Y. Liu

UniSA STEM, University of South Australia, Adelaide, SA, Australia
e-mail: Yan.Zhuge@unisa.edu.au

© The Author(s) 2023 93

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_12
94 W. Duan et al.

1 Introduction

The alkali–silica reaction (ASR) is one of the most severe durability issues in Portland
cement-based materials. Its occurrence in concrete products can cause significant
structural damage and shorten service life. ASR initializes when the reactive (also
known as amorphous) silica from the aggregates is dissolved by the alkaline from
cement [1]. Although using nonreactive aggregates can be the most effective strategy
to prevent ASR, the transportation cost in some locations due to less availability
of these aggregates can be significantly unprofitable. In addition, re-using waste
materials as aggregates, such as crushed glass, which has a high content of amorphous
silica, has been of great interest for decades because it provides an environmentally
friendly solution to disposal of the wastes [2]. Therefore, an economic ASR control
method is necessary to maximize the environmental benefit.
Adding supplementary cementitious materials (SCMs) is a feasible method of
mitigating the ASR. Previous studies have reported the ASR mitigation effects
attributed to SCMs, including metakaolin, coal fly ash (FA) and ground granu-
lated blast-furnace slag (GGBS) [3, 4]. Alum sludge is a byproduct of drinking
water treatment and because alum-based coagulant is usually added to raw water to
remove insoluble particles such as sand and microorganisms, the primary chemical
composition of alum sludge includes aluminum, silicon and organic compounds.
Therefore, alum sludge could be a SCM. Previous studies [5–10] have demonstrated
the successful utilization of alum sludge in the manufacture of mortar and concrete
blocks. The results indicated that up to 10% cement replacement with calcined
and milled alum sludge could improve mechanical performance, but a strength
reduction would occur when higher proportions of cement were replaced. Because
GGBS-incorporated concrete products exhibit considerable strength even with high
volume cement replacement [11], blending GGBS with alum sludge could potentially
improve mechanical performance further when more than 10% cement is replaced.
In addition, a ternary blended system is expected to have satisfactory ASR resistance
due to the high Al content of the sludge.
In the present study, mortars with ternary blended binders containing GGBS
and alum sludge were developed to achieve considerable mechanical performance
and ASR resistance. Compressive strength, the ASR mitigation performance and
microstructural characteristics were investigated for different mixtures of GGBS
and alum sludge.
Effect of Blending Alum Sludge and Ground Granulated Blast-Furnace … 95

CaO
Cement
(a) 10 (b)
80
GGBS
Cement
GGBS 60 ASA
8 ASA
Glass 40
80 SiO2
Volume (%)

6 Others 60
80 20 40
60 20
40
20
4
20
20
40
2
40
60
60
0 80
0.1 1 10 100 1000 80
Particle size (µm)
Fe2O3 Al2O3

Fig. 1 a Particle size of cement, alum sludge ash (ASA), granulated blast-furnace slag (GGBS)
and glass aggregate; b chemical composition of cement, ASA and GGBS

2 Methods

2.1 Materials

General-purpose cement was used as the binder according to AS3972 [12] and clear
crushed glass, which contains high levels of amorphous silica, was used as fine
aggregate to accelerate the ASR. The alum sludge was supplied by a drinking water
treatment plant located in South Australia. The raw sludge was calcined at 800 °C
for 2 h, and then the alum sludge ash (ASA) was milled to pass a 75-μm sieve. The
particle size distributions of the crushed glass, ASA and GGBS are shown in Fig. 1a.
The chemical composition of the calcined alum sludge was analyzed using X-ray
fluorescence (XRF), and the results were compared with those for cement and GGBS
(Fig. 1b).

2.2 Experimental Protocol

2.2.1 Sample Preparation

The mix proportions of mortars were designed according to AS1141.60.1 [13], and
shown in Table 1. The water to cement and aggregate to cement ratios were 0.47
and 2.25, respectively. All mortar samples except those for the ASR tests were cured
in a chamber with temperature and relative humidity controlled at 23 °C and 95%,
respectively, for 28 days before testing.
96 W. Duan et al.

Table 1 Mixture of the mortar samples (kg/m3 )

Sample Cement ASA GGBS Glass Sample Cement ASA GGBS Glass
R0 590 0 0 1328 G20 472 0 118 1328
A10 531 59 0 1328 G30 413 0 177 1328
A20 472 118 0 1328 AG10 531 30 30 1328
A30 413 177 0 1328 AG20 472 59 59 1328
G10 531 0 59 1328 AG30 413 89 89 1328
ASA, alum sludge ash; GGBS, granulated blast-furnace slag

2.2.2 Compressive Strength and Accelerated ASR Tests

The compressive strength was tested according to AS4456.4 [14], and the loading rate
was 0.33 MPa/s. The ASR resistance of the mortar samples was evaluated according
to AS1141.60.1 [13]. The demolded samples were cured in 80 °C water for 24 h. The
initial length of the mortar beams was determined at the end of curing and then the
samples were immersed in 1 M NaOH solution in a water bath with the temperature
set to 80 °C.

2.2.3 Microstructural Analysis

The microstructural characteristics of the samples after the ASR attack were observed
using backscattered electron (BSE) micrographs, and the elemental analysis of the
cement matrix was evaluated using energy-dispersive X-ray spectroscopy (EDS) with
accelerating voltage at 15 kV. X-ray diffraction (XRD) patterns were obtained using
copper Kα radiation at 40 kV and 40 mA.

3 Results and Discussion

3.1 Compressive Strength

The 28-day compressive strength of the control sample was 34 MPa, and the
percentage of strength for other samples relative to the control is shown in Fig. 2. For
the binder with binary blends (ASA and GGBS group), 10% cement replacement
improved the mechanical performance, but greater than this value, a strength reduc-
tion was observed. In contrast, for the ternary blended binders (containing both ASA
and GGBS, labelled as AG group), the mortar with 30% cement replacement still
had a considerable compressive strength compared with the reference, which was
attributed to the extra pozzolanic reaction from the synergy of ASA and GGBS. The
Effect of Blending Alum Sludge and Ground Granulated Blast-Furnace … 97

Fig. 2 Compressive strength 150

relative to the control Cement Replacement:
10% 20% 30%

Relative compressive strength (%)

sample. AG, ASA and
GGBS; ASA, alum sludge 117
112
ash; GGBS, granulated Reference 109 108 109
blast-furnace slag 100 98
94
90
84

50

0
ASA GGBS AG

excessive portlandite (CH) from the GGBS could react with the silica species in the
ASA, contributing to higher mechanical performance at a higher cement replacement
level than with the binary blends.

3.2 ASR-Induced Expansion

The results of ASR-induced expansion and the surface cracking observed in samples
are shown in Fig. 3. Although 30% cement replacement with GGBS kept expansion
less than 0.3% for 21 days, the 14-day expansion exceeded the threshold of 0.1%. As
a comparison, 20% ASA content effectively prevented the ASR. GGBS exhibited
a negligible ASR mitigation effect than ASA for the mortar with binary blended
binders due to the higher Ca content in GGBS, which had the potential to promote
ASR-induced expansion. In addition, high Al content in ASA was beneficial for
binding alkaline. The analysis of the effect of Ca and Al on the ASR is discussed in
Sect. 3.4. For the samples in the AG group, 20% cement replaced with a mix of ASA
and GGBS suppressed ASR-induced expansion to less than 0.3%, and 30% cement
replacement significantly mitigated the ASR. The expansion results were consistent
with the surface visual observation of the samples shown in Fig. 3b, where tree-like
cracks were found in the reference (R0) sample, and no obvious crack could be
identified in AG30.

3.3 Phase Analysis of the Mortars

The XRD spectra of the samples before and after the ASR test are shown in Fig. 4. No
CH peak can be found for sample A30 before the ASR test, which indicated that the
98 W. Duan et al.

Fig. 3 a ASR-induced expansion; b surface visual observation of the mortars. ASA, alum sludge
ash; ASR, alkali–silica reaction; GGBS, granulated blast-furnace slag

CH from cement hydration participated in the pozzolanic reaction and was consumed
by the excess ASA. However, the amount of CH may not be sufficient to generate
considerable pozzolanic C-(A)-S–H. Thus, the cement dilution effect dominated the
mechanical properties, causing a lower strength than the reference.
Incorporating GGBS into the ASA provided an additional CH source and ensured a
high degree of pozzolanic reaction. The cement dilution effect could be overwhelmed
by the generation of pozzolanic products, which contributed to the better mechanical
performance of the mortars in the AG group than those in the ASA and GGBS groups.

Fig. 4 X-ray diffraction spectra of the samples before and after ASR testing. ASR, alkali–silica
reaction
Effect of Blending Alum Sludge and Ground Granulated Blast-Furnace … 99

Fig. 5 BSE-SEM images of a R0, b A10 and c AG30 samples. BSE, backscattered electron; SEM,
scanning electron microscopy

After the ASR test, kaoite (C3 AH6 ) was detected as a new phase in the samples with
ASA content. Kaoite is usually precipitated at a high Al/Si ratio, accompanying C-A-
S-H formation, especially in high-alkaline solution [15]. The later formed C-A-S–H
gels were beneficial for mitigating ASR, attributed to their alkaline binding ability.

3.4 Microstructural Characteristics and Elemental Analysis

Figure 5 shows the BSE images of the R0, A10 and AG30 samples after the ASR
attack. ASR gels could be identified in two locations: interior glass aggregate labeled
as T1 and surrounding the glass aggregate labeled as T2. Both T1 and T2 were found
in R0 and A10, but only T2 was detected in AG30. The major cracks in R0 were
collinear to the T1 gels, indicating that the growth of the ASR gels’ interior aggregates
may be the main factor in cement matrix damage and sample expansion.
The elemental analysis of T1 and T2 was performed using EDS technology to
obtain the atomic ratios of Ca, Al, Si and Na in the ASR gels. The results listed in
Table 2 show that the Ca/Si ratio of the ASR gels surrounding the aggregates (T2) was
significantly higher than that inside the aggregates (T1) for the R0 and A10 samples.
Although the initially formed ASR gels had a low Ca content, the gels surrounding
the aggregates could be more able to absorb Ca2+ than the interior ASR gels and then
transform into Ca-rich ASR gels [16]. Incorporating Ca into ASR gels could increase
their viscosity, making them difficult to transport in pores. In addition, further taking
up Ca from the pore solution could transform ASR gels into C-S-H [17]. The Ca-rich
ASR gels and C-S-H layer surrounding the aggregates will limit the extrusion of the
gels inside the aggregates but cannot prevent alkaline transportation into T1 and ASR
gel precipitation in T1. The growth of the constrained T1 ASR gels finally cracked
the aggregates and cement matrix.
100 W. Duan et al.

Table 2 Atomic ratios in the ASR gel of samples R0, A10 and AG30
Sample T1 T2
Na/Si Ca/Si Al/Si Na/Si Ca/Si Al/Si
R0 0.40 0.28 0.02 0.21 0.58 0.03
A10 0.35 0.24 0.02 0.29 0.43 0.01
AG30 − 0.44 0.28 0.07
ASR, alkali–silica reaction

Fig. 6 Energy-dispersive spectrometry mapping of sample AG30

The Ca/Si ratios of the T2 ASR gels in AG20 were much lower than in R0 and
A10 and similar to those of T1 in R0 and A10. ASR gels in AG30 were more flowable
due to their lower Ca content. The possible reason for this result is that the Al from
the ASA prohibited Ca absorption by the ASR gels, thus eliminating the constrain
effect. The Al layer surrounding the aggregate observed in the EDS mapping shown
in Fig. 1 could be evidence for this explanation (Fig. 6).

4 Conclusions

We developed a mortar with a ternary blended binder containing ASA and GGBS.
This newly developed mortar exhibited satisfactory mechanical performance and
excellent ASR resistance even with high reactive waste glass aggregates. Based on
the results, the following conclusion can be drawn:
• The mechanical performance of mortars can be improved using a mix of ASA and
GGBS as a cement replacement. The mortar with 20% cement replaced in the AG
group exhibited the highest compressive strength of 40 MPa, and 30% cement
replacement resulted in a considerable strength compared with the reference.
Effect of Blending Alum Sludge and Ground Granulated Blast-Furnace … 101

• ASA had a more substantial ASR mitigating effect than GGBS when used as the
SCM and could effectively mitigate ASR. 10% ASA content in the binder could
restrain ASR-induced expansion in mortars within 0.3% in the 21-day test. For the
mortars with ternary blended binder, 30% cement replacement could successfully
suppress ASR.
• The Al content of ASA played a significant role in mitigating ASR. The XRD
result indicated the incorporation of Al in the C-S–H structure, which would limit
silica dissolution. The microstructural and elemental analyses also suggested that
ASA could prevent ASR gels absorbing Ca, eliminate the restriction effect of
ASR gels surrounding the aggregate and impede cracking caused by the growth
of ASR gels inside aggregates.

References

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12. AS, AS 3972 (2010) General purpose and blended cements
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the copyright holder.
Optimisation of Limestone Calcined Clay
Cement Based on Response Surface
Method

G. Huang, Y. Zhuge, T. Benn, and Y. Liu

Abstract Limestone calcined clay cement (LC3) is a new type of cement that
contains Portland cement, calcined clay, and limestone. Compared with traditional
cement clinker, LC3 reduces CO2 emissions by up to 40%, and is a promising tech-
nology for the cement industry to achieve its emission target. We used a numer-
ical approach to predict the optimum composition of LC3 mortar. The experiments
were performed using central composite rotational design under the response surface
methodology. The method combined the design of mixtures and multi-response
statistical optimization, in which the 28-day compressive strength was maximized
while the CO2 emissions and materials cost were simultaneously minimized. The
model with a nonsignificant lack of fit and a high coefficient of determination (R2 )
revealed a well fit and adequacy of the quadratic regression model to predict the
performance of LC3 mixtures. An optimum LC3 mixture can be achieved with 43.4%
general purpose cement, 34.16% calcined clay, 20.6% limestone and 1.94% gypsum.

Keywords Limestone calcined clay cement · Optimum LC3 mixture · Response

surface method · Value-added recycling

1 Introduction

Concrete is the most widely used construction material, and as an essential compo-
nent, Portland cement (PC) production accounts for approximately 8% of annual
anthropogenic greenhouse gas emissions [1]. Because there is no viable economic
alternative to concrete, cement consumption is ever-growing. Currently, using supple-
mentary cementitious materials (SCMs) to achieve a partial clinker replacement is
the most effective strategy to reduce cement production. Nevertheless, the shortage
of traditional SCMs (e.g., fly ash and slag) requires exploration of other types of
cementitious materials [2].

G. Huang · Y. Zhuge (B) · T. Benn · Y. Liu

UniSA STEM, University of South Australia, Adelaide, SA, Australia
e-mail: Yan.Zhuge@unisa.edu.au

© The Author(s) 2023 103

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_13
104 G. Huang et al.

Recently, researchers have worked on limestone calcined clay (LC3) cement,

which consists of PC, calcined clay (metakaolin: MK), and limestone (LS). Such
a combination of ingredients allows higher clinker replacement and leads to the
development of a denser microstructure that gives LC3 excellent mechanical and
durability performance [3]. Because LC3 cement has various components, the single-
factor method is mainly used in the design of material proportions in LC3 cement,
but this can only obtain an optimal mix design for a single performance and cannot
achieve multi-objective optimization. In comparison, the response surface method
(RSM) is especially advantageous for designing a competitive and more sustainable
blended cement binder [4].
In this study, central composite rotation design (CCRD) in RSM was performed.
The content of MK, LS and gypsum (GYP) was taken as the variables. The 28-
day compressive strength, CO2 emission, and materials cost were used as response
values. The mathematical model was established by analysis of variance (ANOVA)
to determine the interaction of various variables and model accuracy. The multi-
objective optimization of LC3 mortar was achieved using the numerical-functional
optimization model.

2 Methods

2.1 Materials

LS, MK, GYP, and general-purpose cement (GPC) were used to create the LC3 binder
mixtures (Table 1). MK was obtained after calcination of kaolin clay at 800 °C for
2 h. The LS, MK and GYP were ground using a horizontal roller mill. Concrete sand
with a particle size ranging from 75 to 2.36 mm and a specific density of 2.64 was
collected from ResourceCO Australia. A polycarboxylate-based superplasticizer was
used to maintain the same consistency for all mixtures (between 185 and 195 mm).
Mortar samples with a binder to aggregate ratio of 1:2.75 were cast in accordance
with ASTM C305.

2.2 Design of Mixtures

CCRD is the most suitable technique in RSM to obtain a highly effective mathe-
matical experiment design and develop a functional relationship between the vari-
ables and responses [5]. For the binder design of LC3 cement, MK content (x1 ), LS
content (x2 ) and GYP content (x3 ) were selected as the independent variables. The
PC replaced by other ingredients was by mass and defined as a weight percentage
of the total binder mass. The variable levels were selected to vary from 0 to 60%
Table 1 Chemical and physical characteristics of the binder materials
Binder Chemical composition (%) Physical characteristic
SiO2 Al2 O3 Fe2 O3 CaO MgO Na2 O K2 O SO3 TiO2 P2 O5 SrO d50 (μm)
GPC 20 4.6 3.1 63.41 1.6 0.19 0.37 2.6 0.3 0.1 0.1 17
LS 5.5 0.8 0.7 50.8 0.8 0.11 0.19 0.1 0.1 < 0.1 < 0.1 22.2
MK 71.18 25.99 0.42 0.31 0.15 0.11 0.04 0.08 1.38 0.06 0.01 18.3
GYP 3.9 1 0.4 30.8 0.3 0.15 0.16 41.5 0.1 < 0.1 0.4 25.4
GPC, general-purpose cement; GYP, gypsum; MK, metakaolin; LS, limestone
Optimisation of Limestone Calcined Clay Cement Based on Response …
105
106 G. Huang et al.

Fig. 1 Schematic illustration of central composite rotational design

for MK, 0% to 30% for LS and 0% to 5% for GYP. The effect of each independent
variable was assessed at five levels with a code value of –alpha, –1, 0, 1, + alpha
(factorial, axial and central points). The spatial schematic illustration of CCRD is
shown in Fig. 1. Three replicates are considered at the central point, resulting in 17
experimental runs. The mix proportions are shown in in Table 2.
The compressive strength of the mortar sample at 28 days (Y1 ), CO2 emis-
sions (Y2 ), and materials cost (Y3 ) were taken as the response variables. The 28-
day compressive strength was recorded experimentally, while the environmental
responses, such as CO2 emissions and materials cost, were evaluated based on per
kg of mortar sample. The relevant information on CO2 emission and the material cost
was collected from previous publications or estimations of local materials’ market
prices. All data used are shown in Table 3. The cost of mortar casting procedures
and the emission from mortar service life were not taken into account.
The Design-Expert 12.0.3.0 (Sat-Ease Inc., Minneapolis, MN, USA) software
was used to develop the mathematical design and ANOVA of the experiments. The
quadratic polynomial model was used to predict the optimal conditions, where Y is
the response value, β is the regression coefficient, ε is the random error, Xi and Xj
are the independent variables, and k is the number of variables.


k 
k 
k 
k
Y = β0 + βi xi + βii xi2 + βij Xi Xj + ε (1)
i=1 i=1 i=1 j>1
Table 2 Central composite design matrix and LC3 mortar mix proportions
Run no Coded values Level of variables (%) Mixture proportion (kg/m3)
x1 x2 x3 MK LS GYP MK LS GYP GPC Sand Water SP
1 −1 −1 −1 12.16 6.08 1.01 7.23 3.61 0.60 48.00 163.44 28.78 0.23
2 1 −1 −1 47.85 6.08 1.01 28.44 3.61 0.60 26.79 163.44 28.78 0.40
3 −1 1 −1 12.16 23.92 1.01 7.23 14.22 0.60 37.39 163.44 28.78 0.26
4 1 1 −1 47.85 23.92 1.01 28.44 14.22 0.60 16.18 163.44 28.78 0.37
5 −1 −1 1 12.16 6.08 3.99 7.23 3.61 2.37 46.23 163.44 28.78 0.27
6 1 −1 1 47.85 6.08 3.99 28.44 3.61 2.37 25.02 163.44 28.78 0.46
7 −1 1 1 12.16 23.92 3.99 7.23 14.22 2.37 35.62 163.44 28.78 0.28
8 1 1 1 47.85 23.92 3.99 28.44 14.22 2.37 14.41 163.44 28.78 0.61
9 −1.68 0 0 0.00 15.00 2.50 0.00 8.92 1.49 49.04 163.44 28.78 0.20
10 1.68 0 0 60.00 15.00 2.50 35.67 8.92 1.49 13.38 163.44 28.78 0.57
11 0 −1.68 0 30.00 0.00 2.50 17.83 0.00 1.49 40.13 163.44 28.78 0.34
12 0 1.68 0 30.00 30.00 2.50 17.83 17.83 1.49 22.29 163.44 28.78 0.28
13 0 0 −1.68 30.00 15.00 0.00 17.83 8.92 0.00 32.69 163.44 28.78 0.33
Optimisation of Limestone Calcined Clay Cement Based on Response …

14 0 0 1.68 30.00 15.00 5.00 17.83 8.92 2.97 29.72 163.44 28.78 0.36
15 0 0 0 30.00 15.00 2.50 17.83 8.92 1.49 31.21 163.44 28.78 0.37
16 0 0 0 30.00 15.00 2.50 17.83 8.92 1.49 31.21 163.44 28.78 0.36
17 0 0 0 30.00 15.00 2.50 17.83 8.92 1.49 31.21 163.44 28.78 0.39
GPC, general-purpose cement; GYP, gypsum; MK, metakaolin; LS, limestone; SP, superplasticizer
107
108 G. Huang et al.

Table 3 Relevant data on emission and cost for LC3 mortar manufacture
Response Binder materials
MK LS GYP GPC Sand Water SP
CO2 emission, kg CO2 /kg mortar 0.25 0.026 0.14 0.79 0.002 0.007 2.96
Materials cost 0.306 0.001 1.320a 0.550a 0.5 0.003a 16.667a
AUD/kg mortar
aAverage price from local suppliers
GPC, general-purpose cement; GYP, gypsum; MK, metakaolin; LS, limestone; SP, superplasticizer

3 Mathematical Modelling and Statistical Analysis

3.1 ANOVA and Adequacy Checking

The ANOVA for the surface response model revealed the model’s accuracy and
reliability. Specifically, the p-value and F-value at a 95% confidence level were
computed to verify the model and the model’s terms significant degree for each
response. A p-value < 0.05 and a larger F-value indicated the model terms have a
significant influence, while p-values ≥ 0.05 are considered insignificant. In addition,
the lack of fit test measured the degree of fitting between the predicted model and
input variables. If the p-value for lack of fit is > 0.05 (non-significant), it implies
that: (1) the generated model fits the experimental data well, and (2) the independent
variables have considerable effects on the response. When the p-value is < 0.05,
seriously insufficient fitting is recorded, indicating the model was not well fitted to
the input data. The ANOVA results for Y1 to Y3 are shown in Table 4. The response
for the model showed a p-value < 0.05 indicating a good fit; however, the response
for the lack of fit had a p-value > 0.05, indicating an insignificant lack of fit.
The regression coefficients (β0 , βi , βii &βij ) were estimated by ANOVA, and the
regression models for each response as functions of %MK, %LS and %GYP are
summarised in Table 5. The coefficient of determination (R2 ) was obtained to ascer-
tain the accuracy of the predicted function. For all responses, the R2 values were
>95%, which indicated a high degree of correlation to the experimental data.

3.2 Interaction Effects on Response

The three-dimensional response surface plots (see Fig. 2) were shown to evaluate the
influence of interactions of different variables on the corresponding response value.
The plot presented the interaction of two variables considering that the third variable
was fixed on the central point in the CCRD design. The significance of the variables
interaction is denoted by the surface steepness and the variable interaction p-value
[6]. Therefore, the variable interaction with the smallest p-value was considered to
study the influence of two independent variables on the responses. Based on the
Optimisation of Limestone Calcined Clay Cement Based on Response … 109

Table 4 Analysis of variables of the test results

Response 28-day compressive CO2 emission Materials cost
strength
Source F-value p-value F-value p-value F-value p-value
Model 19.47 0.0004 3726.21 < 0.0001 26.62 0.0001
Interactions
X1 34.24 0.0006 21,471.1 < 0.0001 18.74 0.0034
X2 77.75 <0.0001 11,880.6 < 0.0001 183 <0.0001
X3 0.043 0.8414 175.77 < 0.0001 29.81 0.0009
X1 X2 5.8 0.0468 0.5253 0.4921 0.474 0.5131
X1 X3 4.41 0.0738 3.43 0.1063 3.34 0.1102
X2 X3 0.08 0.7861 1.67 0.2367 1.62 0.2431
X1 2 51.12 0.0002 0.2767 0.6151 0.227 0.6483
X2 2 10.59 0.014 1.73 0.2303 1.71 0.2321
X3 2 8.75 0.0212 0.2367 0.6415 0.235 0.6427
Lack of fit 2.14 0.3479 14.73 0.0648 14.6 0.0653

Table 5 Regression models for each response

Response Predicted regression function R2 (%) R2 – predicted (%)
28-day compressive strength Y1 = 37.664 + 0.272x1 − 96.16 73.50
0.0732x2 + 2.181x3 +
0.00896x1 x2 + 0.0468x1 x3 −
0.0126x2 x3 − 0.0112x1 2 −
0.0204x2 2 − 0.667x3 2
CO2 emission Y2 = 0.192 − 0.00127x1 − 99.95 99.52
0.00179x2 − 0.00182x3 +
8.757E-07x1 x2 + 0.000013x1 x3
− 0.00019x2 x3 + 2.677E-07x1 2
− 2.675E-06x2 2 − 0.000036x3 2
Materials cost Y3 = 0.472 − 0.000546x1 − 97.16 77.51
0.00122x2 + 0.000246x3 +
4.74E-06x1 x2 + 0.000075x1 x3 +
0.000105x2 x3 + 1.38E-06x1 2 −
0.000015x2 2 − 0.000202x3 2

ANOVA results shown in Table 4, the selected variable interaction is X1 X2 , X1 X3

and X1 X3 for 28-day compressive strength (Y1 ), CO2 emissions (Y2 ) and materials
cost (Y3 ).
110 G. Huang et al.

Fig. 2 Three-dimensional response surface plots. a 28-day compressive strength, b CO2 emissions,
c materials cost. GYP, gypsum; MK, metakaolin; LS, limestone

3.3 Response Surface Optimization Analysis

A multi-objective optimization scheme was conducted to determine an optimum

formulation of %MK, %LS and %GYP, which satisfied the desired strength spec-
ifications while simultaneously minimizing the CO2 emission and materials cost.
The desirability function approach (range from 0 to 1) was used for this purpose. A
desirability value of 1 represents the ideal case, while 0 indicates the responses are
outside their acceptable limits. The optimal design was verified by the experimental
result and shown in Table 6. An overall desirability of 0.72 indicated a reliable result.
Optimisation of Limestone Calcined Clay Cement Based on Response … 111

Table 6 Optimization results and predicted desirability

Optimum Response Goal Predicted value Experimental Desirability
design value
%MK = 28-day Maximize 34.35 35.63 0.72a
34.16 compressive
strength, MPa
%LS = CO2 emission, Minimize 0.109 0.110
20.6 kg CO2 /kg
sample
%GYP = Materials cost, Minimize 0.436 0.437
1.94 AUD/kg sample
a Predicted overall desirability. GYP, gypsum; MK, metakaolin; LS, limestone

4 Conclusion

This study examined the effects of formulation on strength performance, CO2 emis-
sion, and production cost of LC3 cement. The following conclusions were drawn
based on the experimental results.
• ANOVA showed satisfactory fit and accuracy of the developed quadratic models.
• The optimization results confirmed that the optimal percentages of MK, LS and
GYP were 34.16%, 20.6% and 1.94%, respectively.

References

1. Monteiro PJM, Miller SA, Horvath A (2017) Towards sustainable concrete. Nat Mater 16(7):
698–699
2. Scrivener KL, John VM, Gartner EM (2018) Eco-efficient cements: potential economically
viable solutions for a low-CO2 cement-based materials industry. Cem Concr Res 114:2–26
3. Sharma M, Bishnoi S, Martirena F, Scrivener K (2021) Limestone calcined clay cement and
concrete: a state-of-the-art review. Cem Concr Res 149.
4. Perez-Cortes P, Escalante-Garcia JI (2020) Design and optimisation of alkaline binders of
limestone-metakaolin–a comparison of strength, microstructure and sustainability with portland
cement and geopolymers. J Clean Prod 273
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method. Constr Build Mater 341:127874
112 G. Huang et al.

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
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Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Designing Waterborne Protective
Coatings Through Manipulating
the Nanostructure of Acrylic-Based
Nanocomposites

S. Ji, H. Gui, G. Guan, M. Zhou, Q. Guo, and M. Y. J. Tan

Abstract Waterborne coatings with intended functionalities have been designed by

manipulating acrylic-based nanocomposites with different nanostructures. Taking
advantage of the favorable structure of acrylic copolymers, three waterborne coatings
with various desired properties were created through molecular engineering either by
copolymerizing with other components or through nanocomposite formation. This
approach was demonstrated by synthesizing acrylic-based waterborne coatings with
three different nanostructures, namely homogeneous, worm-like, and spherical-like
nanostructures. The properties of coating samples prepared by this new approach
and by traditional physical blending were compared experimentally, which revealed
that the incorporation of 3-methacryloxypropyltrimethoxysilane (MPS)-modified
nanoparticle TiO2 in an acrylic base enabled the formation of a nanocomposite
with nanoparticles uniformly distributed in the acrylic base. The coating film with
this acrylic-TiO2 nanocomposite showed significantly better UV absorption perfor-
mance than the coating made by physical blending. The copolymerization of acrylic
copolymers with an organic polymer (alkyd) created a worm-like nanostructure
of acrylic–alkyd composite that allowed uniform distribution of the acrylic–alkyd
nanocomposite in a more closely packed dense coating film, leading to enhanced
barrier property and significantly improved corrosion resistance as confirmed by

S. Ji · H. Gui · G. Guan · M. Y. J. Tan (B)

Institute for Frontier Materials, Deakin University, Geelong Waurn Ponds Campus, Waurn Ponds,
VIC, Australia
e-mail: mike.tan@deakin.edu.au
M. Y. J. Tan
School of Engineering, Faculty of Science and Technology, Deakin University, Geelong Waurn
Ponds Campus, Waurn Ponds, VIC, Australia
M. Zhou
Jiangsu Shisong New Material Technology Co., Ltd., Changzhou, Jiangsu, China
School of Environmental & Safety Engineering, Changzhou University, Changzhou, Jiangsu,
China
Q. Guo
School of Engineering, Huzhou University, Huzhou, Zhejiang, China

© The Author(s) 2023 113

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_14
114 S. Ji et al.

electrochemical impedance spectroscopy and salt spray tests. The copolymerization

of acrylic monomers with an inorganic polymer (polydimethylsiloxane [PDMS]) led
to a spherical-like nanostructure of acrylic–PDMS composite film. The formation of
this nanostructure arose from the migration of PDMS segments, and a PDMS-rich
phase formed on the film’s surface, which resulted in a coating film with PDMS func-
tionalities such as low dirt-picking behavior. Overall, these three cases demonstrated
that acrylic copolymer are an excellent base for developing various nanocomposite
waterborne coatings with different functionalities through copolymerization and that
the nanocomposites with different nanostructures have a significant influence on the
coatings’ performance.

Keywords Acrylic copolymer · Copolymerization · Multifunctional coating ·

Nanostructure · Waterborne coatings

1 Introduction

Water-based coatings are eco-friendly because they reduce the emission of environ-
mentally unfriendly volatile organic compounds. Waterborne coatings are usually
made by physically blending and emulsifying the binder, pigments, and additives
with water. Copolymers of acrylic, vinyl, and styrene compounds are typically used
as the binder in this type of coating. Of all the types of polymers, acrylic copolymer is
one of most commonly used resins in waterborne coatings. In general, acrylic resins
are synthetic copolymers of acrylic and methacrylic acids or their corresponding
esters created by free-radical polymerization [1]. Acrylic resin is a low-cost mate-
rial compared with other resins such as epoxy resins and polyurethane, and acrylic
copolymers can be stable in water to form aqueous dispersions for application in eco-
friendly coatings [2, 3]. They show excellent performance in film formation, high
gloss, good adhesion, fast drying, outdoor durability, high transparency, and so forth
[4–11]. Acrylic resins have good compatibility with other components and can be
modified by a variety of components, such as polymers, silica nanoparticles, mucin
gel, fiber, etc., to form coatings with improved performance [12–15]. In general,
acrylic resin is a popular base polymer for developing coatings with desired new
properties and for applications in adhesives, construction, automotive and additives,
etc. [16–21]. Unfortunately, acrylic-based waterborne coatings have some weak-
nesses such as low water and corrosion resistance, poor thermal stability and weak
chemical resistance, and waterborne acrylic-based coatings are considered inappro-
priate for application in corrosive environments. If these weaknesses of waterborne
acrylic coatings can be overcome, their applicability could significantly broaden.
The monomers used to synthesize acrylic copolymers are all with vinyl groups,
which makes the structure of acrylic copolymers easy to manipulate. In this regard,
acrylic copolymers can easily react with other components and form covalent
bonds between different components. The acrylic-based copolymers can be mixed
at a molecular scale and can exhibit new functionalities. For instance, anti-icing
Designing Waterborne Protective Coatings Through Manipulating … 115

performance was gained by the acrylic–silicone copolymer used for wind turbine
blades [22]. The introduction of indole derivative groups or tertiary amines into
acrylic resins can produce copolymers with anti-fouling performance [23, 24]. A
biocompatible copolymer was copolymerized by acrylic copolymer and polyhe-
draloligosilsesquioxane [25]. These acrylic-based copolymers exhibit biocompat-
ible stability and could be developed as a denture resin. An acrylic–poly(dimethyl
siloxane) copolymer with improved gas permselectivity was synthesized by an atom
transfer radical polymerization technique [26].
Conventionally, acrylic-based copolymers are prepared with a variety of mate-
rials, including organic polymers (polyurethane, polystyrene and epoxy), organic
silicone, inorganic nanoparticles (clay, silica), etc. In this study we intended to extend
this to the designing of waterborne coatings with intended functionalities through
manipulation of acrylic-based copolymers with different nanostructures. The objec-
tive was to show that the acrylic copolymer can be an excellent base polymer for
developing waterborne coatings by manipulating its nanostructure through the intro-
duction of different components into the acrylic base. Experiments were carried out to
demonstrate this approach by synthesizing several acrylic-based waterborne coatings
with different nanostructures, including homogeneous, worm-like, and spherical-like
nanostructures, by introducing three different components.

2 Methods

2.1 Materials

Acrylic acid (AA), 1,1’-azobis(cyclohexanecarbonitrile) (ACHN), 2-

dimethylaminoethanol (DMEA), 1-methoxy-2-propanol, 2-mercaptoethanol, 3-
methacryloxypropyltrimethoxysilane (MPS), titanium (IV) oxide (TiO2 ), methacry-
loyl chloride, bis(3-aminopropyl) terminated PDMS and 1-methoxy-2-propanol
obtained from Sigma-Aldrich were used directly without further purification.
Styrene (St), n-butyl acrylate (BA), methyl methacrylate (MMA), 2-hydroxypropyl
acrylate (HPA) were supplied by Sigma-Aldrich and were all purified by Al2 O3
(Sigma-Aldrich) chromatographic column. Alkyd was synthesized as described
in our previous study [27]. Curing agent amino resin Resimene 717, a type
of melamine resin with alkoxy groups, was supplied by Jiangsu Shisong New
Materials Technology Co., Ltd.
116 S. Ji et al.

2.2 Preparation of Waterborne Acrylic Copolymer,

Acrylic-Based Coatings and Coating Films

TiO2 was modified by MPS to introduce reactivity [28]. Bis(3-aminopropyl) termi-

nated PDMS and methacryloyl chloride were used to prepare vinyl-terminated
PDMS. The copolymerization process of acrylic and acrylic-based coatings was as
follows. The MPS-modified TiO2 or alkyd or vinyl-terminated PDMS was dispersed
in 1-methoxy-2-propanol and added to a 4-necked flask. A mixture of acrylic
monomers (BA (50 g), HPA (15 g), AA (15 g), MMA (10 g) and St (10 g)), initiator
and solvent was then dropped into the flask and the reaction temperature was 88 °C.
The content of MPS-modified TiO2 or alkyd or vinyl-terminated PDMS was 5 wt%
of the acrylic monomers. The reaction was maintained for 3 h. After copolymer-
ization, the solvent was removed by nitrogen gas. DMEA was used to neutralize
part of the carboxyl groups at 50–60 °C, and water was added to prepare copolymer
aqueous dispersions under mechanical stirring to obtain completely homogeneous
dispersions. The copolymers synthesized were named acrylic, acrylic-TiO2 coating,
acrylic–alkyd coating and acrylic–PDMS coating according to the additional compo-
nent added to the acrylic system. The samples that were a physical blending of acrylic
copolymer aqueous dispersion with MPS-modified TiO2 , alkyd and vinyl-terminated
PDMS separately were also prepared for comparison, and the loading content of other
components was 5%wt of acrylic copolymer. The samples were named PB-acrylic/
TiO2 coating, PB-acrylic/alkyd coating and PB-acrylic/PDMS coating. The coating
films were prepared from copolymer aqueous dispersions, which were mixed with
the curing agent. After mixing, the mixture was poured onto the substrate and a film
adaptor was used to get a wet film of 200 µm. The wet films were dried at room
temperature for 1 h, and then cured at 140 °C for 30 min to get dry films of 20 ±
2 µm.

2.3 Characterization

The absorption performance of the coating films was measured using an

ultraviolet−visible spectrophotometer (Cary Series UV–Vis spectrophotometer,
Agilent, USA) in the range of 200–800 nm. The ultraviolet protection factor (UPF)
value of the coating films was measured by a UPF and UV penetration/projection
measurement system (Model: YG902) from 280 to 400 nm. The mean UPF was
calculated automatically by the test system from eight sets of data obtained from
different test areas. Transmission electron microscopy (TEM) was used to study
the morphologies with a JEOL-2100 microscope under 120 kV. To obtain clear
morphologies of micelles, the acrylic and acrylic–alkyd coating samples were nega-
tively stained with a uranyl acetate substitute (UAR-EMS stain) for 30 min at room
temperature. Atomic force microscopy (AFM) was used to evaluate the surface
morphology and microphases of the coating films. The AFM measurements of the
Designing Waterborne Protective Coatings Through Manipulating … 117

films were performed with a Cypher AFM microscope (Asylum Research). Elec-
trochemical impedance spectroscopy (EIS) tests were conducted with a Bio-Logic
electrochemical workstation at ambient temperature (25 ± 2 °C). A three-electrode
cell arrangement with 3.5%wt (w/v) NaCl solution was used to conduct the tests.
Carbon steel substrates coated with the films were set as the working electrode with
a circular tested area of ~ 1 cm2 . An Ag/AgCl (Sat. KCl) electrode was used as
the reference electrode and a Pt-coated Ti mesh was used as the counter electrode.
The amplitude of the sinusoidal voltage was 10 mV, and the frequency range was
100 kHz to 10 MHz. The EIS data were acquired when the samples were immersed
in salt solution after 2 days. Immersion test was performed in 3.5 wt% (w/v) NaCl
solution at room temperature. The edges and back sides of the samples used for the
immersion test were all covered with epoxy resin. The surface morphologies of the
coating films were detected by scanning electron microscopy (SEM) with a Zeiss
Supra 55 VP under 5 kV. The samples were coated with 5-nm Au film for good
conductivity. The contact angle of the coating films was evaluated by Tensiometer
KSV CAM 101 (KSV Instruments Ltd, Finland) at room temperature and water was
used as the test liquid. The dirt-picking performance was measured by placing black
carbon particles on the coating film and using water to clean the films.

3 Results and Discussion

3.1 Enhanced UV Protection Through the Formation

of an Acrylic-Based Nanocomposite

An example of creating a functional waterborne coating with desirable nanostruc-

ture is the incorporation of MPS-modified nanoparticle TiO2 in an acrylic base to
enable the formation of a homogeneous nanostructure with nanoparticles uniformly
distributed in the acrylic base. UV protection is an important property for waterborne
coatings that are exposed to intense UV environmental conditions. It is well known
that incorporating UV absorbers into the polymer matrix is a method of improving
the UV protection of polymers. Nanoparticle TiO2 is a commonly used UV absorber
for improving the protection performance of acrylics. In the present study, the UV
protection performance of the acrylic base coating and the coating with nanoparticle
TiO2 was evaluated by UV absorbance and UPF value tests, as shown in Fig. 1a,
b. The absorbance of the PB-acrylic/TiO2 coating film slightly increased compared
with acrylic film, which we believed to be due physical blending being unable to
achieve a coating with uniformly distributed nanoparticle TiO2 . The distribution of
nanoparticles and compatibility between nanoparticles and polymer matrix need to
be improved to obtain well-dispersed mixer for high efficiency of UV protection [29,
30]. This was achieved by the acrylic–TiO2 coating film that incorporated nanopar-
ticle TiO2 in the acrylic base by copolymerization, which enabled the formation of a
118 S. Ji et al.

nanocomposite with a homogeneous nanostructure of uniformly distributed nanopar-

ticles in the acrylic base. As shown in Fig. 1a, the coating film with the acrylic–
TiO2 nanocomposite showed significantly better UV absorption performance than
the coating film made by physical blending, especially in the wavelength between 315
and 400 nm (UV-A). In addition, the UPF value of the acrylic–TiO2 coating film was
more than threefold higher than that of the PB-acrylic/TiO2 coating film (Fig. 1b).
The remarkable improvements in UV absorbance and UPF value of the acrylic–TiO2
coating can be explained by the uniformity of the coating film. The appearance of
the PB-acrylic/TiO2 and acrylic–TiO2 coating films is shown in Fig. 1c. It is obvious
that the acrylic–TiO2 coating film has a homogeneous distribution of nanoparticles.
The difference in UV absorbance can be explained by the acrylic–TiO2 coating
prepared through copolymerization having improved compatibility between the
acrylic base and the nanoparticle TiO2 . In this case, the copolymerization of acrylic
and MPS-modified nanoparticle TiO2 achieved the uniform distribution. From the
TEM image in Fig. 2b, it is obvious that the TiO2 nanoparticles are wrapped by
the acrylic copolymer. Moreover, from the AFM images in Fig. 3, the acrylic–TiO2
coating film has a relatively homogeneous nanostructure, therefore the filling of TiO2
nanoparticles by copolymerization leads to a uniform coating film compared with

Fig. 1 a UV absorption spectra of acrylic base coating film and acrylic-based coating films by UV–
vis, b UPF value of coating films in the wavelength of 280–400 nm, c images of the PB-acrylic/
TiO2 and acrylic-TiO2 coating films
Designing Waterborne Protective Coatings Through Manipulating … 119

Fig. 2 Transmission electron microscopy images of copolymer water dispersions, a acrylic

copolymer, b acrylic-TiO2 copolymer c acrylic–alkyd copolymer, and d acrylic–PDMS copolymer

physical blending. In general, the enhanced UV absorbance and UPF value of the
acrylic–TiO2 coating film suggest that the uniform distribution of nanoparticle TiO2
in the acrylic base plays an important role in improving UV protection.

3.2 Enhanced Corrosion Resistance of Acrylic–Alkyd

Coating with Worm-Like Nanostructure

Acrylic waterborne coatings often have weak corrosion resistance in humid envi-
ronments [31, 32]. A possibility method of enhancing their corrosion resistance is
to incorporate other components with high hydrophobicity such as alkyd in order to
produce anti-corrosion waterborne coatings [33]. Unfortunately, waterborne coatings
made by physical blending of acrylic and alkyd polymer failed to show increased
anti-corrosion performance. As shown in Fig. 4, the corrosion resistance of PB-
acrylic/alkyd coating film, as indicated by the impedance values from the Nyquist
plots and impedance spectrum of the coating, actually reduced compared with the
acrylic coating. This result suggested that physical blending of acrylic with alkyd
cannot improve the corrosion resistance of the acrylic base, possibly because phys-
ical blending of acrylic and alkyd cannot obtain a uniform coating film with a dense
120 S. Ji et al.

Fig. 3 Atomic force microscopy topographic and phase images of acrylic copolymer film and
acrylic-based coating films

film surface due to incompatibility. In contrast, as shown in Fig. 4, the anti-corrosion

performance of the acrylic–alkyd coating film was significantly enhanced, which
we believe was due to the uniform distribution of alkyd in the acrylic base, better
compatibility and the formation of a dense coating film. For the acrylic–alkyd coating
film, the aqueous dispersion formed nanoparticles of ≈100 nm as shown in Fig. 2c.
In addition, the acrylic–alkyd coating film formed a nanocomposite with a uniform
and worm-like nanostructure shown in Fig. 3. The worm-like nanostructure was
more closely packed, which created a better barrier that reduced the penetration of
corrosive salt solution into the coating film, thereby enhancing the anti-corrosion
performance of the acrylic base. The dense film surface of the acrylic–alkyd coating
film was confirmed in the SEM image shown in Fig. 4f as compared with the loose
surface of the acrylic film shown in Fig. 4e. Moreover, the immersion test confirmed
the EIS result.

3.3 Multifunctional Acrylic–PDMS Coating

with Spherical-Like Nanostructure

Acrylic waterborne coatings are also commonly used to protect surfaces from dirt and
graffiti. PDMS has low surface energy and high hydrophobicity, which is often used
to creating easy-clean coatings. As another example of creating a desirable water-
borne coating through manipulation of acrylic-based nanocomposites with different
Designing Waterborne Protective Coatings Through Manipulating … 121

Fig. 4 Impedance spectra of the acrylic copolymer film and acrylic-based coated carbon steel in
3.5%wt NaCl aqueous solution: a Nyquist plots, b enlarged part of the Nyquist plots, c Bode plots,
d immersion tests of coating films in 3.5 wt% NaCl solution at room temperature; scanning electron
microscopy images of e acrylic film, and f acrylic–alkyd coating film

nanostructures, the PDMS component was incorporated to prepare an acrylic-based

copolymer with a high dirt resistance. The surface performance of the coating films
was evaluated, as shown in Fig. 5a. The contact angle of the acrylic–PDMS coating
film increased compared with acrylic film and the PB-acrylic/PDMS coating film
formed by blending the hydrophobic PDMS component. In addition, the dirt-resisting
performance of the coating films was tested and shown in Fig. 5b. From the images
of the PB-acrylic/PDMS coating film, it is obvious that the black carbon particles
122 S. Ji et al.

Fig. 5 a Water contact angle of acrylic and acrylic-based coating films, and b images of dirt-picking
property of the coating films

were not washed off completely, whereas the acrylic–PDMS coating film was totally
cleaned by water washing.
The dirt-resistance performance of the PB-acrylic/PDMS coating film was unsat-
isfactory because the PDMS component is hydrophobic and cannot be uniformly
distributed in the acrylic base by physical blending due to the poor compatibility
between acrylic and PDMS. In contrast, copolymerization improved the compati-
bility between the acrylic and PDMS components and contributed to the formation
of a uniform and spherical-like nanostructure, which can explain the excellent dirt-
resistant property of the acrylic–PDMS copolymer. From Fig. 2d, it is clear that the
acrylic–PDMS copolymer formed a nanocomposite with particles at the nanoscale.
Figure 3 shows spherical-like nanostructure of the acrylic–PDMS coating. In this
nanostructure, the PDMS component tended to migrated to the top of the surface
due to its low surface energy and hydrophobicity. In this case, the acrylic–PDMS
film surface is a PDMS-rich phase that can show similar properties to the PDMS
component. These properties contributed to the improvement in the dirt-picking
performance of the acrylic–PDMS coating film.

4 Conclusion

We carried out experiments to demonstrate the design of multifunctional water-

borne coatings through manipulation of acrylic-based copolymers with different
nanostructures. Three components were chosen to study–an inorganic nanoparticle
(TiO2 ), an organic short polymer (alkyd) and an inorganic polymer (PDMS)–to
manipulate the acrylic nanostructure in order to create waterborne coatings with
desirable functionalities and improve the compatibility between different compo-
nents. The copolymerization of MPS-modified TiO2 and acrylic monomers achieved
a uniform acrylic–TiO2 coating film. The uniform distribution of nanoparticles led
to better UV protection property compared with the physical blended sample. Due
Designing Waterborne Protective Coatings Through Manipulating … 123

to thermodynamic incompatibility of acrylic and alkyd, the acrylic–alkyd coating

film formed a worm-like nanostructure, which contributed to its dense film surface
that had better barrier performance and enhanced the corrosion resistance. The
acrylic–PDMS coating film formed a spherical-like nanostructure and the film surface
was a PDMS-rich phase, which showed similar performance to the PDMS compo-
nent compared with PB-acrylic/PDMS. Therefore, the properties of the acrylic–
PDMS coating changed significantly: the contact angle increased to 103° and the
coating film showed enhanced resistance to dirt-picking. Generally, by copolymer-
izing acrylic with other components, it is easy to manipulate the nanostructure of
acrylic-based copolymers and improve compatibility between different components.
Moreover, the nanostructure of acrylic-based copolymers plays an important role
in the coating’s performance. Acrylic-based copolymers with an organized nanos-
tructure can exhibit desirable functions. This approach to designing acrylic-based
copolymers with an organized nanostructure could open opportunities for developing
waterborne acrylic-based copolymers with multifunctionality.

5 Data Availability Statement

The raw/processed data required to reproduce these findings cannot be shared at this
time as the data also forms part of an ongoing study.

Acknowledgements Financial support from Australian Research Council funding for ARC
Research Hub for Nanoscience-Based Construction ARC ITRH-JSNMT-2016-2021 is gratefully
acknowledged.

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Analysis of Categories That Delay Global
Construction Projects

M. Abonassrya, M. Alam, and A. Saifullah

Abstract Delay in construction projects is a significant issue and concern for most
construction companies. Many studies have addressed this issue by identifying the
top-ranked causes, which vary according to project type, location, and the research
method used. The combined factors of delay/time overrun need further analysis
to understand the top-ranked factors considering the project context. We identified
360 delay/time overrun factors of construction projects from articles published in
the past 10 years in top-quality journals ranked as per scientific journal ranking
(SJR). The factors were then coded and classified into categories based on their
impact and the description in NVIVO software. Finally, the categories were analyzed
and ranked by the relative importance index using SPSS software to identify the
critical ones in global construction projects. In addition, the developed and developing
countries affected by these delay factors were determined. The results revealed that
the top five important factors are located under the following categories: orders and
requirements; experience and productivity; financial problems; planning; and lastly
both external and management categories. These categories were the highest ranking
among the five top factors found in the reviewed studies and affect both developed
and developing countries.

Keywords Construction · Delays · Economic impact · Overrun

M. Abonassrya (B) · M. Alam

Department of Civil and Construction Engineering, Swinburne University of Technology, VIC,
Australia
e-mail: mabonassrya@swin.edu.au
A. Saifullah
Department of Mechanical Engineering and Product Design Engineering, Swinburne University
of Technology, VIC, Australia

© The Author(s) 2023 127

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_15
128 M. Abonassrya et al.

1 Introduction

Delay/time overrun is a common issue in construction projects worldwide, and the

success of construction projects can be considered a measure of the political perfor-
mance of countries because it is related directly to their economy [1]. Faridi and
El-Sayegh state that project delays negatively affect the quality and safety of project
work and reflect negatively on the services provided to the community [2]. There-
fore, in the present study we performed a literature review focused on identifying
and ranking the delay factors, although classification of these factors differs among
studies. The relative importance index (RII) analysis was commonly used and many
types of construction projects were researched and discussed. The delay factors were
classified under different categories, but the most prominent were the contractor and
client categories, which relate to financial problems, management, materials and
equipment, experience, planning, and methods of construction. In addition, the delay
factors in developing countries were greater than in developed countries. The delay
factors were many and varied according to the project type, location, and the research
method used in the study. Our main focus was to identify the critical categories of
delay factors found by previous studies over the past 10 years published in jour-
nals ranked under the scientific journal rankings (SJR). The factors identified were
classified into categories that were analyzed by Cronbach’s alpha for reliability veri-
fying and RII using SPSS software for ranking. The top five categories were: orders
and requirements; experiences and productivity; financial problems; planning; and
external and management.

2 Literature Review

The research methods used in the studies differed, but usually followed two trends:
identifying and analyzing the factors [3]. However, some studies also discussed the
effects of delay factors [4]. The delay factors identified were generally classified
into categories applicable for each study. For example, Bajjou and Chafi identified
the delay factors in African countries and classified them into eight categories [5].
Similarly, Wuala and Rarasati analyzed and classified the factors in Southeast Asia
into five categories [6]. In addition, studies discussed the delay factors in different
types of projects [7]. For example, tunnels [8], residential projects [9–11], roads [12–
14], railways [15, 16], sport facilities [17], oil projects [18, 19], as well as transport,
power, building, and water irrigation projects [20]. These various delay factors were
then also variably categorized. For example, financial problems were identified as the
contractor category [21–24], but also as the client category [7, 25–28]. Management
factors were placed in the contractor category [29, 30], but also related to shortages
in equipment and material factors [6, 31]. Some studies stated that the contractor
category related to experience, planning, and construction method factors [32–34],
whereas in others the client category was related to decision-making factors [35, 36],
Analysis of Categories That Delay Global Construction Projects 129

or to changes factors [19, 37]. In addition, there were many other categories related
to other delay factors, such as the external category related to weather, policy, and
security factors [8, 10, 11, 16, 38, 39]. The delay in construction projects affects
both developed and developing countries. Rivera et al. studied the delay factors in
25 developing countries and confirmed that 50% of them have similar delay factors
[32]. On the other hand, developed countries are also affected by delay factors but
less so, according to analysis of three countries: Portugal, the UK, and the USA
[40]. In addition, other studies identified global delay factors for both developed and
developing countries [17, 23, 41–43]. We reviewed all the studies that addressed
and identified delay factors to analyze the top-ranked factors, and then identified the
developed and developing countries affected by these factors.

3 Research Method

Figure 1 details the stages in our research. We began with the research scope to
select relevant studies, which were then refined. The factors (delay/time overrun)
were extracted from the selected studies, classified into categories and ranked based
on their importance from the researchers’ point of view. Next we analyzed the factor
categories by Cronbach’s alpha and the RII. Lastly, the resulting categories were also
classified to the relevant developed/developing country.

3.1 Research Scope and Select Studies

Using the Scopus database, we collected and sorted studies that addressed the relevant
factors (i.e., delay/time overrun) in construction projects. Screening was limited to
the title of the study. Next we selected studies published in the 10 years of 2012–
2021 with the word “delay” in the title and 410 studies were identified. All steps
were repeated for studies with “time overrun” in the title and 41 were identified.
The 451 studies were then screened for publication in a high-quality journal or
conference ranked by SJR, resulting in 277 studies, which were further reviewed and
refined to choose those that clearly identified and ranked the factors of delay/time
overrun in different countries. Finally, 71 studies in total were selected: 62 for delay
and 9 for time overrun. Microsoft Excel was used in this stage of the research.

3.2 Factors Identified and Classification into Categories

The studies reviewed in this research depended on delay/time overrun factors iden-
tified in previous studies that other researchers had performed through interviews,
questionnaires, or case studies. The research area included many developing and
130 M. Abonassrya et al.

Fig. 1 Research flowchart

developed countries. These studies ranked the top essential factors of delay/time
overrun, so the top five factors selected from previous studies were included in this
research. The delay factors were not all similar in terms of their importance, as they
were based on the project type, location, and the method of identifying the factors
used in the study. So, some factors were repeated or had a similar description in some
studies, and other factors were utterly different in their description in other studies.
The factors extracted from all the studies reviewed had been classified into cate-
gories based on the type and description of each factor and thus to any category it
Analysis of Categories That Delay Global Construction Projects 131

most likely belonged. All categories and their factors were coded using NVIVO soft-
ware, and factors were ranked inside their categories from 1 to 5 based on the rank
that each factor had in the original study. Some categories were combined into one
category to reduce the number of categories, such as design and work error; experi-
ence and productivity; material and equipment; orders and requirements; and tender
and contract category. In addition, some studies contained some factors belonging
to the same category.

3.3 Analysis of Factor Categories

We applied two types of analysis: Cronbach’s alpha analysis to verify the reliability
of data collected regarding delay/time overrun factors and RII analysis to rank the
critical categories of the top five factors using SPSS software. The various important
factors were identified and ranked based on the project type, location, and research
method of the previous studies, which depended on literature reviews to identify the
factors, then on questionnaires, surveys or interviews to assess the importance of
the factors, and finally, on ranking these factors by analysis. In the questionnaires
or interviews, the answers and importance degree of the factors identified in the
previous studies were collected from respondents and the answers are ranked to
identify the most important using RII analysis. Therefore, we assumed our study was
similar to a questionnaire survey of the previous studies that addressed and ranked
the top five delay factors, whereby categories classified were questions, studies were
respondents, and importance degree of the factors comprised the answers.
We adopted a Likert scale with 5 points to identify the top five essential levels
for the top five factors identified in previous studies. The five levels of importance
were very high, high, moderate, low, and very low, ranked from 5 to 1 degree of
importance. Factors ranked as the first, which was the highest level in the previous
studies, took a very high ‘5’ degree of importance, and factors ranked as second, third,
fourth, and fifth took a high ‘4’, moderate ‘3’, low ‘2’, and a very low ‘1’ degree of
importance respectively. It is good to analyze all factors classified into categories,
but the total of factors (i.e., answers) in all categories (i.e., questions) was not equal.
Therefore, it was necessary to first reduce the factors for each category to be equal to
the category’s lowest value of the factors by ignoring the levels of the least important
factors. However, the total factors selected after reducing the factors of categories
became less than half of the original factors, even with the categories having the
lower second, third and fourth values of the total factors. In addition, the total of
factors was more than half of the original factors when selecting the following lower
fifth and sixth values. However, the sixth lower value of factors had fewer factors
and categories than the fifth lower value and there was no point in proceeding with
other lower values because the total factors became less, as well as the categories,
and it beneficial to analyze the greatest number of factors. So, the best scenario was
using the category having the fifth lower value of factors because the total factors
selected was more than half the original total. In this process, four categories that
132 M. Abonassrya et al.

had factors with the lowest importance levels were excluded from the analysis. The
factors in the remaining categories were analyzed by the two methods stated above
using SPSS software.
In addition, the factors extracted from the previous studies included many different
countries. Therefore, as the last step in this research, the studies of factor categories
resulting from the analysis were classified as developed or developing states. The
developing states were identified as per the list of developing countries updated in
2022 on the website of the Australian Government Department of Foreign Affairs
and Trade [44]. Finally, the percentage of developed and developing states for each
category was calculated and tabulated.

4 Results and Discussion

A total of 360 factors of delay were identified in the 71 studies. The classification of
these factors resulted in 14 categories as per the description of each factor and the
category to which it belonged. In addition, we found that two or three factors in most
studies were mainly classified under the same category. Also, all studies identified
5 factors, except for 5 studies that identified 6 factors due to two factors having the
same rank. Each factor’s degree of importance in 14 categories is detailed in Table 1.
The total factors ranged from 8 in C1, to 49 in C4. There were 72 factors ranked as
the first degree of importance, and 71 factors each for the second and third degrees
of importance. Lastly, there were 73 factors each in the fourth and fifth importance
degrees.
After refining the factors stated in the research method, six scenarios were
proposed and selected (Fig. 2). Scenario 5 achieved the highest number of factors,
and by getting the factors with the highest degree of importance, the factors for each
category was reduced to 19. At the same time, the total of factors was 190, which was
more than half of the 360 original factors. Also, the factors in this scenario belonged
to 10 categories, because four categories with less than 19 factors were excluded.
Table 2 and Fig. 2 detail scenario 5.
Cronbach’s alpha analysis was the first analysis applied and the result was 0.974,
which indicated an excellent level of reliability (>0.8) and confirmed that the data
collected was very acceptable and reliable and can be used for further analysis. Table
3 shows the internal consistency level of Cronbach’s alpha analysis.
Table 4 shows the results and ranks of the 10 categories analyzed after excluding
the categories with the least importance levels of factors (i.e., C1, C2, C12 & C14).
The results of the analysis illustrated the categories of the top critical delay factors
ranked in previous studies for the 10 year period. The ranking of categories confirmed
that the “orders and requirements” (C9) category had the highest level of importance
(RII = 0.905), the “experience and productivity” (C4) category was next (RII =
0.842), the third category was “financial problems” (C6) (RI I = 0.832), followed
by the “planning” category (C11; RII = 0.789). Lastly, both the “external” (C5) and
“management” (C7) categories came fifth (RII = 0.779). Also, it was noted that the
Analysis of Categories That Delay Global Construction Projects 133

Table 1 Frequency of degree of importance of factors

No Category Code Degree of importance Total
1st 2nd 3rd 4th 5th
1 Communication C1 0 4 2 1 1 8
2 Conflict C2 2 1 5 2 0 10
3 Design and work error C3 5 7 2 1 4 19
4 Experience and productivity C4 7 9 12 14 7 49
5 External C5 8 4 4 5 5 26
6 Financial problems C6 6 10 6 7 8 37
7 Management C7 6 5 10 8 3 32
8 Material and equipment C8 5 2 9 6 12 34
9 Orders and requirements C9 11 7 1 8 10 37
10 Payment delay C10 6 4 4 6 3 23
11 Planning C11 4 10 9 9 7 39
12 Project C12 1 2 3 0 6 12
13 Response delay C13 8 3 4 2 4 21
14 Tender and contract C14 3 3 0 4 3 13
Total 72 71 71 73 73 360

200 190 189

180
Number of factors / categories

160 144 143

140 130
120 112

100
80
60
40
19 21
14 10 13 12 12 13 11 10 9
20 8
0
Scenario 1 Scenario 2 Scenario 3 Scenario 4 Scenario 5 Scenario 6

No. of factors for each category No. of categories Total of factors for all categories

Fig. 2 Proposed scenarios for analyzing the categories of factors

importance levels for all 10 categories were between very high and high. The RII
of the first three categories (C9, C4 & C6) was > 0.8, so they all had a very high
level of importance. In contrast, the fourth (C11), and fifth categories (C5 & C7) had
a high importance level with all other remaining categories. Table 5 illustrates the
importance levels in the RII analysis.
134 M. Abonassrya et al.

Table 2 Frequency of importance degree of 19 factors for each category in scenario 5

No Category Code Degree of importance Total
1st 2nd 3rd 4th 5th
1 Design and work error C3 5 7 2 1 4 19
2 Experience and productivity C4 7 9 3 0 0 19
3 External C5 8 4 4 3 0 19
4 Financial problems C6 6 10 3 0 0 19
5 Management C7 6 5 8 0 0 19
6 Material and equipment C8 5 2 9 3 0 19
7 Orders and requirements C9 11 7 1 0 0 19
8 Payment delay C10 6 4 4 5 0 19
9 Planning C11 4 10 5 0 0 19
10 Response delay C13 8 3 4 2 2 19
Total 66 61 43 14 6 190

Table 3 Internal consistency level of Cronbach’s alpha [7]

Scale Reliability level Cronbach’s alpha
1 Poor α < 0.5
2 Satisfactory 0.5 α < 0.7
3 Good 0.7 α < 0.8
4 Excellent 0.8 α

Table 4 Ranking of categories of factors by RII analysis

No Category Code N Mean RII Rank Importance level
1 Design and work C3 19 3.4211 0.684 9 High
error
2 Experience and C4 19 4.2105 0.842 2 Very high
productivity
3 External C5 19 3.8947 0.779 5 High
4 Financial C6 19 4.1579 0.832 3 Very high
problems
5 Management C7 19 3.8947 0.779 5 High
6 Material and C8 19 3.4737 0.695 8 High
equipment
7 Orders and C9 19 4.5263 0.905 1 Very high
requirements
8 Payment delay C10 19 3.5789 0.716 7 High
9 Planning C11 19 3.9474 0.789 4 High
10 Response delay C13 19 3.6842 0.737 6 High
Analysis of Categories That Delay Global Construction Projects 135

Table 5 Importance levels in the RII analysis [7]

Scale Importance level RII
1 Very low 0.0 RII 0.2
2 Low 0.2 < RII 0.4
3 Moderate 0.4 < RII 0.6
4 High 0.6 < RII 0.8
5 Very high 0.8 < RII 1.0

In addition, the effect of the factors varied according to the different states. Thus,
three groups of countries for each category (developed, developing, and global)
were classified. In this classification, we considered and calculated all 360 factors,
not just those refined and analyzed by RII. Also, some studies discussed global
factors and were classified under the global group because they had not specified
developed or developing countries. The factors of developed countries were stated
in 95 studies, 239 reported factors in developing countries, and 26 studies discussed
global factors. Table 6 shows the total factors divided as into those affecting developed
and developing countries, as well as global factors.

Table 6 Categories of factors in developed and developing countries

No Category Code Developed Developing Global Total
1 Communication C1 3 4 1 8
2 Conflict C2 2 7 1 10
3 Design and work C3 10 8 1 19
error
4 Experience and C4 14 29 6 49
productivity
5 External C5 5 20 1 26
6 Financial problems C6 7 28 2 37
7 Management C7 10 20 2 32
8 Material and C8 6 23 5 34
equipment
9 Orders and C9 11 25 1 37
requirements
10 Payment delay C10 4 19 0 23
11 Planning C11 9 26 4 39
12 Project C12 1 11 0 12
13 Response delay C13 9 11 1 21
14 Tender and contract C14 4 8 1 13
Total 95 239 26 360
136 M. Abonassrya et al.

We also ranked the categories in each group of countries and overall, as shown in
Table 7. We excluded global factors in order to identify just those within developed
and developing countries groups. The category of “experience and productivity” C4
was the highest rank in terms of the number of factors stated in both developed and
developing countries groups and as well as overall rank. Followed by C9 second,
C3 & C7 third, C11 & C13 fourth, and C6 as the fifth category in the developed
countries group ranking. In the developing countries group, C6, C11, C9, and C8 were
ranked one by one. In the overall ranking, the category of “orders and requirements”
(C9) was second; “financial problems” (C6) and “planning” (C11) were third, and
“management” (C7) and “material and equipment” (C8) were fourth and fifth.
In total, 334 factors were classified within these groups: 95 for developed countries
and 239 for developing countries. Therefore, the percentage of developed countries
affected by delay factors was 28%, which was less than the percentage of developing
countries affected, which was 72%. Table 7 details the ranks of categories affecting
developed and developing countries and the overall rank.
Finally, the results clearly showed that the ranks of categories in terms of the
top critical factors in construction projects were very close to the overall ranking

Table 7 Ranking of categories in developed and developing countries

No Category Code Developed Developing Overall
Factor Rank Factor Rank Factor Rank
1 Communication C1 3 9 4 11 7 12
2 Conflict C2 2 10 7 10 9 11
3 Design and work C3 10 3 8 9 18 9
error
4 Experience and C4 14 1 29 1 43 1
productivity
5 External C5 5 7 20 6 25 6
6 Financial C6 7 5 28 2 35 3
problems
7 Management C7 10 3 20 6 30 4
8 Material and C8 6 6 23 5 29 5
equipment
9 Orders and C9 11 2 25 4 36 2
requirements
10 Payment delay C10 4 8 19 7 23 7
11 Planning C11 9 4 26 3 35 3
12 Project C12 1 11 11 8 12 10
13 Response delay C13 9 4 11 8 20 8
14 Tender and C14 4 8 8 9 12 10
contract
Total 95 239 334
Analysis of Categories That Delay Global Construction Projects 137

of developed and developing countries affected by these factors, where the critical
categories were the same in both the important factors ranking, and the overall ranking
of countries, although these categories were not same as a ranking from 1 to 5.
However, some categories shared the same rank; for example, category C5 was
shared as ranking five in categories of essential factors with C7, whereas C6 and C11
shared in the same rank of three in the overall ranking of countries.

5 Conclusion

Delay remains a contentious issue in construction projects, and it is due to many

different factors. We identified 14 critical categories classified from the top five delay
factors identified in previous studies. The critical categories with the highest ranking
were orders and requirements; experience and productivity; financial problems; plan-
ning; and both external and management factors. All categories of essential delay
factors identified in this research have affected both developed and developing coun-
tries, although the effect was 28% for developed countries and 72% for developing
countries.

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Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
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The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Chloride Penetration in Low-Carbon
Concrete with High Volume of SCM:
A Review Study

C. Xue and V. Sirivivatnanon

Abstract Low-carbon concrete (LCC) uses supplementary cementitious material

(SCM) to partially replace cement as a method for reducing its carbon footprint.
Previous laboratory and field studies had provided substantial support and experi-
ence for using LCC in marine structures, which are the most susceptible to chloride-
induced corrosion. Some short-term test methods have provided reliable assessment
of the ability of LCC to resist chloride penetration, but the long-term chloride pene-
tration depends on a great many factors and thus could differ from the results obtained
from laboratory tests. However, the lack of a correlation between the data from short-
term and long-term tests has limited the use of abundant laboratory results for service
life design of LCC. This study presents an overview of results obtained when LCCs
were exposed to chlorides. The key outcome of this study is a broader synthesis of the
available data regarding the relationship between the mix design and the performance
of LCCs in various chloride environments, which helps find the possible correlation
and fully appreciate the value of the short-term tests.

Keywords Chloride penetration · Low carbon concrete · Marine structures

1 Introduction

Approximately 5–8% of the global CO2 emissions is attributed to the production of

ordinary Portland cement (OPC) [1]. To reduce the carbon footprint of the concrete
industry, substitution of cement by supplementary cementitious materials (SCM)
without compromising performance is an efficient solution [2]. Another indirect
way to contribute to sustainability is prolonging the service life of infrastructure by
using durable concrete. From these two viewpoints, ground granulated blast-furnace

C. Xue (B) · V. Sirivivatnanon

School of Civil and Environmental Engineering, University of Technology Sydney,
Ultimo, NSW, Australia
e-mail: Caihong.Xue@uts.edu.au
UTS-Boral Center for Sustainable Building, Sydney, NSW, Australia

© The Author(s) 2023 141

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_16
142 C. Xue and V. Sirivivatnanon

slag (GGBFS) and fly ash (FA) are great choices among other SCMs. One of the
major benefits of blending OPC with GGBFS or FA is improved resistance to chloride
penetration, which has been evidenced by both short-term laboratory tests and long-
term field tests. The former includes the widely used rapid chloride migration test
method standardized in ASTM C1202 [3] and the chloride diffusion test given in
NT Build 492 [4]. However, for application, concrete structures need to be designed
for a specific service life, and this requires long-term quantitative field performance
assessment, which is not always practical [5]. Therefore, finding the link between
the results from laboratory and field tests is important for promoting efficient use of
LCCs, but is nevertheless challenging because a high concentration of deleterious
species in laboratory tests could have already altered the deterioration processes and
the laboratory curing conditions deviate significantly from on-site conditions. The
purpose of this paper was to review of the factors affecting chloride penetration in
LCC made with GGBFS or FA and the correlations between results from different
test methods.

2 Chloride Penetration in LCC

2.1 Effect of SCM Content on Diffusion Coefficients

The replacement of cement with SCM is usually no more than 50% for GGBFS and
30% for FA, due to the reduction in strength with increasing SCM content, as shown
in Fig. 1, in which the short-term laboratory test results from Dhir et al. [6] were
adopted to demonstrate the effect of GGBFS or FA content on D as well as the 28 day
cube strength. It can be seen that the diffusion coefficients continuously decrease with
increasing GGBFS content up to 65%; in the range of 30–50%, D is insignificantly
affected by FA content but the strength reduction is more pronounced at higher FA
dosages. On the other hand, for high-volume FA (HVFA) concrete with >50% FA as
cement replacement and a considerably high amount of superplasticizer, the HVFA
concrete has proven to yield higher long-term strength and resistance against chloride
penetration than OPC concrete, despite the early-age properties of the former being
less competitive [7–10]. Thomas et al. [11, 12] found that HVFA concrete with 50%
FA had a significantly lower D and a slightly higher compressive strength than OPC
concrete after being exposed to the field marine environment for up to 10 years.
Moreover, Moffatt et al. [13] reported the D of a HVFA concrete with 56–58% FA
after 24 years of exposure to a harsh field environment where high tides and freeze–
thaw cycles occurred, was only 1.5 × 10–13 m2 /s compared with 3.6 × 10–12 m2 /s
for the counterpart OPC concrete.
Chloride Penetration in Low-Carbon Concrete with High Volume … 143

(a) GGBFS content

(b) FA content

Fig. 1 Effects of a ground granulated blast-furnace slag (GGBFS) and b fly ash (FA) content as
cement replacement on chloride penetration and other properties assessed by short-term laboratory
tests [6, 14, 15]
144 C. Xue and V. Sirivivatnanon

2.2 Effect of Curing Conditions on Chloride Penetration

It has been well established that partial replacement of cement by GGBFS or FA

improves the microstructure of concrete and thus the resistance to chloride pene-
tration. In LCC, the GGBFS or FA reacts with calcium hydroxide (formed by the
hydration of cement) and water to produce C–S–H and a portion of calcium alumi-
nate phases [16–18]. Additionally, depending on the specific surface area of particles,
GGBFS or FA can act as a filler to fill pores and as nucleation sites to enhance hydra-
tion [19, 20]. Apart from the products formed by normal cement hydration, the addi-
tional hydrates due to either enhanced hydration or hydration of GGBFS/FA reduce
capillary porosity (>30 nm) and thus block the chloride diffusion paths; however,
these benefits of GGBFS and FA depend on the curing conditions, especially for LCC
with GGBFS [21]. Moreover, it is possible that blended cement concrete will perform
no better than OPC concrete when structures are exposed to prolonged drying and
carbonation [22]. Figure 2 compares the D of concretes cured in wet and dry condi-
tions, from which it can be seen that OPC concrete with a low strength grade is more
sensitive to the curing conditions, and the influence of curing conditions diminishes
with increasing concrete strength and exposure duration [23, 24]. Irrespective of the
curing conditions, at the same grade the blended concretes consistently outperform
OPC concrete in resisting chloride penetration. In this regard, Bamforth [25] exam-
ined the D of dry-cured (indoor), membrane-cured and water-cured concrete blocks
(40 MPa at 28 days) located in the splash/spray zone on the south coast of the UK for
8 years, and found that the effect of curing conditions on the D of different concrete
mixes was inconsistent, but that 70% GGBFS-blended concrete yielded the lowest
average D compared with 30% FA-blended concrete and OPC concrete. Addition-
ally, a lower grade of blended cement concrete is more durable than a higher grade
OPC concrete at the later age, which was also confirmed by Thomas et al. [11].

2.3 Effect of Test Methods on Chloride Penetration

There are only limited published data on relating laboratory test results to long-
term field performance of concrete with regard to resistance to chloride penetration.
Figure 3 shows the correlation between D from a long-term field test (Dfield on the
x-axis) and counterpart D from laboratory diffusion tests (Dlab on the left y-axis) and
coulombs (right y-axis) from the rapid chloride permeability test (RCPT), using data
from previous studies [13, 26, 27]. Thomas et al. [26] conducted the RCPT as well as
diffusion tests (using 16.5% NaCl as per ASTM C1556) on uncontaminated GGBFS-
blended concrete exposed in the field to a tidal zone for 25 years, and calculated the
chloride diffusion coefficient (Dfield ) from the chloride profiles in the field-exposed
concrete. Moffatt et al. [13] obtained Dfield and coulombs (RCPT) of high-volume
FA concrete exposed to the marine environment for 19–24 years. Compared with the
results from the RCPT, Dlab from laboratory diffusion tests following the procedures
Chloride Penetration in Low-Carbon Concrete with High Volume … 145

Fig. 2 Effect of curing conditions on the D of different grades of concrete exposed to the tidal zone
of Cape Peninsula with water temperature between 12 and 15 °C. The wet-cured (wet) concrete
was exposed to 6 days’ moist curing (23 °C and 90% relative humidity) after demolding, while the
dry-cured (dry) concrete was stored in an open area (23 °C and 50% relative humidity). The concrete
was exposed to marine environment at age 28 days [23]. GGBFS, ground granulated blast-furnace
slag; FA, fly ash; OPC, ordinary Portland cement

given in ASTM C1556, NT Build 443 or other similar procedures, could better
indicate the ability of concrete to resist chloride penetration. Although there are
synergies between the two diffusion coefficients (Dlab and Dfield ), the quantitative
relationship between them varies. Figure 4 shows the correlation between Dlab (x-
axis) and coulombs from the RCPT (left y-axis), and the non-steady-state (Dnssm )
or steady-state (Dssm ) chloride migration coefficient (Dm on the right y-axis) from
accelerated migration tests reported in previous studies [28–32]. Note that these
previous studies used different NaCl concentrations and exposure durations, which
are summarized in Table 1. When Dlab was used as the reference, the Dm > 2 × 10–12
(m2 /s) and RCPT coulombs >800 could be more reliable for ranking concretes in
terms of the resistance to chloride penetration.
146 C. Xue and V. Sirivivatnanon

Fig. 3 Correlation between the D from marine exposure (Dfield ) and the D from laboratory diffusion
tests (Dlab ) and coulombs from the rapid chloride permeability test (RCPT) [13, 26, 27]

3 Conclusions

(1) Replacing cement with up to 65% GGBFS or 30% FA improves the resistance of
concrete to chloride penetration but decreases early-age strength development.
LCC with GGBFS or FA could achieve higher resistance to chloride penetra-
tion at equivalent strength or binder content as compared with OPC concrete,
indicating that efficient use of LCC requires a performance-based service life
design approach.
(2) The influence of curing conditions on chloride diffusion coefficients diminishes
with increasing concrete strength grade, but could be significant for low-strength
concrete. At strength grade ≥40 MPa, the difference in chloride diffusion coef-
ficients arising from the change in curing conditions is much smaller in LCC
than OPC.
(3) Although short-term laboratory diffusion and accelerated migration tests are
reliable in distinguishing parameters that affect chloride penetration in concrete,
the correlations between results from different methods are difficult to establish.
Chloride Penetration in Low-Carbon Concrete with High Volume … 147

Fig. 4 Correlation between the D from laboratory diffusion tests (Dlab ) and coulombs from the
rapid chloride permeability test (RCPT) and the chloride migration coefficient from migration tests
(Dm ) [28–32]. The gray lines indicate the RCPT results. The red lines indicate the steady-state
chloride migration coefficient (Dssm ) and blue lines indicate non-steady-state chloride migration
coefficient (Dnssm )

Table 1 Laboratory test methods for assessing resistance to chloride penetration shown in Fig. 4
Authors Laboratory diffusion tests Accelerated migration tests
Standard NaCl Days Standard NaCl Hours/voltage
Thomas ASTM C1543 3% 90
et al. [28]
Maes et al. NT Build 443 16.5% 30 NT Build 492 10% 24/30–60
[29]
Boddy et al. AASHTO T259 3% 90
[30]
Chiang and AASHTO T259 3% 90 ACMT 0.52 M 24/60
Yang [31]
Mao et al. NT Build 443 16.5% 90 NT Build 492 10% 24/60
[32]
Wang and 0.1 M 42–49 0.1 M 24–48/12
Lui [33]
148 C. Xue and V. Sirivivatnanon

Acknowledgements The authors acknowledge the support of the UTS-Boral Centre for Sustain-
able Building for the opportunity to review the performance-based testing of the resistance of LCC to
chloride penetration. This will promote greater and systematic use of LCC in marine environments.

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Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
A Compact Review on the Waste-Based
Lightweight Concrete: Advancement
and Possibilities

M. M. U. Islam, J. Li, R. Roychand, and M. Saberian

Abstract Lightweight concrete (LWC) has been used for more than 2000 years, and
the technical development of waste-based LWC is still proceeding. Notably, the very
first representative concrete mix of infrastructural LWC was introduced for building
a family house in Berlin, Germany, a few decades ago. The unique and distinc-
tive combination of waste-based LWC successfully creates an appealing alternative
to traditional concrete aggregates in terms of durability, robustness, cost, energy-
saving, transportation, environmental advantages, innovative architectural designs
and implementations, and ease of construction. Numerous researchers have attempted
to utilize waste materials to produce LWC, aiming to bring both ecological and
economical solutions to the construction industry over the past few decades. Waste
materials, such as crushed glass, waste tire rubber, masonry rubber, chip rubber,
plastics, coconut shells, palm oil fuel ash, palm kernel shells, fly ash, and rice husks,
possess lower specific gravity than traditional concrete aggregates. Thus waste-based
LWC can be a significant replacement for conventional raw materials (cementitious
material and aggregates) as it requires less strength than conventional concrete for
both structural and non-structural applications. Although waste-based LWC is well
recognized and has proven its scientific potential in a broad range of applications,
there are still uncertainties and hesitations in practice. Therefore, the primary objec-
tive of this study was to demonstrate the current state-of-the-art understanding and
advancement of waste-based LWC over the past decades. Furthermore, an equally
critical discussion is reported to shed light on the potential benefits of LWC. We
highlight how the performance of LWC has been enhanced significantly over the
period, and understanding of the properties of waste-based LWC has advanced.

Keywords Lightweight concrete · Lightweight concrete aggregates · Mechanical

properties · Structural bond behavior · Thermal conductivity · Waste materials

M. M. U. Islam · J. Li (B) · R. Roychand · M. Saberian

Discipline of Civil and Infrastructure Engineering, School of Engineering, RMIT University,
Melbourne, VIC, Australia
e-mail: jie.li@rmit.edu.au
M. M. U. Islam
e-mail: momeen.structure.um@gmail.com

© The Author(s) 2023 151

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_17
152 M. M. U. Islam et al.

1 Introduction

Knowledge of lightweight concrete (LWC) dates back almost 3000 years [1]. A
number of LWC-based superstructures can be found around the Mediterranean, the
most noteworthy of which are the Pantheon Dome and the Port of Cosa, both of which
were built in the early age of the Roman Empire. LWC has a considerable number
of advantages, such as better fire resistance, thermal insulation, and low density.
Notably, implementation of LWC has been extensively explored as a non-structural
and structural material.
In the past few decades, LWC has become a significant and resourceful material
that has been considerably developed, thanks to scientific endeavors [2, 3]. Indeed,
LWC is one of the exciting materials in the contemporary construction sector due to
its immense advantages, both ecologically and structurally. In the case of structural
advantages, LWC has crucial applications, particularly in heavy structure design,
where the dead load governs the total weight and that dead load is substantially
greater than the anticipated service load, such as for bridges and multistorey build-
ings. The lessened self-weight that originates from the utilization of the LWC in struc-
tures delivers accountable cost savings and flexibility. LWC also improves fire resis-
tance, and seismic structural response offers longer structural spans, reduces rein-
forcement ratios, and lowers the cost of foundation materials. Furthermore, precast
elements manufactured using LWC reduce placement and transportation costs [4, 5].
For bridges, LWC facilitates longer spans and more lanes. For instance, LWC can be
used on one side of a cantilever bridge, while normal-weight concrete (NWC) is used
on the other side to facilitate the weight balance for a longer span. From the ecolog-
ical aspect, LWC possesses lower thermal conductivity than NWC and hence plays
a significant role in saving energy when introduced as a thermal insulation material.
Moreover, implementation of LWC produced from controlled thermal lightweight
materials lessens the energy consumption by air acclimatizing in both warm and
cold countries. Nowadays, energy-shortage problems are escalating at an alarming
and upsetting rate, and it has become a worldwide concern. Most importantly, the
waste materials produced from the agricultural and industrial sectors can be utilized
to manufacture the LWC following an eco-friendly and economical approach, which
assists in mitigating climate change. In this review, we focused on compiling the
standpoints and previous scientific efforts related to LWC to improve knowledge of
the entire scenario and identify the research gaps. Notably, this compilation sheds
light on the advancement and possibilities of LWC to inspire new researchers for
further progress.
A Compact Review on the Waste-Based Lightweight Concrete … 153

2 Lightweight Aggregate Concrete (LWAC)

LWAC can be manufactured using artificial or natural lightweight aggregates (LWAs)

to substitute the conventional aggregates in NWC [6]. There are a number of LWAs
that possess various physical and mechanical characteristics allowing LWAC to be
produced with varying ranges of strengths and densities [7]. Different types of LWAs
are also commercially available, which has pushed researchers to investigate and
compare them in the production of high-strength LWACs [8]. As an example, several
natural materials (perlite or vermiculite), recycled waste materials (crushed glass,
waste tire rubber, masonry rubber, chip rubber, and plastics), and argillaceous mate-
rials (clay, slate, and shale) have been used by previous researchers in the manufac-
ture of LWAC. Moreover, a significant number of research has been conducted into
utilizing agricultural wastes, such as coconut shells, palm oil fuel ash, palm kernel
shells, and apricot shells. Including these wastes in the manufacture of LWAC brings
sustainable solutions to significant environmental issues [9, 10].
It is well known that the LWAs possess higher porosity than normal-weight aggre-
gates (NWAs) [11] and thus lower strength values together with larger deformations
[12]. These characteristics suggest that LWAs will be the weakest constituent, having
a significant role in the ultimate performance of the manufactured concrete mix.
Furthermore, LWAs usually constitute >50% total volume of the concrete [13]. It is
therefore crucial to carefully implement LWAs to improve the overall performance
of the LWAC mix in both the fresh and hardened phases. A significant number
of researchers have conducted detailed investigations of LWAs properties, such as
grading, absorption, and particle size, on the thermal and mechanical behaviors of
LWAC. Table 1 shows the requirements for compressive strength and splitting tensile
strength of LWAs according to ASTM C331 [14].

Table 1 Requirements for compressive and splitting tensile strength of lightweight aggregates
(LWAs)
Concrete type Dry density, kg/ Minimum splitting tensile Minimum compressive
m3 strength at 28 days (MPa) strength at 28 days (MPa)
LWA 1760 2.2 28
1680 2.1 21
1600 2.0 17
Mixture of 1840 2.3 28
LWAs and 1760 3.1 21
NWAs
1680 2.1 17
NWAs, normal-weight aggregates
154 M. M. U. Islam et al.

3 Structural Bond Behavior of LWC

It is necessary to have adequate bonding between the concrete matrix and reinforcing
bars for (i) gaining an efficient beam action, (ii) crack control, and (iii) improving
ductility [15]. Furthermore, all the empirical equations in the standard codes greatly
depend on sufficient bonding between the cement matrix and reinforcing bars [16].
Hence, a reduction in bonding may lead all the design basics towards invalida-
tion. There are two mechanisms for achieving improved bond strength: mechan-
ical (bearing and friction action) and physiochemical (linkage/adhesion) [17]. The
adhesion/linkage force originates from the chemical reaction between the surface of
the reinforcement and the cementitious matrix. The friction forces come from the
bearing force and rough contact, resulting from interlocking between the reinforcing
ribs and the concrete matrix [18].
Several researchers have investigated the bond strength behavior of LWC and
described the factors that can adversely affect the bond strength. These crucial factors,
such as the water–cement ratio (w/c), aggregate types, the diameter of the reinforce-
ment bars, admixtures, surface texture and type of reinforcing bars, types of lateral
confinement, and bond strength, significantly control the bond strength [19]. Many
equations have been developed to predict the bond strength of LWC, and three are
presented below [20–22]:
  2   
h h
τ = 171.9 − 24.2 + 1.29 f c' (1)
d d
 
37.5
− 9.4 f c'
0.5
τ= (2)
(d + ld )0.25
ρd
τ = K · [44.5 − 60(w/ c)] · (3)
2200
where h is the reinforcing rib height, d is the reinforcing bar diameter, ld is the length
of embedment, f ' c is the compressive strength, w/c is the water–cement ratio, and ρ d
is the dry density of LWC.

4 State of the Art

Extensive research on LWAC (Table 2) has led to many structural applications, such
as high-rise buildings, long-span bridges, and buildings where the foundation condi-
tions are vulnerable, and also in highly dedicated applications, such as offshore and
floating structures. Moisture content and density are the major factors controlling
the thermal conductivity properties of LWAC, whereas the mineralogical properties
Table 2 Engineering properties of waste-based lightweight aggregate concrete
Waste Replacement Aggregate Content Mechanical properties at 28 days Thermal Remarks References
material type size (%) Compressive Splitting Flexural Modulus of conductivity
strength tensile strength elasticity (W/m K)
(MPa) strength (MPa) (GPa)
(MPa)
Rice Cement Powder 0–35 42–24 2.6–1.8 4.1–2.7 32–29 – Decreased the [33]
husks mechanical
properties by
the addition
of rice husk
Cement Powder 0–20 68–48 5.1–4.4 8.1–6.9 – – 20% [34]
replacement
of cement
with rice husk
improved the
properties
Cement Powder 0–15 36–41 4.5–4.9 – – 1.21–0.99 Improved [35]
thermal
conductivity
A Compact Review on the Waste-Based Lightweight Concrete …

Cement Ultrafine 10 47 – – – – Improved [36]

compressive
strength
Palm oil Cement 300 µm 0–25 41.8–36 3.8–3.1 6.6–4.2 15.8–13.9 – 10% POFA [2]
fuel ash improved the
(POFA) mechanical
properties of
LWC
(continued)
155
Table 2 (continued)
156

Waste Replacement Aggregate Content Mechanical properties at 28 days Thermal Remarks References
material type size (%) Compressive Splitting Flexural Modulus of conductivity
strength tensile strength elasticity (W/m K)
(MPa) strength (MPa) (GPa)
(MPa)
FA 0.1–250 µm 0–15 27–35 2.2–3.3 2.1–3.9 60–78 – Addition of [37]
15% POFA
content
improved the
mechanical
properties
Cement 300 µm 0–30 32.6–24.9 2.8–2.6 – – Varied from Thermal [38]
27 to 40 °C conductivity
for 26 h improved for
up to 30%
POFA content
Oil palm CA 2.36–9 mm 100 41.8–36 3.8–3.1 6.6–4.2 15.8–13.9 – Adopting [2]
shell OPS as CA
(OPS) can reduce the
consumption
of NCA
(continued)
M. M. U. Islam et al.
Table 2 (continued)
Waste Replacement Aggregate Content Mechanical properties at 28 days Thermal Remarks References
material type size (%) Compressive Splitting Flexural Modulus of conductivity
strength tensile strength elasticity (W/m K)
(MPa) strength (MPa) (GPa)
(MPa)
CA ≤2.36 mm 100 2.1–20 – – – 0.4–0.9 Palm shell [39]
foamed
concrete
CA 8 mm 50 34.2–41.3 2.77–3.2 3.8–4.9 – – OPS [40]
improved the
properties
Waste CA 8–15 mm 100 18 3.34 3.74 15.28 – Proposed new [4]
tire method to
rubber improve the
mechanical
properties
FA ≤4.75 mm 5–25 26–41 3–3.8 3.6–4.4 – – Fracture [41]
energy
increased
A Compact Review on the Waste-Based Lightweight Concrete …

with the
addition of
rubber and
fibers
(continued)
157
Table 2 (continued)
158

Waste Replacement Aggregate Content Mechanical properties at 28 days Thermal Remarks References
material type size (%) Compressive Splitting Flexural Modulus of conductivity
strength tensile strength elasticity (W/m K)
(MPa) strength (MPa) (GPa)
(MPa)
FA 1–2 mm 0–20 20.53–33.94 2.25–3.7 3.42–5.5 21.95–33 – Improved [42]
ductile
properties
FA 1–3 mm 0–15 43–53 3.44–3.8 4.77–6.4 2.52–2.7 – Addition of [43]
rubber
decreased
properties
CA 5–20 mm 40 24.1–28.2 2.4–3.0 4.9–6.1 25.9–28 – Brittleness [44]
index reduced
due to the
addition of
rubber
FA 0.075–4 mm 0–20 22–20 – – – 1.0–0.9 Improved [13]
thermal
conductivity
Waste CA ≤5 mm 0–70 39–45 2.4–3.6 4.4–6.6 – – Decreased the [45]
glass properties
FA ≤2 mm 100 56–57 – – – 1.1–0.9 Improved [46]
conductivity
(continued)
M. M. U. Islam et al.
Table 2 (continued)
Waste Replacement Aggregate Content Mechanical properties at 28 days Thermal Remarks References
material type size (%) Compressive Splitting Flexural Modulus of conductivity
strength tensile strength elasticity (W/m K)
(MPa) strength (MPa) (GPa)
(MPa)
Coconut CA 4.7–12.5 mm 100 13.8–24 1.84–2.98 2.95–5.65 44–67 – Useful to [47]
shell produce
structural
LWC
CA, coarse aggregate; FA, fine aggregate
A Compact Review on the Waste-Based Lightweight Concrete …
159
160 M. M. U. Islam et al.

of the aggregates may affect up to 25% of the thermal conductivity for LWC under
a similar density value [3, 23]. The concrete matrix penetrates into the LWAs during
mixing of the concrete materials [24]. Nevertheless, the penetration rate dramati-
cally depends on the surface layer and microstructure of the aggregates, viscosity of
the concrete matrix, and particle size distribution of the cement. Additionally, both
the chemical and physical properties of LWAs influence the strength of the LWAC
because of the processes occurring at the interfacial transition zone. The compres-
sive strength of LWC increased from 15.5 to 29 MPa for an increase in cement
content from 250 to 350 kg/m3 by maintaining the same density of ≈1500 kg/m3
[25]. Figure 1a shows the correlation between cube strength and density of LWAC.
Another study [26] explored the compressive strength and thermal conductivity of
expanded perlite aggregate-based concrete along with the mineral admixtures. They
mentioned that using fly ash and silica fume as cementitious materials can reduce
the thermal conductivity values by up to 15%, while the compressive strength and
density of the concrete were also lowered by 30%. LWC using diatomite as the LWA
was manufactured with a density ranging from 950 to 1200 kg/m3 and compressive
strength from 3.5 to 6 MPa, and thermal conductivity was found to increase from
0.22 to 0.30 W/(m K) for a cement content of 250–400 kg/m3 . LWC associated with
expanded glass and clay as the LWAs exhibited higher resistance to chloride ion pene-
tration and water with a cement content of 500 kg/m3 and unit density of 1400 kg/
m3 , and the compressive strength of the LWAC reached 24 MPa at 28 days. LWAC
produced with dredged silt as the LWA exerted densities from 800 and 1500 kg/m3
for the varying binder content of 364, 452, and 516 kg/m3 [27, 28]. The dredged
silt-based LWAC exhibited compressive strength from 18 to 42 MPa with thermal
conductivity of 0.5–0.7 W/m K at 28 days. It was observed that the density of the
LWC crucially affects the strength of the concrete for similar cement and water
content. The thermal conductivity of LWC is significantly influenced by various
factors such as cement content, water content, and the type and content of LWA
[29, 30]. LWC bricks were manufactured using rice husk ash (RHA) and expanded
polystyrene as LWAs, and the maximum cement replacement by RHA was 10% by
weight [31]. The effect of zeolite inclusion for autoclaved concrete was investigated
by using aluminum to introduce a pore-forming agent, where zeolite was used with
a total content of 535 kg/m3 . The highest compressive strength of 3.3 MPa was
attained with 50% replacement, and the thermal conductivity was 0.18 W/(m K).
In another study, autoclaved concrete was produced using bottom-ash as the LWA
and as fractional replacement of cement, where the bulk density was ≈1400 kg/m3
with the increase in compressive strength ranging from 9 to 11.6 MPa, and thermal
conductivity varying from 0.5 to 0.61 W/(m K) [32]. Figure 1b shows the correlation
between thermal conductivity and dry density of LWAC.
A Compact Review on the Waste-Based Lightweight Concrete … 161

Fig. 1 Correlation between a cube strength and dry density for lightweight concretes with different
compositions at 28 days [48], and b thermal conductivity and dry density for LWC [49]

5 Possibilities

The compilation of previous research gives a clear indication that the thermal and
mechanical behaviors of LWAC have been extensively developed and investigated.
However, there have been few studies related to the structural applications of LWAC.
We suggest widening the scope of LWAC for investigation of its structural appli-
cation and behaviors, so complete guidelines can be provided to design engineers,
assisting them with all required design aids and data. Further studies, especially
on the mechanics of lightweight materials, would establish confidence and trust in
the potential applications of LWAC as a structural element. Another limitation (i.e.,
the higher time-dependent deformations of LWAC) requires further investigation.
Figure 2 shows some of the superstructures that have been built with LWAs.

Fig. 2 a LWC: Schweiz 2003 (Gartmann house), with density 1,100 kg/m3 , strength 12.9 MPa,
and thermal conductivity 0.32 W/m K [50], b Infra LWC: Berlin 2007 (Schlaich house) with density
760 kg/m3 , strength 7.4 MPa, and thermal conductivity 0.18 W/m K [51], c LWC: Stuttgart 2012
(house H36) with density 1,000 kg/m3 , strength 10.9 MPa, and thermal conductivity 0.23 W/m K
[50], d Infra LWC: TU Eindhoven 2015 (Pavilion) with density 780 kg/m3 , strength 10 MPa, and
thermal conductivity 0.13 W/m K [50]
162 M. M. U. Islam et al.

6 Conclusions

LWAC is an exceptionally versatile material that can be implemented in a broad range

of applications. Though LWAC has been used for the past two millennia, there are
still some limitations that have been reported and discussed in this review. Clear and
straightforward definitions have been given for different types of LWAC, and a critical
analysis of the mix design, aggregates properties, testing, and classifications of LWC
was presented. Several limitations of LWAC have been discussed and presented as
crucial information on the state of the art. Infra LWC is a new research area to be
explored. We recommend further essential studies to develop a strong design and
reliable construction methodology to confirm product quality.

Acknowledgements The authors acknowledge RMIT University, Melbourne, Australia, for

financial assistance received through RTP scholarships.

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Influence of Reinforcement
on the Loading Capacity of Geopolymer
Concrete Pipe

S. Dangol, J. Li, V. Sirivivatnanon, and P. Kidd

Abstract Geopolymer concrete is emerging as a sustainable construction material

due to utilization of industrial by-products, which greatly reduces its carbon foot-
print. Past studies of the mechanical properties and resistance to sulfuric acid reac-
tion of cement-less geopolymer concrete indicated its suitability for precast concrete
pipes over ordinary Portland cement (OPC) concrete. In the present study, a three-
dimensional finite element (FE) model of reinforced concrete pipe was developed
using commercial software ANSYS-LSDYNA. The load-carrying capacity of rein-
forced and non-reinforced geopolymer concrete pipes under the three-edge bearing
(TEB) test was investigated and compared with OPC concrete pipes. The results indi-
cated geopolymer concrete with comparable compressive strength to OPC concrete
showed higher loading capacity in a pipe structure due to its better tensile perfor-
mance. The effect of steel reinforcement area on the loading capacity of geopolymer
concrete pipes was quantitatively analyzed, and they met the specified strength
requirement for OPC concrete in the ASTM standard, with up to 20% reduction
in the reinforcement area.

Keywords Geopolymer concrete pipe · Pipe loading capacity · Numerical

modelling

1 Introduction

As an integral part of civil infrastructure, concrete pipes are used as conduit for
sewage and storm water. Ordinary Portland cement (OPC) concrete pipes have
demonstrated reliable long-term performance over years of usage. Their structural

S. Dangol · J. Li (B) · V. Sirivivatnanon

School of Civil and Environmental Engineering, University of Technology Sydney, Sydney,
NSW, Australia
e-mail: jun.li-2@uts.edu.au
P. Kidd
Cement Australia Pty. Ltd., Brisbane, QLD, Australia

© The Author(s) 2023 165

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_18
166 S. Dangol et al.

performance is evaluated in terms of test load (Dpeak ) corresponding to load causing a

0.3 mm crack and by ultimate load (Dult ) corresponding to the load supported before
failure of the pipe. The three-edge bearing (TEB) test is a standardized test described
in AS/NZS-4058 [1] and ASTM-C76M [2] to examine the mechanical strength of
a pipe, wherein a line load is applied to the crown of the pipe while the base of
the pipe is supported by two bearers. However, carrying out such destructive tests
is uneconomical and often inefficient, considering the need for human judgement of
crack formation during the TEB test. Hence, numerous researchers have performed
numerical modelling of concrete pipes to investigate the load-carrying capacity and
load–deflection behavior of concrete pipes.
de la Fuente et al. [3] and de Figueiredo et al. [4] simulated the TEB test for
steel fiber-reinforced concrete (SFRC) pipe using a MAP (mechanical analysis of
pipes) model to study mechanical behavior. The results from the numerical simulation
were compared with experimental results, concluding the efficiency of the numerical
model to design fiber-reinforced concrete pipes because the model gave an average
error of 7%, which was within the anticipated contingency. Similarly, Ferrado et al.
[5] used the commercial software ABAQUS to simulate SFRC pipes. In their study,
the behavior of the pipes was defined by compression and the uniaxial tension curve
based on theoretical formulation in the existing literature. The load–deflection curve
and stress distribution of the pipes from experimental and numerical analysis were
found to be in good agreement. Likewise, a numerical modelling of concrete pipes
with different diameter and reinforcement configuration, was conducted by Younis
et al. [6] to predict the service load and ultimate load. Following the concrete damage
plasticity (CDP) model equation developed by Alfarah et al. [7], the non-linear
behavior of concrete in compression and tension was defined. Based on their analysis
and experimental results, the average prediction error was ≈6% for both service load
and ultimate load, suggesting the reliability of numerical modelling for designing
concrete pipes.
With the growing interest in sustainable construction materials, numerous studies
of geopolymer concrete have been carried out, because it utilizes industrial by-
product as its source material. Geopolymer is a cement-less binder formed as a
result of reaction between aluminosilicate compounds with alkali [8–10]. Studies
exploring the material properties of geopolymer concrete have shown significant
development of strength at early age when cured at elevated temperature, high
compressive strength, high flexural strength, and resistance to chemical attack [8,
11, 12]. Results of studies conducted to investigate indirect tensile strength and flex-
ural strength of geopolymer have also showed higher indirect tensile strength and
flexural strength in comparison with OPC concrete of the same compressive strength
[13–16]. It is reported that the geopolymer possessed 1.4- and 1.6-fold higher value
for indirect tensile strength and flexural strength, respectively, compared with OPC
concrete [17]. Such properties of geopolymer concrete can be used to test the strength
of pipes with reduced reinforcement bars. Although much efforts has been made to
study the structural behavior of OPC SFRC to enhance the load-carrying capacity,
study of the structural performance of geopolymer concrete pipes has not been carried
out.
Influence of Reinforcement on the Loading Capacity of Geopolymer … 167

We investigated the load-carrying capacity of geopolymer concrete pipes based on

a three-dimensional (3D) finite element (FE) model developed to simulate the TEB
test. For this study, fly ash (FA)/slag-based powder form geopolymer GeocemTM
developed by Cement Australia with different ratio of FA and slag was utilized. The
general purpose geopolymer binder we termed “geocem 1” contained 50% of FA
and 32% of ground granulated blast-furnace slag (GGBFS), and the high strength
geopolymer binder termed “geocem 2” comprised 30% FA and 50% GGBFS. The
developed FE model was updated for both types of geopolymer concrete based
on their mechanical properties. Subsequently, the FE model was used to eval-
uate the load-carrying capacity of the geopolymer pipes and the effect of reduced
reinforcement area was evaluated.

2 FE Modelling of TEB Test

Our 3D FE model to simulate the TEB test for the pipe was based on commer-
cial software ANSYS LS-DYNA (Fig. 1). The model comprised three components:
concrete part, reinforcement steel bars, and bearing strips. Pipe of diameter 450 mm,
length 1000 mm, and wall thickness of 42 mm were modelled. The concrete pipe
and bearing strips were modelled using a 3D solid element (SOLID164). Similarly,
a beam element (BEAM161) was used to model the reinforcing steel bars. Discrete
steel formulation was used and perfect bond condition between the reinforcement bar
and concrete was assumed. The bearing strips were modelled to mimic the bound-
aries in the TEB test: the lower bearing strips were fixed at the bottom to prevent
translational and rotational degrees of freedom, and the upper bearing strips were
restricted in all directions except for vertical displacement movement to allow for
displacement-controlled loading on the pipe. The interaction between the pipe and
the bearing strips was defined by an automatic contact surface. For the simulation of
the test, displacement-controlled loading was defined as applied downward displace-
ment on the upper bearer. The load–deflection curve was obtained for the analyzed
pipe and is presented in terms of design load (N/m/mm) as specified in ASTM-C76M
[2] and deflection in millimeters.

3 Material Modelling

Due to the complex material behavior of concrete, which includes elastic, non-linear
plastic behavior, and material damage, the available concrete damage models for
numerical modelling of concrete structures are often quite complex because these
material models often contain parameters for which values are difficult to obtain
from simple tests or have only mathematical meaning and no physical meaning [18].
To date, there are a lot of material models available to simulate concrete damage
behavior [19, 20]. Among them, one simple concrete damage model implemented in
168 S. Dangol et al.

Fig. 1 a Three-edge bearing (TEB) test setup; b finite element model of concrete pipe for TEB
test simulation

LS-DYNA to model concrete behavior is the Karagozian and Case (K&C) concrete
model (Fig. 2). A key merit of the K&C concrete model for numerical simulation
of concrete behavior is its reliance on just one main input parameter of unconfined
compressive strength. Schwer & Malvar [21] stated that the K&C concrete model can
be utilized for analysis involving new concrete materials with no detailed information
available to characterize the concrete beside its compressive strength, owing to the
fact that the unconfined compressive strength of the concrete not only describes the
elastic response, but also accounts for the plastic response including shear failure,
compression, and tensile failure.

Fig. 2 Three failure surface of K&C concrete model [22]

Influence of Reinforcement on the Loading Capacity of Geopolymer … 169

The material constitutive behavior of the K&C concrete model comprises three
parts; for initial loading, the stress is elastic until it reaches the yielding point, after
which it increases further till the limit surface, called the maximum yield surface.
Following the maximum yield surface, perfectly plastic, or softening behavior up
to the residual yield surface is observed. These shear failure surfaces are mutually
independent and can be formulated as [22, 23]:
p
Fi ( p) = a0i + (1)
a1i + a2i p

where i stands for either yield strength surface (y), maximum strength surface (m)
or residual strength surface (r), p is the pressure calculated as −I3 1 , and the variables
aji ( j = 0, 1, 2) are the parameters calibrated from test data.
The resulting failure surface is interpolated between the maximum strength surface
and either the yield surface or the residual strength surface as per the following
equations:
   
F(I1 , J2 , J3 ) = r (J3 ) η(λ) Fm ( p) − Fy ( p) + Fy ( p) for λ ≤ λm (2)

= r (J3 )[η (λ) (Fm ( p) − Fr ( p)) + Fr ( p)] for λ ≥ λm (3)

where I 1 , J 2 , and J 3 are the first, second, and third invariants of deviatoric stress
tensor, λ is the modified effective plastic strain or the internal damage parameter,
η(λ) is the function of the internal damage parameter λ, with η(0) = 0, η(λm ) = 1
and η(λ ≥ λm ) = 0, and r(J 3 ) is the scale factor in the form of the William–Warnke
equation [24].
The K&C concrete model considers the effect of strain rate, failure, and different
mechanical–physical properties in compression and tension and hence is suitable
for concrete modelling [18]. Based on the uniaxial compressive strength, material
parameters are generated, requiring definition of only a few parameters for the func-
tionality of the material model, and more parameters can be defined if required. The
model requires 49 parameters to be defined, as well as equation of state, which is
complicated because many parameters have only mathematical meaning. Hence, the
developers advocate using parameter generation if the data to define the material are
not available. The default parameters in the K&C concrete model were calibrated
using uniaxial, biaxial, and tri-axial test data available for well characterized concrete
and using the relationship such as tensile strength or modulus of elastic as the func-
tion of compressive strength [21]. Hence, the K&C concrete model was used for both
OPC concrete and geopolymer concrete modelling for FE analysis (FEA).
For the reinforcement bar, an elastic–plastic constitutive relationship for rein-
forcement bar, with or without strain hardening, is commonly adopted for numerical
analysis. However, the elastic-perfectly plastic assumption shown in Fig. 3a often
fails to capture the steel stress at high strain, and accurate assessment of the strength
170 S. Dangol et al.

Fig. 3 Typical stress–strain curve of steel reinforcement representing a ideal elastic-perfectly

plastic model; b bilinear elastic–plastic model with linear strain hardening [25]

of structure at large deformation cannot be made [26]. Hence, more accurate ideal-
ization of the stress–strain curve as shown in Fig. 3b was used. The Piecewise Linear
Plasticity model used to represent the steel reinforcement behavior in LS-DYNA
considers the plastic deformation, strain rate effects and failure [19]. In the Piece-
wise Linear Plasticity model, the stress–strain curve for the reinforcing steel is treated
as bilinear by defining the tangent modulus [27]. The steel response is thus defined
by parameters such as Young’s modulus (E s ), yield strength ( f sy ) and hardening
modulus (E st ). The magnitude of E st in the plastic regimen is commonly set at 1%
of E s [28, 29].

4 Load–Deflection Behavior of Concrete Pipes

A concrete pipe model was used for our study of the load–deflection behavior of
geopolymer concrete pipes with respect to OPC concrete pipes. For the design of
450 mm reinforced concrete pipe, a reinforcement area of 175 mm2 /m was adopted
based on minimum reinforcement area criteria for class II 450 mm concrete pipe
defined in ASTM-C76M [2] in order to meet the design load criteria. Table 1 provides
the details of the mechanical properties of the materials for the study based on the
experimental results.
Comparing the results obtained from numerical analysis with the design require-
ment specified for OPC concrete of 50 N/m/mm and 75 N/m/mm for peak and
ultimate load respectively in ASTM-C76M [2] with the geocem 1 and geocem 2
FEA results, it was evident from the load–deflection curve shown in Fig. 4 that the
geopolymer concrete exhibited better load-carrying capacity than OPC concrete. The
peak load and ultimate load value of the OPC pipe were 85 N/m/mm and 113 N/m/mm
respectively, and for the geocem 1 and geocem 2 pipes, the peak load value was 100 N/
m/mm and 105 N/m/mm and the ultimate load value was observed to be 119 N/m/
Influence of Reinforcement on the Loading Capacity of Geopolymer … 171

Table 1 Material properties

Material property Geocem 1 Geocem 2 OPC
of the pipe model
Compressive strength (MPa) 50 50 50
Tensile strength (MPa) 3.6 3.8 2.3
Poisson’s ratio 0.2 0.2 0.2
Rebar yielding stress (MPa) 500 500 500
OPC, ordinary Portland cement

Fig. 4 Load–deflection
curves of the 450 mm pipes

mm and 121 N/m/mm respectively. The load-carrying capacity of both geopolymer

concretes outperformed the OPC concrete load requirement by ≈15% for peak load
capacity and ≈5% for ultimate load capacity. The high tensile strength of geopolymer
concrete benefited the load-bearing capacity of the pipe.
In the load–deflection plot of the pipe, a drop in the load was noticed after the
pipe reached its peak load capacity. Such a drop in loading capacity for a single-cage
model was also observed by Tehrani [30], Peyvandi et al. [31], and Younis et al. [6].
The rise in the load capacity following the drop after peak load signifies the load
stress being carried by the steel reinforcement.

5 Effect of Change in Reinforcement Area

To investigate the effect of steel reinforcement area on the load–deflection behavior of

the reinforced OPC, geocem 1 and geocem 2 concrete pipes, the steel reinforcement
was reduced by 20%, 40% and 50% from the total reinforcement area. The main
objective was to the test the load capacity in geopolymer concrete pipes under reduced
172 S. Dangol et al.

reinforcement and evaluate against the strength requirement as specified in ASTM-

C76M [2].
Figure 5 shows the load–deflection plot of 450 mm pipes under reduced rein-
forcement conditions. It is obvious for both geopolymer concrete types (Fig. 5a, b)
that the load-carrying capacity of the pipes decreased with the reduction in the rein-
forcement area, especially the ultimate load-carrying capacity. As the development
of crack in concrete structures is dependent on the tensile strength of the concrete,
changing the steel reinforcement does not change the service load capacity of the
pipe, but alters the ultimate load capacity. When sufficient tensile stress develops in
the concrete surface causing the concrete to crack, the pipe loses its capacity, which
is marked by the drop in the load capacity. As the tensile stress in the concrete pipe
is transferred to the steel reinforcement, regaining load capacity is observed until it
reaches its ultimate load capacity, after which the pipe fails. With the reduction of
reinforcement area by 40 and 50%, a significant reduction in the post-crack loading
capacity of the pipe was observed; however, the reduction of reinforcement area by
20% showed decrease ultimate load capacity by 7% and 6% for geocem 1 and geocem
2, respectively, but still satisfied the design requirement specified in the ASTM stan-
dard. The strength gained by fully reinforced OPC concrete pipe was approximately
equivalent to the strength gained by geopolymer concrete pipes with 20% reduced
reinforcement area. This result suggested the geopolymer concrete pipes could resist
load without failure with up to 20% reduction in the area of steel reinforcement.
Furthermore, in the comparative study of the load–deflection behavior of unre-
inforced OPC concrete pipe against unreinforced geocem 1 concrete pipe shown in
Fig. 5c, the test load capacity of the unreinforced OPC concrete pipe was 69 N/m/
mm and the test load capacity of geocem 1 was 96 N/m/mm. AS/NZS-4058 specifies
the test load for 450 mm unreinforced concrete pipe as 30 KN/m (or 67 N/m/mm)
[1]. The load capacity of the geocem 1 pipe was 37% higher than that of the OPC
concrete pipe. Moreover, considering the AS/NZS-4058 requirement of 67 N/m/mm
as the test load and 100 N/m/mm as the ultimate load capacity for class 3 reinforced
450 mm concrete pipe [1], unreinforced geopolymer concrete pipe can be used to
meet the design requirement of class 3 reinforced concrete pipe.

Fig. 5 a–c Effect on load–deflection behavior of 450 mm pipes due to change in reinforcement
steel area
Influence of Reinforcement on the Loading Capacity of Geopolymer … 173

6 Conclusion

We developed a 3D FE model of 450 mm concrete pipe to simulate the TEB test

in LS-DYNA. The K&C concrete model was used to characterize the OPC and
geopolymer concrete behavior, and a bilinear elastic–plastic model with linear strain
hardening was used to model steel reinforcement in the pipe FE model. In the compar-
ative analysis of geopolymer and OPC concrete pipes, the peak load and ultimate
load capacity exhibited by both types of geopolymer concrete pipes were notice-
ably higher than by OPC concrete pipe, which can be attributed to the high tensile
strength of geopolymer concrete. The FE model was used to study the effect of
changing the reinforcement area on the load-carrying capacity of geopolymer and
OPC concrete pipes. The change in steel reinforcement area affected the post-crack
behavior of the pipes, because reinforcement mostly contributes to strength develop-
ment in cracked concrete sections. Based on the numerical study, the reinforcement
bar area in concrete pipes can be reduced by 20% and still meet the specified design
load criteria. Further, comparing the loading capacity of unreinforced geopolymer
concrete pipe against OPC concrete pipe revealed that the unreinforced geopolymer
concrete pipe can satisfy the load criteria for AS/NZS class 3 450 mm reinforced
concrete pipe. Thus, with the use of geopolymer concrete, reinforcement requirement
for concrete pipe can be reduced to a certain percentage and still meet the specified
design load criteria.

Acknowledgements This research was funded through an Australian Research Council Research
Hub for Nanoscience Based Construction Materials Manufacturing (NANOCOMM) with the
support of Cement Australia. The authors are grateful for the financial support of the Australian
Research Council (IH150100006) in conducting this study.

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Creep of Slag Blended Cement Concrete
with and Without Activator

H. T. Thanh, M. J. Tapas, J. Chandler, and V. Sirivivatnanon

Abstract Partly replacing Portland cement (PC) with lower carbon footprint cemen-
titious materials such as ground granulated blast furnace slag (slag) is considered as
a practical method for reducing CO2 emissions in the cement concrete industry. To
mitigate the slow reactivity of slag in a cementitious system and enhance early-age
strength, the addition of a chemical activator is a solution. However, the effect of the
activator on creep behaviour of slag-blended cement concretes remains unclear. This
work presents the effect of sodium sulfate (Na2 SO4 ) activator on the compressive
creep of PC concrete blended with 50 and 70 wt% slag. Four concrete mixes (with
and without 2.5% Na2 SO4 activator) containing 395 kg of cementitious material
were prepared. The creep strain measurements were conducted on 150 × 300 mm
cylindrical specimens for 140 days under sustained compressive load. The results
showed that the 70% slag concrete had lower creep strain than 50% slag-blended
cement concrete. The presence of Na2 SO4 helped reduce the creep strain of 50%
slag concrete but slightly increased that of 70% slag-blended cement concrete. In
addition, the applicability of the predictive model in AS3600:2018 for the creep
behaviour of high slag content concrete was assessed.

Keywords Activator · Compressive creep · Slag concrete

1 Introduction

Cement-based materials have an essential role in civil infrastructure worldwide.

However, the production of clinker, the major constituent of Portland cement (PC),
accounts for ≈7% of global CO2 emissions [1], which has a destructive impact on

H. T. Thanh (B) · M. J. Tapas · V. Sirivivatnanon

School of Civil and Environmental Engineering, University of Technology Sydney (UTS),
Ultimo, NSW, Australia
e-mail: Hai.Tran@uts.edu.au
J. Chandler
Boral Limited, North Ryde, NSW, Australia

© The Author(s) 2023 177

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_19
178 H. T. Thanh et al.

environmental sustainability. Reducing the clinker to cement ratio by partly replacing

PC with lower carbon footprint cementitious materials such as ground granulated
blast furnace slag (GGBSF or slag) is considered a practical technique for reducing
CO2 emissions in the cement industry sector. The direct cumulative CO2 emissions
reduction by this method is estimated to reach almost 37% by 2050 [2]. Using slag
in cement concrete also enhances its compressive and flexural strengths at later ages
and improves the durability properties of concrete [3, 4].
As a pozzolanic-hydraulic activity or latent hydraulic behaviour, slag only
provides additional surface area to enhance PC hydration by filler effect in the early
days of hydration, without chemical reaction [5]. Hence, at high levels of slag replace-
ment, the slow reactivity and dissolution of slag in the cementitious system causes
slower strength gain at early ages. To mitigate this, adding an activator such as
sodium sulfate (Na2 SO4 ) is adopted [6, 7]. This addition accelerates slag’s reactivity
and enhances the early strength. However, in the presence of Na2 SO4 , a decrease in
the compressive strength growth rate and increased capillary porosity at a later age
have been reported [8, 9].
Creep of concrete is a time-dependent property that significantly affects the perfor-
mance of concrete structures. Generally, creep in concrete includes basic and drying
creeps. Under constant stress, gel particles slide and consolidate as water molecules
rearrange in capillary pore structures, and microcracks form in the interfacial tran-
sition zones between the aggregate and cement paste. These phenomena primarily
govern the basic creep of concrete [10]. In terms of creep of cement concrete using
slag as a cement replacement, only a few conducted studies are reported. Khatri et al.
[11] and Shariq et al. [10] concluded that creep of slag-blended cement concrete
increased when the slag replacement increased. In addition, the creep strains were
higher in slag cement concrete compared with plain concrete [10]. However, the creep
behaviour of high slag-blended cement concretes with the presence of an activator
has not yet been reported.
To understand the effect of Na2 SO4 activator on the creep behaviour of slag-
blended cement concrete, we conducted experiments on compressive creep. In this
study, four mixes were cast with 50 and 70 wt% slag replacement. A dose of
2.5 wt% by total binder of Na2 SO4 activator was used for activating two of the
four slag-blended cement concrete mixes. The measurement data were analyzed and
compared with the obtained values from the predictive model in AS3600:2018 to
assess the capability of model prediction for creep behavior of high slag-blended
cement concrete.
Creep of Slag Blended Cement Concrete with and Without Activator 179

Table 1 Concrete mix proportions and properties

Mix ID W/B Cement Slag Aggregate Sand HRWR Slump f c 28 days MoE
(kg/m3 ) (kg/m3 ) (kg/m3 ) (kg/m3 ) mL/m3 (mm) (MPa) (GPa)
M1_50 0.54 197.5 197.5 1035 850 1685 140 43.0 36.3
M2_50A 0.51 197.5 197.5 1035 850 2130 120 43.0 44.6
M3_70 0.52 118.5 276.5 1035 850 1685 110 38.0 35.8
M4_70A 0.52 118.5 276.5 1035 850 2130 130 35.5 35.7
f c , compressive strength; HRWR, high-range water reducer; MoE, modulus of elasticity

2 Experiments

2.1 Mixture Proportions and Properties

The four concrete mixes containing 395 kg of cementitious material were prepared
using shrinkage limited (SL) cement and slag provided by Boral Cement (Maldon,
Australia). The mix proportions and mix IDs are shown in Table 1. In the mix ID,
the numbers 50 and 70 signify the replacement percentage of slag, and the letter A
denotes the use of 2.5% Na2 SO4 activator for that mix.
Four cylinders of 150 × 300 mm and seven cylinders of 100 × 200 mm were cast
for each concrete mix. Of the four 150 × 300 mm cylinders, two were used for the
creep test, and two were used as control samples for measuring free drying shrinkage
strain. The compressive strength test and the determination of elastic modulus were
conducted on the 100 × 200 mm cylinders. The average results of these two tests
are shown in Table 1.

2.2 Creep Test Setup

Before the testing day, demountable mechanical gauge (DEMEC) points were
attached with 200 mm gauge length in three-gauge lines that were uniformly drawn
around the perimeter of each cylinder. The DEMEC strain gauge is used for measuring
creep/shrinkage strains of specimens. The vertical alignment of specimens and creep
rigs were carefully checked with water level instruments before applying load.
Figure 1a shows the creep rig used to perform compressive creep tests, and Fig. 1b
presents the controlled samples for free drying shrinkage measurements.
One creep rig was used to test two mixes of the same replacement percentage of
slag with and without activator. The loading started at 28 days, and the applied load
was set at 40% of the average 28 day compressive strength of the concretes. Because
two mixes were tested within a creep rig, the load was applied at 40% of the lower
strength mix. The tests were conducted in a controlled environment (23 ± 2 °C and
180 H. T. Thanh et al.

Fig. 1 a Creep rigs and loaded samples, b control samples

relative humidity 50%). During the testing period, the applied load was monitored
by a pressure gauge and maintained at a value not less than 5% of its initial value.
The referenced readings were obtained just before applying load. The creep/
shrinkage strains were recorded immediately after loading, 2 and 6 h for the first
loading day, daily for the first week, weekly until 2 months, and then monthly. Here
we report the findings after sustained load had been applied for 140 days. The creep
strains of each mix were averaged from two specimens and three-gauge lines of each
specimen.

3 Creep Model in AS3600:2018

As creep of concrete is a time-dependent behavior, using an empirical model to predict

the creep strain of concrete is a practical and useful approach for designing concrete
structures. From the viewpoint of building construction, it is not practical to perform
a long-term creep test for every concrete mix. In AS3600:2018, the predictive creep
strain of a sample bearing constant sustained stress σo at any time t can be generally
expressed as:
Creep of Slag Blended Cement Concrete with and Without Activator 181

( ) σo
εcc (t) = ϕcc t, tload , f c' , th , E env , ϕcc.b (1)
Ec

where ϕcc is the time-dependent creep coefficient function depending on the loading
time tload , characteristic compressive strength f c' at 28 days, hypothetical thickness of
sample th and the exposure environmental condition E env . The basic creep coefficient
ϕcc.b and elastic modulus E c of a sample loaded at 28 days can be determined based on
the characteristic compressive strength f c' . Assuming that we only know the 28 day
compressive strength of the four mixes described in Sect. 2.1, the characteristic
compressive strength f c' was selected equal to 30 MPa and 35 MPa for 50% and 70%
slag-blended cement concretes, respectively. Hence, in the predictive model, the
values of E c = 29.1 GPa and ϕcc.b = 3.6 were used for 50% slag-blended cement
concretes, and E c = 31.1 GPa and ϕcc.b = 3.2 were used for 70% slag-blended
cement concretes. In addition, the interior environment (E env = 0.65), which was
nearest to the test environment, was also selected.

4 Results and Discussion

4.1 Total Deformation

The total deformation due to elastic, creep, and shrinkage strains of the four mixes are
plotted in Figs. 2 and 3. It can be observed that the presence of Na2 SO4 significantly
reduced the creep and shrinkage strains of the 50% slag-blended cement concrete. A
similar trend can also be observed with the shrinkage of 70% slag-blended cement
concrete, but the total deformation of the loaded samples is comparable. A significant
difference in deformation at initial loading due to the higher elastic modulus in the
activated mix was also notable in the 50% slag-blended cement concrete mixes. The
creep strain rate of the two mixes was, however, almost identical.
In addition, it can be seen that the 70% slag-blended cement concrete exhibited
lower creep strain than the 50% slag-blended cement concrete. At 140 days, the creep
strains of the 50% slag-blended cement concretes was ≈1400 µm/m, while those
of 70% slag-blended cement concretes were only slightly greater than 1000 µm/m.
This finding contradicts the reports in [10].

4.2 Creep Strain Measurement and Model Prediction

The creep strains (including instantaneous elastic strain) were obtained by subtracting
the shrinkage strain from the total deformation of the loaded specimens. The devel-
opment of creep strains of the four tested mixes are shown in Figs. 4 and 5. Although
182 H. T. Thanh et al.

1800
1600 Total deformation
1400
Total strain ( m/m)

1200
1000 M1_ 50
800 M2_ 50A
600 M1_ 50 (Shrinkage)
M2_ 50A (Shrinkage)
400
200
0
0 20 40 60 80 100 120 140
Measurement duration (days)

Fig. 2 Total deformation including creep and shrinkage strains of 50% slag-blended cement
concrete, with and without Na2 SO4

1800
1600
1400
Total strain ( m/m)

Total deformation
1200
1000
M3_ 70
800 M4_ 70A
600 M3_ 70 (Shrinkage)
400 M4_ 70A (Shrinkage)

200
0
0 20 40 60 80 100 120 140
Measurement duration (days)

Fig. 3 Total deformation including creep and shrinkage strains of 70% slag-blended cement
concrete, with and without Na2 SO4

the presence of Na2 SO4 appeared to slightly reduce the creep strains of 50% slag-
blended cement concrete, an increase, though within measurement error range, in
that of 70% slag-blended cement concrete was observed.
The predictive results from the creep model in AS3600:2018 are also plotted in
Figs. 4 and 5. It can be seen that the predictive values from the model are lower than
the measured values for all mixes. Except for the initial deformation or instantaneous
elastic strain, the model underestimated the development of creep strain for both 50
and 70 wt% slag replacement cement concretes.
Creep of Slag Blended Cement Concrete with and Without Activator 183

1600
1400
1200
Creep strain ( m/m)

1000
800
600 M1_ 50
400 M2_ 50A
200 AS3600:2018

0
0 20 40 60 80 100 120 140
Measurement duration (days)

Fig. 4 Creep strain of concretes with 50 wt% slag replacement and predictive results

1600
1400
1200
Creep strain ( m/m)

1000
800
600
M3_ 70
400 M4_ 70A
200 AS3600:2018

0
0 20 40 60 80 100 120 140
Measurement duration (days)

Fig. 5 Creep strain of the concrete with 70 wt% slag replacement and predictive results

5 Conclusions

From the results of 140 day compressive creep testing of PC concrete blended with
50 and 70 wt% slag, the main findings can be summarized as:
(1) 70% slag-blended cement concrete had lower creep strain than 50% slag-
blended cement concrete, which suggested that increasing the slag content
improved creep performance
(2) adding 2.5% Na2 SO4 activator had negligible effects on the creep strain of
slag-blended cement concretes
(3) the activator slightly decreased the shrinkage strain of slag-blended cement
concretes
184 H. T. Thanh et al.

(4) the predictive model in AS3600:2018 underestimated the creep strain of high
slag concretes regardless of the presence of Na2 SO4 .
The monitoring of long-term creep behavior of these slag-cement concretes is
ongoing, and the results will be reported in the future.

Acknowledgements This study was supported by the UTS-Boral Centre for Sustainable Building
under funding from the Innovative Manufacturing CRC (IMCRC).

References

1. Voldsund M, Gardarsdottir SO, De Lena E, Pérez-Calvo J-F, Jamali A, Berstad D, Fu C, Romano

M, Roussanaly S, Anantharaman R (2019) Comparison of technologies for CO2 capture from
cement production—part 1: technical evaluation. Energies 12(3):559
2. Plaza MG, Martínez S, Rubiera F (2020) CO2 capture, use, and storage in the cement industry:
state of the art and expectations. Energies 13(21):5692
3. Özbay E, Erdemir M, Durmuş Hİ (2016) Utilization and efficiency of ground granulated blast
furnace slag on concrete properties—a review. Constr Build Mater 105:423–434
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5. Joseph S, Cizer Ö (2022) Hydration of hybrid cements at low temperatures: a study on Portland
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SJ, Waite TD (2021) A microstructural investigation of a Na2 SO4 activated cement-slag blend.
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220
Creep of Slag Blended Cement Concrete with and Without Activator 185

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Partially-Unzipped Carbon Nanotubes
as Low-Concentration Amendment
for Cement Paste

S. Iffat, F. Matta, J. Gaillard, M. Elvington, M. Sikder, M. Baalousha,

S. Tinkey, and J. Meany

Abstract Partially-unzipped multiwalled carbon nanotubes (PUCNTs) combine the

chemical structure and basic mechanical properties of multiwalled carbon nanotubes
(MWCNTs) and graphene nanoplatelets (GNPs) with exceptionally high graphene-
edge content. As a result, PUCNTs can be effectively oxidized for dispersion in
aqueous solutions and have a specific surface area that is larger than that of MWCNTs
with a comparable aspect ratio. Thus, the incorporation of relatively small concentra-
tions of PUCNTs in cement composites may result in significant physicomechanical
enhancements. In the proof-of-concept study presented here, cement paste specimens
were manufactured with oxidized PUCNT concentrations of 0, 0.001, and 0.005%
in weight of cement (wt%), that is, one order of magnitude smaller than lower-bound
concentrations for MWCNTs reported in the literature. Stable dispersion in water
was verified through dynamic light scattering analysis. Physicomechanical changes
and PUCNT dispersion in the cement matrix were investigated through uniaxial

S. Iffat (B) · F. Matta

Department of Civil and Environmental Engineering, University of South Carolina, Columbia,
SC, USA
e-mail: iffats@farmingdale.edu
S. Iffat
Department of Civil Engineering, Bangladesh University of Engineering and Technology,
Dhaka, Bangladesh
Department of Civil Engineering Technology, Farmingdale State College (SUNY), Farmingdale,
NY, USA
J. Gaillard · M. Elvington · S. Tinkey
Savannah River National Laboratory, Aiken, SC, USA
M. Sikder
Solenis, Wilmington, DE, USA
M. Baalousha
Center for Environmental Nanoscience and Risk, University of South Carolina, Columbia,
SC, USA
J. Meany
Ionic Flask Materials Group, Atlanta, GA, USA

© The Author(s) 2023 187

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_20
188 S. Iffat et al.

compression tests on 25 × 25 × 76 mm prism specimens, and visual inspection of

scanning electron microscopy (SEM) micrographs, respectively. The incorporation
of 0.001 wt% and 0.005 wt% of PUCNTs resulted in an average increase in compres-
sive strength of 10% and 29%, respectively, compared with plain cement paste. In
both instances, representative SEM micrographs show the preferential formation of
cement hydrates in cement paste that accorded with well-dispersed PUCNTs.

Keywords Cement paste · Dispersion · Partially-unzipped multiwalled carbon

nanotubes · Nanoparticles · Strength

1 Introduction

Partially-unzipped multiwalled carbon nanotubes (PUCNTs) are synthesized by

longitudinal unzipping of the outer walls of multiwalled carbon nanotubes
(MWCNTs) [1, 2], resulting in a portion having a morphology similar to that of
graphene nanoplatelets (GNPs). MWCNTs are characterized by a relatively high
aspect ratio, which may be beneficial when using these nanoparticles as reinforce-
ment in cement composites [3]. Mechanical enhancements have also been reported
for nano-amended cement composites with GNP concentrations between 0.03% [4]
and 0.5% [5] in weight of cement (wt%). Related mechanisms may be associated
with improved cement hydration through the presence of carboxyl and hydroxyl
functional groups on the surface of oxidized MWCNTs and GNPs [6, 7], with a
reduction in total porosity [8, 9], pore-size refinement [8, 10], and densification of
the cement matrix [11].
PUCNTs have appealing properties compared with their fully-unzipped counter-
parts; elongated strips of graphene with more functional group-edge sites compared
with MWCNTs or GNPs, but also a smaller aspect ratio [12]. Partial unzipping of
the outer walls of MWCNTs results in a radical increase in the open surface area
[12], and oxidation produces a relatively larger amount of functional groups [1], thus
enhancing the chemical affinity with the cement matrix. Through this process, a few
outer layers of a given MWCNT are partially unzipped without affecting the inner
core walls, resulting in a hybrid graphitic structure together with the inner (intact)
carbon nanotubes [13]. Therefore, the resulting PUCNTs exhibit: (a) larger surface
area and greater functionality than their precursor MWCNTs, which may enhance
dispersibility in aqueous solutions and in the cement matrix, and (b) a relatively high
aspect ratio, which may enhance the mechanical properties of the cement composite.
We present a feasibility study of cement paste amended with low concentrations of
oxidized PUCNTs, namely, 0.001 and 0.005 wt%, which is one order of magnitude
smaller than successful lower-bound concentrations reported in the literature for
MWCNTs and GNPs.
Partially-Unzipped Carbon Nanotubes as Low-Concentration … 189

Fig. 1 Partially-unzipped multiwalled carbon nanotubes (PUCNTs) used in cement paste speci-
mens: a scanning electron microscopy image showing PUCNTs resulting from oxidative unzipping
of multiwalled carbon nanotubes; b aqueous suspensions

2 Methods

2.1 PUCNT Aqueous Suspensions

The PUCNTs were prepared at Savannah River National Laboratory (Aiken, SC,
USA) using an oxidative unzipping process modified from Kosynkin et al. [14].
Defects in the pristine MWCNTs were created by soaking in concentrated sulfuric
acid and o-phosphoric acid, with potassium permanganate serving as the oxidant.
By controlling the amount of oxidant, heating time, and soaking time, the PUCNTs
shown in Fig. 1a were produced, and then suspended in deionized (DI) water. Stock
solution of 3 g/L of these PUCNTs in DI water was diluted with additional DI
water to prepare 0.02 and 0.1 g/L suspensions, to be used for the manufacturing of
cement paste specimens with PUCNT concentrations of 0.001 wt% and 0.005 wt%,
respectively. A small amount of NaOH (<0.01 mol/L) was added to the suspensions
to reach an approximate pH of 12, thereby enhancing the stability of acid-treated
(i.e., low pH) PUCNTs in water. The suspensions were ultrasonicated for 15 min
using an ultrasonic bath sonicator (model CPX 2008, Branson Ultrasonics Corp.,
CT, USA).

2.2 Cement Paste Specimens

Prismatic specimens with dimensions 25 × 25 × 76 mm were prepared per ASTM

C305 [15] using Type I ordinary Portland cement (OPC) and PUCNT aqueous
suspensions, with a water-to-cement ratio of 0.5 (in weight). Plain cement paste
specimens to be used as the benchmark were also cast using Type I OPC and DI
water. All specimens were cast in acrylic molds from which they were removed after
24 h, and then placed in saturated lime water for 28 days.
190 S. Iffat et al.

2.3 Procedure

The dispersibility of the functionalized PUCNTs in DI water solutions was quan-

titatively assessed via dynamic light scattering (DLS) testing using a particle and
molecular size analyzer (model Zetasizer nano ZS, Malvern Panalytical Ltd.).
For each PUCNT concentration, a suspension sample of 1 mL was used to measure
the zeta potential. Each measurement was repeated 10 times to assess the variability
of the results. Representative aqueous suspensions used for the manufacturing of
cement paste specimens with PUCNTs in concentrations of 0.001 and 0.005 wt%
are shown in Fig. 1b.
Uniaxial compression tests were performed on the 28 day cured prism specimens
using a servo-hydraulic loading frame. The load was applied in displacement-control
mode at a rate of 0.3 mm/min and was measured using a 20-kip load cell. Thin
polytetrafluoroethylene inserts were placed between the specimen surfaces and the
loading platens to minimize friction. A representative test setup is shown in Fig. 2a.
Four specimens were tested for each PUCNT concentration (0, 0.001 and 0.005 wt%).
The fractured specimens were used for scanning electron microscopy (SEM) anal-
ysis to investigate the incorporation and dispersion of PUCNTs in the cement paste
structure. Specifically, SEM micrographs were used to visually assess the presence
of undesirable PUCNT agglomerates as a result of ineffective dispersion. To this
end, specimens’ fragments <5 mm in length in any direction were used. Following
the compression tests, these samples were extracted from the prisms, air-dried for
24 h, and placed in a vacuum-suction chamber for 1 h to remove the excess moisture.
The samples were gold-spattered before being examined under a SEM (Ultraplus
Thermal Field Emission Scanning Electron Microscope, Zeiss).

Fig. 2 Uniaxial compression testing of cement paste prism specimens: a test setup, b summary of
results (histogram bars and error bars indicate mean and standard deviation)
Partially-Unzipped Carbon Nanotubes as Low-Concentration … 191

3 Results and Discussion

3.1 Dispersibility in Aqueous Suspension

The effective dispersion of PUCNTs in the aqueous suspensions was assessed by

zeta potential measurements made as part of the DLS tests. The zeta potential values
obtained were –39.5 ± 1.30 mV and –39.6 ± 1.07 mV for the suspensions used to
manufacture cement paste specimens with 0.001 wt% and 0.005 wt% of PUCNT
concentration, respectively. Because suspensions characterized by a zeta potential
smaller than –30 mV are considered stable [16], and also based on visual inspection
(Fig. 1b), it was concluded that the oxidized PUCNTs were effectively dispersed in
the aqueous solutions used in this research.

3.2 Compressive Strength

The mean and standard deviation of the 28 day compressive strength were 32.0 ±
3.10 MPa, 34.7 ± 1 0.93 MPa, and 41.4 ± 7.4 MPa for cement paste with PUCNT
concentrations of 0, 0.001 and 0.005 wt%, as summarized in Fig. 2b.

Fig. 3 Scanning electron microscopy image of plain cement paste structure

192 S. Iffat et al.

Our results showed that the incorporation of very small amounts of PUCNTs
resulted in significant compressive strength enhancement, namely 29% on average
for a concentration of 0.005 wt%, compared with plain cement paste.
Evidence collected through SEM image analysis was used to better understand
plausible strengthening mechanisms. In fact, the micrographs shown in Figs. 3, 4
and 5 suggest that the incorporation of PUCNTs resulted in preferential formation of
amorphous (C–S–H) and crystalline (calcium hydroxide) cement hydrates (Figs. 4
and 5), with a less porous structure than plain cement paste (Fig. 3).
In addition, the lack of visible clusters of PUCNTs suggested that the methodology
used will produce cement composites with well-dispersed and chemically-affine
PUCNTs, and thus with a meso-scale homogeneity comparable to or better than that
of plain cement paste. However, the significantly larger standard deviation of the
compressive strength results for the 0.005 wt% concentration suggested that larger
concentrations of PUCNTs may be more difficult to disperse in the cement matrix.

Fig. 4 Scanning electron microscopy image of 0.001 wt% PUCNT-amended cement paste
structure. PUCNTs, partially-unzipped multiwalled carbon nanotubes
Partially-Unzipped Carbon Nanotubes as Low-Concentration … 193

Fig. 5 Scanning electron microscopy image of 0.005 wt% PUCNT-amended cement paste
structure. PUCNTs, partially-unzipped multiwalled carbon nanotubes

4 Conclusions

The following conclusions are presented.

• It is feasible to incorporate well-dispersed and chemically-affine PUCNTs to
enhance the physicomechanical properties of cement paste. The unzipping process
is similar in chemical approach (and thus cost) to the well-known Hummer’s
method used to generate graphite oxide.
• Due to the morphology and high graphene-edge content of PUCNTs, dispersibility
in aqueous solutions was easily achieved.
• A significant increase in the compressive strength of cement paste occurred with
a PUCNT concentration of 0.005 wt%, which is one order of magnitude smaller
than lower-bound concentrations reported in the archival literature for MWCNT-
and GNP-amended cement pastes.
• Microscopic inspection of the cement paste nano- and micro-structures suggests
that plausible contributing mechanisms for strength enhancement include the
preferential and well-distributed formation of cement hydrates in the vicinity of
PUCNTs—also resulting in a less porous structure—compared with plain cement
paste.
194 S. Iffat et al.

Acknowledgements This material is based on collaborative work supported by the U.S. Depart-
ment of Energy Office of Science, Office of Basic Energy Sciences, and Office of Biological and
Environmental Research, under award number DE-SC0012530; Savannah River National Labo-
ratory; and, the University of South Carolina (USC) Advanced Support for Innovative Research
Excellence (ASPIRE) program. Special thanks are extended to Ms. Erika Rengifo (Ph.D.) and Ms.
Sarah Riser (undergraduate research assistant) at the UofSC Department of Civil and Environmental
Engineering, and personnel of the Center for Environmental Nanoscience and Risk at the Arnold
School of Public Health, and the Electron Microscopy Center, for their technical assistance.

Author Contributions The authors confirm contributions as follows: study conception and design:
Iffat S., Matta F., Gaillard J., Baalousha M.; data collection: Iffat S., Levington M., Tinkey S., Meany
J.; analysis and interpretation of results: Iffat S., Matta F., Gaillard J., Sikder M., Baalousha M.;
draft manuscript preparation: Iffat S., Matta F. All authors reviewed the results and approved the
final version of the manuscript.

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2. Cheng Y, Zhang S, Li J, Sun J, Wang J, Qin C, Dai L (2017) Preparation of functionalized
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nanotube reinforcements on the mechanical properties of cementitious nanocomposites. Constr
Build Mater 35:647–655
4. Lv SH, Deng LJ, Yang WQ, Zhou QF, Cui YY (2016) Fabrication of polycarboxylate/graphene
oxide nanosheet composites by copolymerization for reinforcing and toughening cement
composites. Cement Concr Compos 66:1–9
5. Zohhadi N (2014) Functionalized graphitic nano-reinforcement for cement composites. PhD
thesis, University of South Carolina, Columbia
6. Kaur R, Kothiyal NC (2019) Comparative effects of sterically stabilized functionalized carbon
nanotubes and graphene oxide as reinforcing agent on physico-mechanical properties and
electrical resistivity of cement nanocomposites. Constr Build Mater 202:121–138
7. Makar JM, Chan GW (2009) Growth of cement hydration products on single-walled carbon
nanotubes. J Am Ceram Soc 92(6):1303–1310
8. Li GY, Wang PM, Zhao X (2005) Mechanical behavior and microstructure of cement compos-
ites incorporating surface-treated multi-walled carbon nanotubes. Carbon 43:1239–1245
9. Nochaiya T, Chaipanich A (2011) Behavior of multi-walled carbon nanotubes on the porosity
and microstructure of cement-based materials. Appl Surf Sci 257:1941–1945
10. Hu Y, Luo D, Li P, Li Q, Sun G, Key S (2014) Fracture toughness enhancement of cement
paste with multi-walled carbon nanotubes. Constr Build Mater 70:332–338
11. Liu J, Wang L, Li Q, Xu S (2019) Reinforcing mechanism of graphene and graphene oxide
sheets on cement-based materials. J Mater Civ Eng 31(4):04019014
12. Jeong YC, Lee K, Kim T, Kim JH, Park J, Cho YS, Yang SJ, Park CR (2016) Partially unzipped
carbon nanotubes for high-rate and stable lithium–sulfur batteries. J Mater Chem A 4:819–826
13. Shende RS, Ramaprabhu S (2016) Thermo-optical properties of partially unzipped multiwalled
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15. ASTM International (2014) Standard practice for mechanical mixing of hydraulic cement pastes
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Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Effect of Fine Aggregates and Test
Settings on the Self-sensing Response
of Cement-Based Composites
with Carbon Nanotubes as Conductive
Filler

T. C. dos Santos, P. A. Carísio, A. P. S. Martins, M. D. M. Paiva,

F. M. P. Gomes, O. A. M. Reales, and R. D. Toledo Filho

Abstract Cement-based self-sensing composites with carbon nanotubes (CNT)

have attracted attention due to their multifunctional properties and great potential
for their application in the smart monitoring of concrete structures. In this study, the
self-sensing properties of one paste and three mortars containing 0.50 and 0.75 wt%
of CNT, and 1.5 and 1.0 sand/cement ratio were investigated, aiming to evaluate
their impact on the piezoresistive response of the composites. The inclusion of sand
in the cement paste with CNT led to a reduced gauge factor and a higher electrical
noise response. The inert aggregates modified the compressive loading mechanical
response of the composites and possibly acted as barriers to electronic mobility,
by increasing the CNT conductive paths’ tortuosity or even interrupting them. The
mortar containing 0.50% of CNT showed a higher electrical resistivity and, at the
same time, greater sensitivity and a more linear self-sensing response than the one
with 0.75% CNT, which can be explained by the CNT content being closer to its
percolation threshold in the first. In this way, a lower CNT concentration generated
a conductive network with a higher capacity to be rearranged under loading, gener-
ating significant changes in resistivity, but a higher CNT concentration presented a
more stable and conductive network. The results suggested that both the conduc-
tive and non-conductive phases affect the detection performance of the composites
and, therefore, must be dosed appropriately. Additionally, the test setup modifica-
tions positively affected the self-sensing response signal, which is particularly useful
to reduce the deleterious effects of the sand additions in the matrix. This overall
approach can make the use of self-sensing mortars in structural monitoring a viable
option.

T. C. dos Santos (B) · P. A. Carísio · A. P. S. Martins · M. D. M. Paiva · O. A. M. Reales ·

R. D. Toledo Filho
Civil Engineering Program, Universidade Federal do Rio de Janeiro, Rio de Janeiro, Brazil
e-mail: thais.santos@numats.coc.ufrj.br
F. M. P. Gomes
Research Center, Eletrobras Furnas, Rio de Janeiro, Brazil

© The Author(s) 2023 197

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_21
198 T. C. dos Santos et al.

Keywords Carbon nanotubes · Self-sensing mortars · Structural monitoring

1 Introduction

Self-sensing cementitious composites change their electrical properties in response

to mechanical loadings and, therefore, are becoming a novel and useful class of mate-
rial for structural health monitoring when compared with conventional technologies
[1–3]. An essential element in these composites is the functional filler that forms an
interconnected conductive network across the cement matrix. In this context, carbon
nanotubes (CNT) are particularly promising when compared with other conductive
fillers, given their excellent electrical, physical, and mechanical properties [4]. Once
incorporated into the composites, in an adequate and well-dispersed concentration,
it is possible to obtain nanostructured cementitious materials, electrically sensitive
to stress–strain that can ultimately be applied as excellent piezoresistive sensors for
monitoring structures.
Yoo et al. [5] investigated the performance of cement pastes containing different
carbon-based fillers, and CNT were found to be the most effective filler for increasing
conductivity when compared with other particles at the same concentration. Han et al.
reported the CNT/cement-based materials’ ability to detect both stress–strain in the
elastic regime, as well as structural damages based on irreversible change in their
electrical resistance responses [6]. Nalon et al. verified the ability of mortars to detect
fire damage and exhibit a residual piezoresistive response, even after their exposure
to high temperatures [7]. Le et al. [8] produced ultrahigh-performance concretes
with improved self-sensing properties by combining micro- and nanoscale fibers and
functional aggregates. In that case, CNT improved the conductive network at the nano
level and, consequently, the conductivity and piezoresistivity of the corresponding
composites.
However, the effective application of this material in structural monitoring requires
more investigation, which motivated the present study. Several authors report that
aggregates, especially the coarse, create obstacles to electrical current flow, extending
or interrupting the conductive paths and negatively affecting the self-sensing response
[1, 2, 9]. Tian et al. reported that from the paste to the concrete, both the number
of publications and the self-sensing performance decreased [3]. On the other hand,
the claims regarding the use of aggregates are not unanimous in the literature. Addi-
tionally, other researchers suggest that aggregates do not significantly affect the
percolation threshold of CNT composites and that their proper incorporation can
reduce the matrix voids and promote more contact points between the fillers [4, 10].
Thus, the electromechanical performance of the CNT composites can be improved.
The effects of aggregates on the properties of self-sensing composites are complex
and therefore require further investigation for better understanding.
Both Han et al. [9] and Wang and Aslani [4] point out that CNT composites with
aggregates are much more profitable in real applications than those made of cement
paste. According to Dong et al., both the conductive and non-conductive phases are
Effect of Fine Aggregates and Test Settings on the Self-sensing … 199

key factors to obtaining self-sensing composites with higher performance [2]. Han
et al. reinforced that the proper selection of constituents and the determination of their
proportions are crucial for designing self-sensing composites [1]. Han et al. consider
that one of the challenges for the development and future implementation of self-
sensing composites is their fabrication and, therefore, suggest that upcoming research
should include investigations related to the design, optimization, and production of
this type of composite, especially those containing aggregates [9].
In this context, aiming to contribute to the smart monitoring of concrete struc-
tures, we present the self-sensing properties of paste and mortars containing CNT
and evaluate the effects of both conductive filler content and fine aggregates. In addi-
tion, the self-sensing test settings and their influence on the quality of the acquired
electrical signals were investigated.

2 Methods

2.1 Materials and Sample Preparation

The materials used for this work were high initial strength Portland cement
(LafargeHolcim), NC7000 Multi-Walled Carbon Nanotubes (MWCNT; Nanocyl),
MasterRheobuild 1000 naphthalene sulfonate-based superplasticizer (BASF Chem-
icals), crystalline silica flour (S325; Mineração Jundu), and MasterMatrix UW 410
cellulose-based viscosity-modifying agent (VMA; BASF Chemicals). As the fine
aggregate, a mixture of fine (#16) and coarse (#100) quartz sands was used in the
volumetric fractions of 61.53% and 38.47% respectively.
The MWCNT used in this study consisted of a powder formed by CNT agglomer-
ates, with an average diameter of 9.5 nm, an average length of 1.5 μm, and a surface
area between 250 and 300 m2 /g, according to the manufacturer. They were dispersed
in deionized water using ultrasonication (Sonics Vibra cell, VCX 500) and super-
plasticizer as the dispersing agent. In the dispersion procedure, 30 g of suspension
with a 2 wt% CNT concentration and 2 wt% of superplasticizer solids was subjected
to 20 sonication cycles with an average energy of 110 J/g at an amplitude of 40%
and pulses of 20 s. In each cycle, the suspension of CNT was homogenized in a
low-temperature bath to cool the mixture. The UV–Vis absorbance spectra of the
CNT/superplasticizer aqueous solutions were measured before and after the disper-
sion procedure. For the measurements, an aliquot of the suspensions was diluted in
deionized water in the ratio of 1:2000 (% v/v).
To investigate the influence of sand addition on the self-sensing properties of the
composites, samples containing 0.75% of CNT were produced, without and with sand
(P1 and M1, respectively) at a 1.5 sand/cement (or s/c) ratio, as shown in Table 1. In
addition, to evaluate the effect of the CNT content on the mortars’ piezoresistivity, a
sample of the same mix design was produced incorporating 0.5% of CNT dispersed
by mass of cement (M2). In the second part of the study, to investigate the effects
200 T. C. dos Santos et al.

of the test configurations and the sand content, the M3 trait was defined, with an s/c
ratio of 1.0.
The P1, M1, and M2 samples were produced as cubic specimens with 50 mm
edge and 4 copper electrodes with dimensions of 0.3 × 30 × 50 mm (thickness x
width x length) embedded equidistantly (Fig. 1). The M3 sample was cubic with
40 mm edge and copper electrodes with 1050 mm2 (30 × 35 mm) of embedded
area. All mixtures were mixed manually, using a glass rod, into which the previously
homogenized fine solids (cement, mineral admixture, and VMA) were incorporated
into the aqueous dispersion of CNT and additional water and mixed for 10 min. The
sand, when included, was incorporated into the mixture after the previous process and
homogenized for another 3 min. Finally, compaction was performed by vibration,
before and after the copper electrode insertions. After demolding, the samples were
cured in a humid chamber at room temperature.

Table 1 Mix designs of the self-sensing cementitious composites

Sample w/c S325 VMA s/c CNT Mix proportions (kg/m3 )
(%a ) (%a ) (%a ) Cement S325 Sand Water CNT
P1 0.5 13.0 0.1 0.0 0.75 1139.42 148.12 0 569.71 5.70
M1 1.5 0.75 669.25 87.01 381.52 334.63 5.02
M2 1.5 0.50 670.00 87.11 381.95 335.00 3.35
M3 1.0 0.50 776.98 101.04 776.98 388.49 3.88
a By mass of cement. CNT, carbon nanotubes; VMA, viscosity-modifying agent

Fig. 1 Copper electrodes (a) and views of b mold setup and c a typical produced sample
Effect of Fine Aggregates and Test Settings on the Self-sensing … 201

2.2 Self-sensing Test

Self-sensing tests were executed by applying cyclic compressive loadings with simul-
taneous acquisition of electrical signals using the embedded copper plates as elec-
trodes. Before the self-sensing tests, all samples were dried at 60 °C for 3 days
to eliminate the effect of humidity. In addition, the specimens were instrumented
with strain gauges and electrically isolated from the machine using insulating tape.
The electrical resistance of the matrices was evaluated using the four-probe method,
in which the DC voltage is applied to the samples through the external electrodes,
and the voltage measured in the sample is recorded by a data acquisition system
connected to the internal electrodes. The test setup is shown in Fig. 2.
Prior to the mechanical loading, DC voltage was applied to each sample for 20 min,
to achieve electrical signal stabilization and to mitigate the sensor capacitive effects.
During the test, the electric current of the circuit was acquired by using a shunt
resistor with a known electrical resistance connected in series. Using the voltage
of the shunt recorded throughout the test and applying Ohm’s 1st law (Eq. 1), the
electrical current was obtained. With the current and voltage between the internal
electrodes, the electrical resistance of the matrix was determined.
For the correlation with the strains, the fractional change in resistance (FCR) in
response to the applied loads were obtained according to Eq. 2. The initial electrical
resistance of the sample R0 was measured after a pre-load of 0.5 kN at the end of the
electrification time of 20 min. On the other hand, electrical resistance R corresponds
to the electrical resistance value over time, under a given compressive loading cycle
in the linear elastic regimen of material. Finally, the electrical resistivity ρ of the

Fig. 2 Experimental setup for the self-sensing test

202 T. C. dos Santos et al.

samples, in Ω.cm, was obtained through Ohm’s 2nd law (Eq. 3), considering the
distance between the electrodes and their contact area with the cement composite.

V = R×i (1)

R − R0
FC R = (2)
R0
R0 × A
ρ= (3)
L
where: V is the voltage (V ); R is the electrical resistance (Ω); i is the intensity of
the DC current (A); R0 is the initial electrical resistance of the sample, in Ω, which
corresponds to the value recorded at the end of the electrification period and before
the loading cycles; L is the distance between the copper electrodes (cm); and A is
the embedded area of the electrodes (cm2 ).
In the first sequence of tests, in which samples P1, M1, and M2 were evaluated,
the described test setup was used, applying an electrical voltage of 5 V and shunt
resistance of 100 Ω. In the second part, using the sample M3 and aiming to investigate
the effects of the test configurations, both the applied DC voltage and shunt resistance
were varied, respectively, from 4 to 12 V and from 6 to 100 kΩ.

3 Results and Discussion

3.1 Materials Characterization

Figure 3 shows the particle size distribution, obtained by laser diffraction analysis,
of the cement, silica flour (S325), and sand fractions. Table 2 lists the chemical
composition of the raw materials and their main physical properties. After CNT
dispersion by sonication, there was a relative increase in the area under the UV–Vis
spectrum and an increase in the absorbance peak, corresponding to the wavelength
range around 260 nm (Fig. 4). Before sonication, the absorbance intensity of 0.16 at
the peak demonstrated a low degree of dispersion of the nanotubes. After ≈66,000 J
of sonication energy, the highest value of 1.31 at the peak suggested effective CNT
dispersion in the water/superplasticizer solution.
Effect of Fine Aggregates and Test Settings on the Self-sensing … 203

Fig. 3 Particle size

distributions of the raw
materials

Table 2 Chemical composition and main physical properties of the raw materials
Chemical composition (wt%) Cement S325 Fine sand Coarse sand CNT
CaO 66.37 – ND ND 0.111
SiO2 14.043 97.140 0.532
Al2 O3 3.681 1.570 7.013
Fe2 O3 4.171 0.175
SO3 3.933 0.813 0.456
K2 O 0.484 – –
TiO2 0.294 – –
SrO 0.275 – –
MnO 0.082 – –
ZnO 0.033 0.011 –
CuO 0.011 0.012 –
ZrO2 – 0.008 –
Co2 O3 – – 0.087
P2 O5 – – 0.441
Tm2 O3 – – 0.003
LOI 6.626 0.445 91.18
Physical properties
Density (g/cm3 ) 3.0676 2.6902 2.50 2.55 1.5
D50 (μm) 12.2 15.5 ND ND ND
Maximum size of aggregate (mm) ND ND* 0.6 2.4 ND
Fineness module ND ND* 0.9 2.9 ND
CNT, carbon nanotubes; ND, not determined
204 T. C. dos Santos et al.

Fig. 4 UV–Vis spectra of

carbon nanotubes in the
water/superplasticizer
solution before and after
sonication process

3.2 Effect of Fine Aggregate and CNT Content on Electrical

and Piezoresistive Properties

The initial resistivity ρ0 of the composites P1, M1, and M2 was 100, 600, and 700
kΩ.cm, respectively. Figure 5 shows the strain and FCR variations as a function of
time for these samples, under cyclic compressive loadings from 2 to 8; and 2 to 12
kN (0.8–3.2 and 4.8 MPa). Figure 6 shows the relationship between strain and FCR
values during the self-sensing tests of samples P1, M1, and M2. No smoothing was
applied to the obtained data.
A piezoresistive response of all cementitious composites can be seen in Fig. 5, once
the FCR values decreased with the increase in both loading and strain, and increased
with the reduction of strain, during the unloading step. During cyclic loading, under
compression, the CNT came closer and conductive paths formed due to the contact
and tunneling conductivities, allowing more electric current flow and reducing the
electrical resistance of the sample. Upon unloading, the conductive network returned
to its initial state and recovered its electrical resistance, increasing the FCR value
[2, 11].
However, it should be noted that there was an increase in the electrical resistivity
ρ0 , as well as a change in the amplitude and quality of the piezoresistive response with
the addition of fine aggregate and the reduction in CNT content. Comparing paste
P1 and mortar M1 (Fig. 5a, b), both with 0.75% of CNT content, it is notable that
the aggregate addition affected the FCR variation. In P1, the load cycles generated
a higher maximum amplitude of FCR, equal to 0.33, while for M1 a less sensitive
response to loading was obtained, with a maximum FCR value of 0.15, besides a
greater degree of noise. Naturally, this also affected the gauge factor (GF) values,
equal to 1076 for P1 and 375 for M1, which is equivalent to the slope of the FCR–
strain curve (Fig. 6) and represents the sensitivity of the composites. In addition, the
highest coefficient of determination (R2 ) was obtained in paste P1 (0.91), meaning
that the FCR–strain curve was less scattered compared with mortar M1—noisier and
with a lower R2 of 0.52.
Effect of Fine Aggregates and Test Settings on the Self-sensing … 205

P1
(a) Paste with 0.75%
of CNT

M1
(b) Mortar with 0.75% of
CNT and s/c = 1.5

M2
(c) Mortar with 0.5% of
CNT and s/c = 1.5

Fig. 5 Fractional change in resistance (FCR), stress, and strain responses as a function of time
of a P1, b M1, and c M2 under cyclic compressive loading. CNT, carbon nanotubes; s/c, sand to
cement ratio

This behavior is possibly explained by the addition of the fine aggregate, which
besides modifying the mechanical response of the composite to loading, disturbed the
conductive paths and electron mobility due its insulating nature. Previous evidence
from the literature suggests that aggregates can constitute obstacles to current flow,
negatively affecting the sensitivity and noise of the self-sensing response [2–4, 9,
12]. Therefore, the results obtained by this study reiterate that both the conductive
206 T. C. dos Santos et al.

Fig. 6 Fractional change in resistance (FCR) versus strain diagrams of samples a P1, b M1, and
c M2 under cyclic compressive loads

(CNT) and non-conductive phases (aggregate) affect the self-sensing performance

of the composite.
Chiarello and Zinno observed that the magnitude of the conductivity of the system
containing conductive fibers decreases exponentially with the increase of the s/c ratio
[13]. García-Macías et al. [14] found variations in the resistivity of paste tenfold
greater than those found for mortar and concrete. Dong et al., in a literature review,
explained that the presence of the aggregate weakens the sensitivity of the composite,
due to the imperfect crack controlling capacity of conductive fibers under the interac-
tion with aggregates [2]. Moreover, due to the tendency of aggregates to separate the
connectivity of the CNT conductive network. Both D’Alessandro et al. [12] and Han
et al. [15] reported other adverse effects with the use of aggregates, such as reduced
sensitivity, repeatability, and signal quality of self-sensing composites containing
CNT.
The detection ability of the composites was also affected by the CNT concentra-
tion, as can be seen when comparing the mortars M1 and M2, with 0.75% and 0.5%
of CNT, respectively. The maximum FCR amplitude obtained for these samples,
corresponding to the compressive stress of 4.8 MPa, was equal to 0.15 and 0.29
respectively, suggesting a higher sensitivity of the composite containing 0.5% of
CNT. It can be seen that, compared with M1, the self-sensing response of sample
M2 showed higher linearity (with R2 = 0.73) and detection sensitivity (GF 883).
This behavior is probably due to the CNT content of M2 being closer to the perco-
lation threshold, which generated a conductive network with a greater capacity to
rearrange under loading, resulting in more significant changes in electrical resistivity
and greater sensitivity. In the case of samples with 0.75% of CNT, the concentra-
tion of conductive filler was closer to the conductivity zone, in which conductive
networks are more stable and denser, leading to a lower sensitivity [1–4, 9]. Thus,
the self-sensing properties are also dependent on the CNT concentration in the matrix
and its distribution, which is negatively affected by the distribution of the aggregates,
once they are the non-conductive phase.
Effect of Fine Aggregates and Test Settings on the Self-sensing … 207

Fig. 7 Fractional change in resistance (FCR) response under 10 kN cyclic compressive load with
a varying voltage from 4 to 8 V and shunt resistance from 6 to 9 kΩ and b varying voltage from 4
to 12 V and shunt resistance from 14 to 100 kΩ

3.3 Effect of Test Setup on Electrical and Piezoresistive

Properties

Figure 7 shows the variation in FCR for sample M3, under 10 kN of cyclic compres-
sive loading, when both the applied DC voltage and shunt resistance were varied
from 4 to 12 V and from 6 to 100 kΩ, respectively. In Fig. 7a the lower values of
electrical resistance of the shunt resistor led to noisier, intensified, and even distorted
responses. From 14 kΩ of shunt resistance, as seen in Fig. 7b, the amplitude of the
FCR variation over time became approximately equal, regardless of the voltage and
shunt value adopted, suggesting that the response to loading in those cases was solely
dependent on the matrix properties.
Supposedly, lower shunt resistances interfere with their acquired voltages, nega-
tively affecting the current and matrix resistance values derived from these measure-
ments. Higher shunt resistances, compatible with the magnitude of resistance vari-
ations observed in the cement matrix, seemed to cause a better resolution of the
acquired data. Therefore, the present results suggested that the higher noise level
observed was not only a property of the matrix but also dependent on the test config-
uration. The latter is easily adjustable to allow a more efficient acquisition of electrical
signals.
Previous investigations suggest that factors related to the measurement of elec-
trical signals in the self-sensing test can affect the intensity and stability of the piezore-
sistive response—for example, the configuration of the electrodes or the current type
and its magnitude—and therefore should be properly adjusted [2, 10, 16, 17]. Galao
et al. applied varying electrical currents to the self-sensing composites (0.1, 1.0,
and 10 mA) and found that the piezoresistive response was better with increasing
current intensity, achieving better signal stability and correlation with the strain of
208 T. C. dos Santos et al.

material [17]. On the other hand, among the applied electrical voltages of 10, 20,
and 30 V, Konsta-Gdoutos and Aza obtained an optimal voltage of 20 V, suggesting
that high electrical current intensities can also be harmful [16]. Ding et al. evaluated
self-sensing cementitious composites using reference resistors with 1000 Ω of elec-
trical resistance [18], and other researchers used the same circuit model but without
mentioning the value of shunt resistance [19].
Based on these results, our setup was defined with a shunt resistor of 27 kΩ and
a DC voltage of 4 V to evaluate mortar M3. Figure 8a shows the results under cyclic
loading of 3.2 MPa for these conditions. Comparing an excerpt of the M2 test using
the first setup (Fig. 8b), a clear change in the self-sensing response can be seen,
with considerable noise reduction of the signal. The better self-sensing response
suggested that not only the concentrations of conductive filler and fine aggregate
affect the electrical signals acquired in self-sensing tests. The improved response
can be explained in part by the decrease in the s/c ratio but seems to be mainly due
to the change in the test settings.
The R2 coefficient observed in Fig. 9b, equal to 0.96, was even higher than that
obtained for the paste P1, equal to 0.91, which had more CNT and no aggregate. The
mortar M3 also showed good detection sensitivity, represented by GF of 693.
Yoo et al. [5] investigated the self-sensing performance of cement pastes and
verified that the sample containing 1.0% by volume of CNT presented a maximum
amplitude of the FCR equal to 0.26 under a compressive load of 40 kN. The GF of
the composites ranged from 77.2 to 95.5 with a minimum R2 of 0.9382. Yin et al.
[11] produced cement pastes with 1.7% CNT, by volume, with a high coefficient of
sensitivity to deformation (1500), high linearity (R2 = 0.97), and maximum FCR
equal to 0.19 under cyclic compressive loading of 10 MPa. The hybrid combination of

Fig. 8 Fractional change in resistance (FCR), stress and strain responses as a function of time of
samples a M3 and b M2 under cyclic compressive loading
Effect of Fine Aggregates and Test Settings on the Self-sensing … 209

Fig. 9 Fractional change in resistance (FCR) versus strain diagrams of a M3 and b M2 mortars
under cyclic compressive loading

CNT and nickel nanofibers resulted in the best piezoresistive sensitivity and response
linearity under the same load, with a maximum FCR equal to 0.24, GF of 1880, and
R2 of 0.99.

4 Conclusions

In this study, the self-sensing properties of cementitious composites containing fine

aggregate and varied content of CNT were evaluated under compressive cyclic
loading. In the first step, the paste P1 with 0.75% CNT led to higher linearity and
sensitivity of the self-sensing response and a low degree of noise in the electrical
signal, when compared with mortars M1 and M2. The electrically inert aggregate
possibly modified the mechanical response of the composite to compressive load and
constituted an obstacle to electron flow, increasing tortuosity or even interrupting the
conductive paths of the CNT. Comparing the mortars M1 and M2, with 0.75% and
0.5% of CNT respectively, higher detection sensitivity and linearity were verified
for the composite with lower CNT content—closer to the percolation threshold. At
this CNT concentration, a conductive network formed with a greater capacity to rear-
range under loading, resulting in more significant changes in electrical resistivity. On
the other hand, higher concentrations of CNT, especially closer to the conductivity
zone, formed denser and more stable conductive networks, leading to lower detection
sensitivities. These results confirmed that both the conductive and non-conductive
phases affect the electrical and self-sensing behavior of cementitious composites
containing CNT. The conductive network is, therefore, formed by an appropriate filler
concentration, and the choice of aggregate, as well as the appropriate proportions,
210 T. C. dos Santos et al.

is extremely important for the fabrication of self-sensing cementitious composites

with good detection sensitivity and signal stability.
Additionally, the test setup configuration should be also taken into account when
investigating the self-sensing performance of cementitious composites. We verified
that the test settings can allow more efficient acquisition of electrical signals and
even reduce the noise caused by the addition of aggregates. In a circuit with a refer-
ence resistor, lower values of shunt resistance led to noisy, scattered, and distorted
electrical responses, whereas shunt resistances from 14 to 100 kΩ led to a better
self-sensing response signal, with reduced noise and higher linearity. Under these
conditions, the FCR under loading was always less scattered and with the same
amplitude, reinforcing its capability of reflecting the piezoresistive properties of the
matrix. This suggests that strategic adjustments in the self-sensing test settings may
be particularly useful to reduce the deleterious effects of added sand in the matrix,
enabling its use as aggregate in self-sensing composites while maintaining satisfac-
tory detection performance. This approach may benefit the implementation of smart
mortars in structural health monitoring. In addition, it can contribute to the massive
use of aggregates in this type of composite, consequently lowering the consump-
tion of raw materials with higher cost and environmental impacts such as CNT and
Portland cement.

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Effect of Carbonation
on the Microstructure and Phase
Development of High-Slag Binders

M. J. Tapas, A. Yan, P. Thomas, C. Holt, and V. Sirivivatnanon

Abstract The drive for sustainable concrete production favors the use of high
replacement levels of supplementary cementitious materials (SCMs) in the concrete
mix. The use of SCMs such as fly ash and slag, however, although they improve
the sustainability of concrete production as well as most concrete durability prop-
erties, increases the carbonation rate. Carbonation decreases the pH of the concrete
pore solution, making the steel reinforcement susceptible to corrosion. The effect of
carbonation is, however, not confined to the change in pH of the pore solution. We
investigated changes in the microstructure and phases of high-slag binders due to
carbonation. The carbonation resistance of mortars with 50 and 70% slag replacement
were investigated at exposure conditions of 2%CO2 , 50%RH, 23 °C. The carbonated
and non-carbonated parts of the mortars were subjected to various characterization
techniques to investigate the effect of carbonation on microstructure and phase devel-
opment. Results confirmed the absence of portlandite in all the carbonated regions
(“colorless” by phenolphthalein test, which indicated that the change in color of the
phenolphthalein solution was due to the absence of portlandite to buffer the pH).
Significant reduction in the amount of C-S-H, as well as increase in the amount
of calcium carbonate, were been observed in the carbonated regions. Aragonite, a
polymorph of CaCO3 , was very prominent in all the carbonated mortars.

Keywords Carbonation · Microstructure · Slag

M. J. Tapas (B) · A. Yan · P. Thomas · V. Sirivivatnanon

School of Civil and Environmental Engineering, University of Technology Sydney,
Ultimo, NSW, Australia
e-mail: mariejoshua.tapas@uts.edu.au
C. Holt
Boral Innovation, Maldon, NSW, Australia

© The Author(s) 2023 213

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_22
214 M. J. Tapas et al.

1 Introduction

Concrete, the most utilized construction material and the second most used substance
on earth after water, creates significant amount of CO2 emissions during produc-
tion [1]. The CO2 emissions primarily originate from the calcination of limestone
(CaCO3 ) to produce cement (main binder material in concrete), releasing CO2 in
the process. As CO2 contributes to global warming, there has been a strong focus
on the decarbonation of concrete, and reducing the amount of cement per cubic
meter of concrete is a proven approach to deliver on the target CO2 reduction.
Reducing the cement content can be achieved through partial cement substitution with
supplementary cementitious materials (SCMs) such as slag or fly ash. Slag, being
hydraulic in nature like cement, can be used at higher substitution rates than other
SCMs (usually ≥ 50%), translating to better CO2 reduction. However, although slag
notably improves the later-age strength as well as most concrete durability proper-
ties, including chloride ingress [2], the alkali–silica reaction [3–5], sulfate resistance
[6, 7], and delayed ettringite formation [8], its use results in increased susceptibility
of the concrete to carbonation [9–11].
Carbonation refers to the ingress of CO2 into the binder system, which results in the
formation of carbonic acid (H2 CO3 ) that further dissociates into H+ and CO3 2− and
reacts with calcium ions in the pore solution, resulting in the precipitation of calcium
carbonate (CaCO3 ) and a decrease in the pH of the pore solution. Low concrete
pH (≤9.5) resulting from carbonation is detrimental to steel-reinforced concrete
because steel begins to lose its passivation layer at low pH, making it susceptible
to corrosion [10]. Thus, carbonation is a serious durability concern, particularly for
steel-reinforced concrete [12].
Phenolphthalein is an indicator used to assess the depth of carbonation. It is
colorless at lower pH values (≤9), whereas at pH > 10.5, it presents a characteristic
purple or magenta. Because mortar or concrete that has been carbonated has pH ≤ 9,
phenolphthalein is used to visually confirm the drop in pH due to carbonation [13].
The effect of carbonation is, however, not confined to the change in pH of the pore
solution. Although it has been reported that carbonation in general is beneficial and
results in an increase in compressive strength due to the conversion of Ca(OH)2 to
CaCO3 , which increases the volume of the binder and reduces porosity [10], it appears
that this may not be true for all binder systems. It has been reported that although
moderate carbonation can improve the mechanical properties of the concrete, exces-
sive carbonation impairs mechanical strength due to the decalcification of the C-S-H
[14]. High-slag concrete is also particularly susceptible to carbonation shrinkage [9].
Therefore, due to variability in the reported effect of carbonation and considering the
increasing levels of slag being used in concrete production, a better understanding
of the effect of carbonation on microstructure and phase development is required.
We investigated the effect of carbonation on the microstructure and phase devel-
opment of high-slag mortars (50% and 70% slag replacement). The mortars were
characterized after being subjected to accelerated carbonation conditions (2%CO2 ,
23 °C, and 50% relative humidity (RH) for 112 days.
Effect of Carbonation on the Microstructure and Phase Development … 215

Table 1 Mortar mixes

Mix details Normen sand (g) Cement (g) Slag (g) Total binder (g) Water (g)
OPC 1350 450.0 0.0 450.0 171.0
OPC + 50% slag 1350 225.0 225.0 450.0 171.0
OPC + 70% slag 1350 135.0 315.0 450.0 171.0
OPC, ordinary Portland cement

2 Methods

2.1 Raw Materials

We used General Purpose cement and slag that complied with AS3972 and AS3582.2
respectively. Normen sand (CEN Standard Normsand according to EN196-1) was
used as fine aggregate. The maximum moisture content of sand was 0.2%.

2.2 Carbonation Test

The 40 × 40 × 160 mm mortars were prepared using Normen sand in combination

with ordinary Portland cement (OPC) and OPC + slag binders (slag at 50% and
70% replacement) at 0.45 water to cement ratio. Table 1 shows the mortar mixes
investigated in this study.
The mortars were demolded after 1 day, cured for 28 days inside a sealed moisture
bag and then transferred into the shrinkage room (50%RH, 23 °C) to air cure for 7 days
in preparation for the accelerated carbonation test. At age 35 days (28 + 7 days), the
mortars were transferred into the carbonation chamber running at 2%CO2 , 50%RH
and 23 °C for the accelerated carbonation test. Carbonation depth measurements
using phenolphthalein (1% solution) were carried out after 1 week (7 days), 4 weeks
(28 days), 9 weeks (63 days), and 16 weeks (112 days) exposure of the mortars in
the carbonation chamber.

2.3 Characterization of the Mortars

After 112 days carbonation, the “colorless” and “pink regions” of the mortars were
subjected to thermogravimetric analysis (TG; SDT-Q600 Simultaneous TGA/DSC
equipment, TA Instruments) and scanning electron microscopy (SEM). The mortar
specimens were ground and 50-mg samples of the ground material were transferred
to a platinum crucible, which was placed inside the TG instrument. The thermal
216 M. J. Tapas et al.

analysis was performed in a nitrogen gas atmosphere, within a temperature range of

23–900 °C and at a heating rate of 10 °C/min.
All imaging and elemental analyses of the fractured mortars were performed using
a Zeiss Supra 55VP SEM fitted with a Bruker SDD EDS Quantax 400 system and FEI
Quanta 200 with Bruker XFlash 4030 EDS detector. The microscopes were operated
at 15 kV accelerating voltage and 12.5 mm working distance.

3 Results and Discussion

Figure 1 shows the carbonation depths at 16 weeks (112 days). As expected, the plain
OPC mortar showed the best resistance to carbonation (largest pink region) while
mortars with slag exhibited poorer resistance. The higher the slag replacement, the
poorer the carbonation performance.
Figure 2 shows the increase in carbonation depth over time, with the mortar with
70% slag consistently having the highest carbonation depth and fully carbonated at
63 days.
Figure 3 shows the TG curves of the non-carbonated (pink) and carbonated parts
(colorless) of the plain OPC mortar, mortar with 50% slag and the mortar with 70%
slag. Mass loss in the range of calcium hydroxide (CH) decomposition at ≈400–
500 °C corresponds to the dehydroxylation of CH, (Ca(OH)2 → CaO + H2 O) [15].
Therefore, the area under the curve at ≈400–500 °C corresponds to the amount of
CH, and thus a larger area means more CH. Comparing the CH content of plain OPC
mortar and the 50% slag mortar (pink regions), OPC notably has more CH, as may
be expected, which explains its better carbonation resistance. During carbonation,
portlandite, which is the most soluble source of calcium in the binder, serves as a
buffer and maintains the pH of the pore solution by dissolving and releasing OH− and
Ca2+ ions. The OH− neutralizes the H+ while Ca2+ binds CO3 2− , precipitating CaCO3
[11]. Therefore, due to the lower amount of portlandite, high-slag binders carbonate
much faster (i.e. pH drops faster) than pure cement. Absence of portlandite in the

OPC OPC+50%SL OPC+70%SL

Fig. 1 Photos of the mortars (OPC, OPC + 50% slag and OPC + 70% slag) at 112 days exposure
in the carbonation chamber showing the carbonation depth determined using phenolphthalein
Effect of Carbonation on the Microstructure and Phase Development … 217

Fig. 2 Carbonation depth measurements before carbonation (Day 0) and after 7, 28, 63 and 112 days
exposure to 2%CO2 at 23 °C and 50% relative humidity. OPC, ordinary Portland cement; SL, slag

“colorless regions” of the plain OPC and 50% slag mortars are notable indicating the
full consumption of portlandite in the carbonated regions. There is also no portlandite
remaining in the 70% slag mortar, consistent with it being fully carbonated (top and
middle areas of the mortar were tested). Moreover, consequent to the full consumption
of portlandite in the “colorless regions” of all mortars (i.e., fully carbonated regions),
there was a drastic increase in the amount of CaCO3 . A notable decrease in the amount
of C-S-H, carboaluminates, and ettringite (TG region 0–300 °C) can also be seen,
because once portlandite has been fully consumed, the other calcium-bearing phases
start to react with CO2 and carbonate as well [10]. Decalcification of the C-S–H can
occur, resulting in carbonation shrinkage [12].
Figure 4 shows the amount of CaCO3 in the different binder systems (carbon-
ated regions) calculated from the decarbonation region (CaCO3 →CaO + CO2 ). The
higher the amount of CaCO3 formed, the higher the CO2 binding capacity. CO2
binding capacity is related to the amount of CaO in the binder and because the
higher the slag replacement, the lower the CaO available, the CO2 binding capacity
also decreases.
SEM images of the fractured “carbonated” 50% and 70% slag mortars are shown
in Figs. 5 and 6 respectively. The presence of aragonite (a CaCO3 polymorph) is very
prominent in both systems. The microstructure also appears to be porous, although
the change in porosity due to carbonation should be quantified.
218 M. J. Tapas et al.

Fig. 3 Derivative thermogravimetric curves of the carbonated and non-carbonated part of plain
OPC, 50%SL mortar and 70%SL mortar. OPC, ordinary Portland cement; SL, slag

Fig. 4 Percentage CaCO3 in the plain OPC, 50%SL mortar and 70%SL mortar determined from
the thermogravimetric mass loss measurements. OPC, ordinary Portland cement; SL, slag
Effect of Carbonation on the Microstructure and Phase Development … 219

CaCO3

CSH

Fig. 5 Scanning electron microscopy images of the carbonated 50%slag mortar

CaCO3

CSH

Fig. 6 Scanning electron microscopy images of the carbonated 70%slag mortar

4 Conclusions

In this study we investigated the effect of carbonation on the microstructure and

phase development of high-slag binders. Relevant results are as follows.
1. High-slag binder carbonates faster than plain OPC binder. The higher the slag
content, the higher the carbonation rate.
2. TG analysis confirmed the absence of portlandite and increased CaCO3 in the
carbonated regions (“colorless” regions), which suggests that the trigger for
the drop in the pH (i.e., the change in color of phenolphthalein from “pink”
to “colorless”) is the absence of portlandite to buffer the pH.
3. Significant reduction in the amount of C-S-H was also observed, suggesting
decalcification of C-S-H, and that all calcium-bearing phases are prone to
carbonation.
4. The higher the slag replacement, the lower the CO2 binding capacity, which is
consistent with the reduced CaO content of high-slag binders as well as their
higher carbonation rate.
220 M. J. Tapas et al.

5. SEM images confirmed the prominent presence of aragonite (a CaCO3 poly-

morph) in the carbonated mortars.

Acknowledgements This study was carried with the support of the UTS-Boral Centre for
Sustainable Building under funding from the Innovative Manufacturing CRC (IMCRC).

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A New Dispersion Strategy to Achieve
High Performance Graphene-Based
Cement Material

Z. Zhang, Y. Yao, H. Liu, Y. Zhuge, and D. Zhang

Abstract The addition of graphene and its derivatives can enhance the mechanical
and functional properties of cement-based composites, but most of the current tech-
nologies have limited dispersion and are costly. The creation of a cost-effective
graphene-reinforced cement material with uniform graphene dispersion remains
difficult. We used glucose as an economical carbon source to induce the in-situ forma-
tion of graphene on cement particles. Our proposed method is approximately 80% less
expensive than commercial techniques. Evaluation of the microscopic morphology
demonstrated uniform distribution of graphene in the cement matrix, which improved
the mechanical properties of the cement paste. The compressive strengths of the test
groups with 3% carbon source improved by almostly 38% and 48.9%, respectively,
compared with pure cement paste. This newly established technique is essential for
the future design of excellent graphene-based cement materials and the achievement
of multifunctional cementitious applications.

Keywords Cement · Dispersion effect · Graphene · In-situ growth strategy ·

Mechanical properties

Z. Zhang · Y. Yao (B)

School of Civil Engineering, Xi’an University of Architecture and Technology, Xi’an, China
e-mail: yaoy@xauat.edu.cn
H. Liu
School of Chemistry and Chemical Engineering, Xi’an University of Architecture and
Technology, Xi’an, China
Y. Zhuge
UniSA STEM, University of South Australia, Adelaide, SA, Australia
D. Zhang
College of Civil Engineering, Fuzhou University, Fuzhou, China

© The Author(s) 2023 223

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_23
224 Z. Zhang et al.

1 Introduction

Over the past decades, a significant amount of research has been devoted to manipu-
lating the structure of cement hyadration products and the mechanical properties of
cement at the nanoscale by using a wide range of nanomaterials such as nanoscale
silicon dioxide [1], carbon nanotubes (CNTs) [2], and graphene-based materials [3–
7]). Because graphene-based materials are two-dimensional, they possesses good
physical and chemical characteristics, making them a suitable option for the next
generation of improved cement-based material [8–12]. Many studies [13–17] have
demonstrated that graphene and its derivatives, such as graphene oxide (GO), can
effectively improve the mechanical characteristics of cementitious materials by
increasing the hydration process of the cement and altering the pore distribution
in the matrix. Increased durability of GO-reinforced cement mortar can be achieved
with only a small amount of additional GO [18]. However, the high cost and poor
dispersion of graphene-based compounds prevent their future practical application.
The dispersion of graphene materials in the cement matrix [19, 20] is the primary
factor that determines how well graphene can reinforce cementitious materials.
It is possible to attribute the aggregation of graphene and its derivative GO to the
powerful van der Waals force that exists between nanomaterials as well as the linking
effect that Ca2+ and Mg2+ ions have on GO in the cement environment. This is because
van der Waals forces are known to exist between nanomaterials [21–23]. Graphene
and GO in aqueous solution have been prestabilized using a variety of chemical and
physical techniques in order to address the issue of their poor dispersion. Graphene-
modified cement can be manufactured by combining cement with the prestabilized
aqueous solution [14, 23–26]. However, treatment with ultrasonication for extended
periods of time and functionalization with strong acids have adverse impacts on
graphene materials, which might cause flaws in the graphene structure.
Here, we describe a novel and uncomplicated technique for the synthesis of
graphene–cement (GC) composite. This strategy involves the in-situ development
of graphene in the cement matrix (Fig. 1) by carbonization and calcination [27, 28].
In the course of the synthesis procedure, the glucose used as the carbon source is
thoroughly combined with the cement [27, 29]. In order to obtain advanced GC
material, the mixture is heated further at 800 °C for 2 h, during which the glucose
is converted into graphene on the cement particles, which inhibits aggregation and
ensures well-dispersed graphene in the cementitious matrix.
A New Dispersion Strategy to Achieve High Performance … 225

2 Methods

2.1 Materials

Glucose (C6 H12 O6 , 98%; Sigma Aldrich) and GO (Suzhou TANFENG graphene
Tech) acted as the carbon source and reinforcement material, respectively. Ordinary
Portland cement (P.O. 42.5; Jiuqi Building Components) and ethanol (C2 H5 OH, AR.;
Sinopharm Chemical Reagent) were also used in this work.

2.2 Sample Preparation

It is a well known that incorporating a negligible quantity of GO (0.05% by weight)

into cement materials can dramatically improve the mechanical characteristics of
these materials [5, 13, 30–32]. As a result, we carried out a control experiment of
adding 0.05 wt% of GO to 100 g of cement. After completely mixing the GO with
water, the mixture was sonicated for 30 min by ultrasonic equipment at the highest
possible power setting (500 W). Next, the GO solution was included in the cement
mix. The production of the new cement paste infused with GO (GOP) followed the
same procedure as that of the GC material. In terms of the overall weight of binder, the
quantity of GO that was utilized was equal to 0.05 wt%. These techniques provide
assurance that the cementitious materials will be capable of meeting the casting
standards at the location of the construction project. The GO that is sold commercially
has a thickness of roughly 1 nm and a diameter that may reach a maximum of
10 µm. The substance known as GC was created by heating a combination of cement
and glucose (carbon source) as shown in Fig. 1. The ratio of cement to glucose
powder varied as follows: 100:1, 100:3, and 100:6. The cement and glucose were
homogeneously mixed manner at room temperature for 1 h at a rotational speed of
1000 rpm. The mixture was then placed in a furnace to undergo the reaction under a

Fig. 1 Diagram of in-situ graphene grown on cement particles

226 Z. Zhang et al.

Table 1 Mix proportions of

Group Cement Water GO/glucose
different groups
GP-0 100 35 /
GP-1 100 35 1
GP-3 100 35 3
GP-6 100 35 6
GOP 100 35 0.05
* The names of the samples were based on the graphene oxide
(GO) or graphene concentration. For example, GP-0 denotes pure
cement paste. GOP refers to cement paste containing 0.05% GO.
GP-1 refers to the cement paste containing 1% carbon source prior
to graphitization

specified high-temperature program while being shielded by N2 as shown in Fig. 1.

In this particular investigation, the oven was initially purged with nitrogen at a rate
of 50 mL/min before being heated from 25 to 550 °C at a rate of 5 °C/min over
60 min. The temperature inside the furnace then increased by 5 °C every minute
until it reached 800 °C, where it remained for 2 h before gradually decreasing to the
temperature of the surrounding air. The amount of graphene could be changed by
adjusting the bulk amount of glucose, and the yield of pure graphene was determined
by using the same techniques. The percentage of graphene that can be yielded was
7%. In order to investigate the effect that in-situ grown graphene development had
on the mechanical characteristics of cement paste, we divided the samples into 4
groups: GP-0, GP-1, GP-3 and GP-6 according to the addition amount of glucose
before heat treatment as presented in Table 1. Fresh cement paste containing the
in-situ growing graphene was pressed into plastic molds with dimensions of 50 ×
50 × 50 mm (for the compressive test) and 40 × 40 × 160 mm (for the flexure test).
The molds were removed after 24 h and the samples placed in an environment with
a relative humidity of 95% for 3, 7, and 28 days, respectively.

2.3 Experimental Methods

In the current investigation, a porosimeter capable of detecting pores ranging in

size from 5 nm to 100 µm was utilized (equivalent to pressures of 206 MPa and
345 kPa, respectively, which are the highest and minimum that were applied). Raman
spectra were obtained with a Renishaw spectrometer that included an excitation
laser wavelength of 532 nm. The instrument was also fitted with ×50 lens and was
focused to a spot size that was 1 m in diameter. The scanning electron microscopy
(SEM) examination was performed in order to investigate the microstructure. The
compression test of cubic specimens with dimensions of 50 × 50 × 50 mm and the
three-point flexure test of beam members with dimensions of 40 × 40 × 160 mm are
both common methods for determining the mechanical properties of cementitious
A New Dispersion Strategy to Achieve High Performance … 227

materials. Both of these tests were performed on specimens with dimensions of 50 ×

50 × 50 mm. When calculating the amount of compressive stress, the applied force
is divided by the area that is being loaded. The loading rate was set at 1.2 mm/min
for the compressive strength test, and at 0.05 mm/min for the flexure test.

3 Results and Discussion

3.1 Characterization of GC Material

The conversion of glucose (carbon source) into graphene on cement particles, which
took place during the manufacturing process of GC material, was an essential step
in our process of dispersing graphene evenly throughout the cement composites.
Figure 2 illustrates the morphology of graphene as well as GC. The results of tests
using energy dispersive X-ray spectroscopy (EDS) showed a distinct distribution of
carbon (Fig. 2b, c). EDS mapping of the GC material showed that the elements carbon,
oxygen, calcium, and silicon were all equally distributed throughout the material
(Fig. 2e). SEM and atomic force microscopy revealed that the wrinkled nanosheets
of graphene generated by glocuse had a thickness of 1.1 nm (Fig. 2e, f). Very thin
sheets were positioned on the surface of the cement particles and had a wrinkly
appearance that was analogous to the morphology shown in Fig. 2e. Additional
methods of characterization indicated beyond a shadow of a doubt that glucose was
effectively transformed to graphene. The X-ray diffraction (XRD) patterns of the
GC composite (Fig. 3a) revealed a new peak around 27° representing as-formed
graphene sheets. According to the Raman spectra of the GC material, two additional
peaks were discovered at 1578 and 1360 cm−1 , corresponding to the G-peak and D-
peak of graphene, respectively. These peaks are found at these specific frequencies.
The development of graphitic carbon was further supported by the G-band of the
samples’ 532 nm Raman spectra (Fig. 3b), which was located at 1578 cm−1 . The
sample displayed a wide D-band with its center at 1360 cm−1 , which was indicative
of nanoscale graphite particles and chemically modified graphene flakes. The center
of the band was at 1360 cm−1 . This property, representing the existence of disorder
as well as the boundaries of graphene domains, was detected with high-resolution
SEM. The results of the studies suggested that the GC composite was composed
of cement and graphene. Additionally, the results indicated that the graphene was
equally distributed throughout the cement matrix and was confirmed by the fact that
the GC material passed the GC test.
228 Z. Zhang et al.

Fig. 2 Microscopic shape and structure of graphene, cement, and graphene–cement (GC)
composite. a SEM image of fabricated GC composite in this study. b–d EDS of GC material.
e High-resolusion SEM of GC material. f AFM of fabricated graphene sheet. AFM, atomic force
microscopy; EDS, energy dispersive X-ray spectroscopy; SEM, scanning electron microscopy

Fig. 3 a X-ray diffraction results, b Raman results of graphene, cement and graphene–cement
composites
A New Dispersion Strategy to Achieve High Performance … 229

3.2 Mechanical Properties of GC Material

After being cured for 28 days, the compressive strengths of GOP paste, GP-0, GP-1,
GP-3, and GP-6 were evaluated, and the results are depicted in Fig. 4. When calcu-
lating each reported compressive strength, the average of three duplicate specimens
was used as the basis for the calculation. The compressive strengths were affected
by the different content of the reinforcing materials. The compressive and flexural
strengths of the GC paste were superior to those of GP-0 and GOP. The compressive
strength of GC paste increased in direct proportion to the graphene content. The
GP-3 group demonstrated the strongest compressive strength of all of the groups. In
comparison with the GC paste, GOP had a somewhat lower compressive strength.
After curing for 28 days, the compressive of GP-3 rose by 38.18% in comparison
with GP-0. In contrast, after curing for 28 days, the compressive strength of GOP
fell by almost 0.75%, as shown in Fig. 4.

3.3 Dispersion Effect

The large-scale SEM study of GP-3 and GOP, as well as the related element scanning
tests, were carried out as shown in Fig. 5 for the purpose of confirming consistent
distribution of the graphene. The carbon element distribution map of GCP-3 is shown
in Fig. 5b, d, and the element distribution map of GOP is shown in Fig. 5f, both at
the same scale as Fig. 5b. As can be seen in Fig. 5f, the aggregation of GO resulted
in the formation of carbon element facula. On the other hand, graphene with uniform
dispersion does not create an aggregation zone at the same scale (Fig. 5b), which
demonstrated that the graphene was uniformly disseminated throughout the cement
matrix.

Fig. 4 The 28-day compressive strength tests of GP-0, GP-1, GP-3, GP-6, and GOP and
corresponding rising rates of the samples compared with pure cement paste after 28 days
230 Z. Zhang et al.

Fig. 5 Microscopic morphology and structure of GP-3 and GOP samples. a, c SEM image of GP-3
with different magnification. b EDS map of (a). d EDS map of (c). e SEM image of GOP. f EDS
map of (e). GOP

4 Conclusions

By heating a mixture of glucose powder and cement powder, a new in-situ growth
approach has been devised with the goal of successfully dispersing graphene
throughout the cement matrix in a homogeneous manner. In order to manufacture
high-quality graphene in situ, glucose was used as the carbon source because it
reduces the overall cost of the process. This recently developed synthetic technique
is extensible to the rational design of additional cement-based materials, and it has
already been done. The in-situ growing process that was developed may produce a
low-cost product and improve the dispersion effect of graphene sheets in the cement
matrix. This in turn improves the mechanical properties of cement paste and makes
it more amenable for graphene-based reinforced cement composites to be used in
civil engineering.

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Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Accelerated Mortar Bar Test to Assess
the Effect of Alkali Concentration
on the Alkali–Silica Reaction

B. Boyd-Weetman, P. Thomas, P. DeSilva, and V. Sirivivatnanon

Abstract We report the outcomes of a study into the influence of alkali concentration
on expansion induced by the alkali–silica reaction (ASR), a deleterious reaction that
causes cracking and durability loss in concrete structures. We assessed the effect of
alkali concentration on mortar bar expansion using a modified form of AS1141.60.1,
the accelerated mortar bar test (AMBT). Mortar prisms were prepared with a reactive
aggregate and immersed in alkali solutions of varying concentrations (from 0.4 to
1.0 M NaOH) and saturated limewater at 80 °C. Expansion was monitored for 28 days.
The degree of expansion was observed to increase with increasing alkali concentra-
tion and an induction period prior to expansion was observed for the 0.4 M NaOH. No
expansion was observed for mortar bars immersed in the control saturated lime water
bath. Additionally, no expansion was observed for mortars using blended cements
containing fly ash (FA) and ground granulated blast furnace slag, suggesting the
AMBT is a viable technique for demonstrating the efficacy of mitigation strategies.

Keywords Accelerated mortar bar test · Alkali-silica reaction · Durability ·

Mortar

B. Boyd-Weetman (B)
School of Mathematical and Physical Sciences, University of Technology Sydney (UTS), Sydney,
NSW, Australia
e-mail: brendan.boyd-weetman@uts.edu.au
P. Thomas · V. Sirivivatnanon
School of Civil and Environmental Engineering, University of Technology Sydney (UTS),
Sydney, NSW, Australia
P. DeSilva
Faculty of Health Sciences, Australian Catholic University (ACU), Sydney, NSW, Australia

© The Author(s) 2023 233

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_24
234 B. Boyd-Weetman et al.

1 Introduction

The alkali–silica reaction (ASR) is a deleterious reaction affecting the long-term

durability of concrete structures. ASR proceeds in the pore solution of concrete
when alkali hydroxides and reactive forms of silica are present. It can be produced in
any cementitious system provided that the three reaction components are available:
alkali, reactive silica, and water. Reactive silica, such as strained quartz or amorphous
opal, is introduced by aggregates. These silica forms dissolve in the alkaline pore
solution and subsequently precipitate to form an alkali–silica gel, the ASR gel. The
ASR gel, once precipitated in the concrete, generally within an aggregate, exerts
mechanical stress from within the concrete, causing the development of cracking.
To reduce the potential for deleterious ASR, a number of mitigation strategies are
deployed in Australia, including the use of nonreactive aggregate materials, a limit on
the maximum available alkali, and the incorporation of supplementary cementitious
materials (SCMs). In Australia, the concrete alkali limit is 2.8 kg/m3 Na2 Oeq [1] and
by specifying a cement alkali content limit of 0.6% Na2 Oeq [2].
Two standardized test methods have been adopted for use in Australia to determine
an aggregate’s propensity to cause ASR: AS1141.60.1, the AMBT, and AS1141.60.2,
the concrete prism test (CPT). These tests are carried out under accelerated conditions
designed to provide aggressive reaction environments to promote ASR gel develop-
ment [2, 3]. An alkali threshold is the lowest alkali level in concrete where deleterious
expansion is found [3]. Threshold testing has largely been the topic of investigation in
CPT-based studies, for which RILEM has a recommended testing method published
as protocol AAR-3.2 [3]. Although threshold focus has been on CPT-based studies,
the AMBT could prove valuable as a tool for identifying threshold behavior. The
benefit of assessing aggregate reactivity using the AMBT procedure is that it enables
a rapid evaluation of an aggregate’s reactivity within a 21-day timeframe (Table 1)
and these benefits could also apply to threshold investigation should the method show
viability. ASR gels in AMBT specimens have been observed to have compositions
that vary compared with gel produced in concretes, with AMBT-induced gels having
a higher sodium content compared with similar binder concretes, likely due to the
external 1 M sodium hydroxide bath [4]. Regardless of gel compositional discrepan-
cies, the AMBT remains valuable to industry for aggregate reaction screening due
to its rapid timeframe and this warrants additional investigation of the test and the
properties of the ASR gel it produces. The AMBT is also used to assess aggregate
reactivity in the presence of SCM-incorporated binders to identify the level of SCM
required to mitigate ASR for a particular aggregate. SCMs mitigate ASR through a
variety of mechanisms that reduce the available alkali of the pore solution [5, 6].
The focus of this study was the alkali concentration of the bath solution in modi-
fied AMBT experiments and we report expansion tests for reactive aggregates as a
function of the bath solution concentration. In addition, expansion of mortars incor-
porating SCM-blended cements is reported for cements containing fly ash (FA) and
slag (S).
Accelerated Mortar Bar Test to Assess the Effect of Alkali … 235

Table 1 Reactivity classification for the AMBT as defined by AS1141.60.1

Mean mortar bar expansion % AS1141.60.1 aggregate reactivity classification
Duration of specimens in 1 M NaOH
80 °C
10 days 21 days
– E < 0.10a Nonreactive
E< 10a 0.10 ≤ E < 0.30 Slowly reactive
E ≥ 0.10a – Reactive
– 0.30 ≤ Reactive
a Value for natural fine aggregates is 0.15%

2 Methods

The primary aggregate of focus was reactive river sand classified as reactive by
AS1141.60.1. Reactive river sand contains 10.7% moderately strained quartz, 2%
heavily strained quartz and 1.3% fine microcrystalline quartz within fragments of
indurated meta-greywacke/siltstone and acid volcanic rock. Three binder material
combinations were used to prepare the prisms: GP cement, FA and ground granu-
lated blast furnace slag (S). The cement used was a Portland-type GP cement that
met the specified requirements of AS3972, with an alkali content of 0.47% Na2 Oe
determined by X-ray fluorescence analysis. For the FA and S incorporated mixes,
cement replacement percentages of 25% and 65%, respectively, were chosen because
they are representative of the recommended SCM substitution rates for mitigating
ASR by AS HB79 [1]. To mix the mortar, the procedure outlined in AS1141.60.1 was
followed. Gauge studs were placed within the mold prior to mixing and calibrated to
have a gauge length of 250 ± 1 mm. Fine aggregate was prepared in its natural unal-
tered grading by oven drying at 110 °C before cooling for mixing. Potable tap water
was used for mixing. The mortar prisms were cured in three gang molds for 24 h
before demolding and were then immersed in tap water at room temperature prior to
heating to 80 °C for 24 h for equilibration prior to zero day length measurement and
subsequent immersion in respective alkali baths equilibrated at 80 °C.
To assess the effect of different external alkali environments and to observe poten-
tial threshold behavior, four alkali concentrations were used as immersion baths for
the mortar bars: 0.4, 0.7 and 1 NaOH and a bath of saturated Ca(OH)2 solution, which
was used as the control bath with no external source of alkali. Distilled water was
used to prepare the immersion solutions. The baths were kept at 80 °C throughout the
duration of the test. Mortar bars were vertically oriented within the bath, supported
by a stainless steel grid so that no contact with the gauge pins occurred.
To determine the comparative length change of the specimens, all comparative
expansion measurements were conducted on a steel frame comparator equipped
with a Mitutoyo digital micrometer. All expansion measurements are in reference
to a 295-mm invar reference bar that was placed with identical positioning for each
236 B. Boyd-Weetman et al.

measurement and checked on a regular basis between mortar bar measurements. The
mortar bars’ comparative length measurements were recorded at day 0 (immedi-
ately after removal from 80 °C water bath), then at 1, 3, 7, 10, 14, 21 and 28 days
following immersion in the alkali baths. To measure relative expansion, mortar bars
were removed from the alkali baths, placed in the comparator for recording to a preci-
sion of 1 micron. These comparative length measurements were carried out within
10 s of removal from the bath and were measured in the same orientation within the
comparator at each age.

3 Results and Discussion

An overview of the GP cement mortar bars (no SCM) is shown in Figs. 1 and 2
displays the expansion curve plots for each binder composition over time while
immersed in the respective alkali immersion baths listed. The concentrations of the
solutions in each bath were measured by titration and the pH calculated is listed in
Table 2, which also lists the expansion for each GP cement mortar bar (no SCM).
Expansion is a strong function of the alkali content of the solution concentration.
Little or no expansion was observed over the timeframe for mortar bars exposed to
the saturated Ca(OH)2 solution. Expansion for the alkali solutions increased with
increasing bath concentration over the time frame of the experiment. An induction
period was apparent for 0.4 M NaOH where expansion was negligible up to 14 days
followed by a notable increase in expansion. Expansion appeared to be increasing at
28 days. Further measurements will yield a limit for the expansion and discriminate
the concentration effects on ASR. Threshold definitions have yet to be applied to
AMBT-based studies. For standard reactivity assessment of the AMBT, if expansion
is equal to, or exceeds 0.1% at 10 days (0.15% for natural sands such as the river
sand used in the present study) or 0.3% at 21 days in bars exposed to 1 M NaOH,
then the aggregate is classified as reactive (Table 1). For the GP mix immersed in
1 M and 0.7 M NaOH we observed the mortar bars exceed this expansion limit,
indicating a classification of reactive. For the mortar bars exposed to 0.4 M NaOH,
the expansion was within the limit, with an expansion of 0.230% observed at 21 days.
Although the expansion was below the expansion limit designated for classification
as reactive, the test was nonstandard and some expansion due to the reactivity of the
aggregate was observed. A delay in significant expansion was observed after 14 days,
suggesting an induction period prior to expansion (a delay in the onset of expansion)
for the 0.4 M NaOH bath, rather than an alkali threshold, because expansion tends
to be >0.3% at 28 days. For the mortar bars exposed to 0.7 M and 1 M NaOH, it
appeared that, within the resolution of the measurements taken, the induction period
was relatively similar, with the onset to significant expansion occurring between 3
and 7 days. When the expansion rates were compared (Fig. 1), it could be seen that
increasing the NaOH concentration increased the expansion rate. It remains to be
seen with this aggregate whether threshold behavior is seen in maximum expansion
with respect to the change in alkali bath concentration.
Accelerated Mortar Bar Test to Assess the Effect of Alkali … 237

1
0.9 1M NaOH
0.8 0.7M NaOH
% Expansion 0.7
0.4M NaOH
0.6
Saturated Ca(OH)
0.5
0.4
0.3
0.2
0.1
0
0 5 10 15 20 25 30
Day

Fig. 1 Percentage expansion of natural sand mortar bars immersed in 0.4 M, 0.7 M and 1.0 M
NaOH solutions and saturated Ca(OH)2 solution over 28 days

Fig. 2 Expansion of mortars submerged in a 0.4 M NaOH, b 0.7 M NaOH, c 1.0 M NaOH and
d saturated Ca(OH)2 solutions at 80 °C for 28 days

No significant expansion was observed for the reactive aggregate mortar bars
prepared with blended cements containing FA or S (Fig. 2). These SCMs appeared
to sufficiently mitigate ASR within the timeframe of the experiment. Measurements
to extended ages will identify whether this is the result of complete mitigation or if
the expansion-free region is an induction period where the SCM acts as an inhibitor,
delaying the onset of expansion once consumed in the pozzolanic reaction.
238 B. Boyd-Weetman et al.

Table 2 Percentage expansion of mortar bars at age 28 days for 100% GP cement binder and
calculated pH for each alkali immersion bath
Alkali bath
Saturated Ca(OH)2 0.4 M NaOH 0.7 M NaOH 1 M NaOH
pH (calculated) 12.45 13.6 13.85 14
% Expansion of GP mix at 0.034 0.048 0.048 0.335
10 days
% Expansion of GP mix at 0.0248 0.230 0.576 0.796
21 days
% Expansion of GP mix at 0.0356 0.374 0.658 0.895
28 days

In summary, expansion increased with increasing concentration of NaOH solu-

tion, demonstrating that reactivity increases with increasing availability of alkali.
Additionally, the rate of expansion (slope of the expansion curves) was observed
to increase with increasing alkali concentration in the bath solution. An induction
period was observed for bars exposed to 0.4 M NaOH, and although this is not tradi-
tional threshold behavior, it does indicate that there is a mechanism affecting the
time for expansion to occur at lower alkali testing environments. As data collection
was limited to 28 days, the maximum possible expansion has not been determined.
Further measurements at greater ages will be carried out to identify any long-term
influence of alkali solution concentration on expansion to attempt to identify the
origins of the expansion behavior observed, with the aim of differentiating between
threshold behavior and the degree of reaction controlled as the reactivity or concen-
tration of the immersion solution. The incorporation of SCMs in the mortar mix
significantly reduced expansion, which indicated mitigation, but further measure-
ment is required to ensure that SCMs are not mitigating the reaction only in the short
term and delaying the onset of expansion.

4 Conclusions

As the alkali solution concentration increases, mortar bar expansion increases, which
indicates a relationship between alkali hydroxide concentration and ASR reaction
severity. Threshold behavior, as defined by RILEM, was not observed for this aggre-
gate at lower alkali levels; however, an induction period before the onset of significant
expansion was observed at the lower alkali bath concentration of 0.4 M. The expan-
sion data presented here is limited to the timeframe of the AMBT testing criteria,
so further extending AMBT experiment timeframes may clarify the relationship
between alkali concentration and extent of the ASR reaction with time, including
the maximum observed expansion. The investigation demonstrated that both FA and
S mitigate ASR-induced expansion in aggressive accelerated reaction environments
during testing at incremental concentrations up to 1 M NaOH.
Accelerated Mortar Bar Test to Assess the Effect of Alkali … 239

Acknowledgements This research was funded through an Australian Research Council Research
Hub for Nanoscience Based Construction Materials Manufacturing (NANOCOMM) with the
support of Cement Concrete and Aggregates Australia. The authors are grateful for the financial
support of the Australian Research Council (IH150100006) in conducting this study. This research
was supported by an Australian Government Research Training Program Scholarship.

References

1. Standards Australia (2015) Alkali Aggregate reaction—guidelines on minimising the risk of

damage to concrete structures in Australia. In: HB79
2. Australian Technical Infrastructure Committee (2019) ATIC-SP43. In: Section SP43(2019)
Cementitious Materials for Use with Concrete. ATIC, ATIC-SPEC
3. Nixon PJ, Sims I (2016) RILEM recommendations for the prevention of damage by Alkali-
aggregate reactions in new concrete structures, vol 17. Springer Netherlands, pp XVI, 168
4. Gavrilenko E et al (2007) Comparison of ASR-gels in concretes against accelerated mortar bar
test samples. Mag Concr Res 59(7):483–494
5. Vollpracht A et al (2016) The pore solution of blended cements: a review. Mater Struct
49(8):3341–3367
6. Shafaatian SMH et al (2013) How does fly ash mitigate alkali-silica reaction (ASR) in accelerated
mortar bar test (ASTM C1567)? Cement Concr Compos 37:143–153

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Development of High-Strength
Light-Weight Cementitious Composites
with Hollow Glass Microspheres

X. Li, Y. Yao, D. Zhang, Z. Zhang, and Y. Zhuge

Abstract Achieving both light-weight and high-strength cementitious composites

(HSLWCCs) is challenging. In this study, hollow glass microspheres (HGMs) were
used to develop a HSLWCC. Different amounts of HGMs were incorporated in
the cement mixture and the associated effects on the engineering properties and
microstructure were investigated. The results showed that the density and strength
decreased with increasing HGM content. Compressive strength of the HSLWCC
decreased significantly when the HGM content increased from 30 to 40% and
decreased slightly with further increasing HGM content, while the density generally
reduced linearly with increasing HGM content. Structural efficiency of the HSLWCC
increased when the HGM content was 30% and then decreased significantly at HGM
content of 40%. In particular, a floatable cementitious composite with a density of
~970 kg/m3 and compressive strength of ~31 MPa was developed by incorporating
60% of HGMs. Additionally, two failure modes (i.e., (i) debonding of interface
and (ii) crush of HGM) were found in the high-strength light-weight cementitious
composite (HSLWCC), with the former dominating in HSLWCC with high HGM
content and the later dominating in HSLWCC with low HGM content.

Keywords High-strength light-weight cementitious composites · Hollow glass

microspheres · Structural efficiency

X. Li · Y. Yao (B) · Z. Zhang

School of Civil Engineering, Xi’an University of Architecture and Technology, Xi’an, China
e-mail: yaoy@xauat.edu.cn
X. Li · Z. Zhang
XAUAT UniSA An De College, Xi’an University of Architecture and Technology, Xi’an, China
D. Zhang (B)
College of Civil Engineering, Fuzhou University, Fuzhou, China
e-mail: zhangdong_ce@fzu.edu.cn
Y. Zhuge
UniSA STEM, University of South Australia, Adelaide, SA, Australia

© The Author(s) 2023 241

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_25
242 X. Li et al.

1 Introduction

To date, cementitious composites are still one of the most widely used construc-
tion materials in the world [1, 2]. With increasing demand for long-span structures,
offshore platform, and precast prefabricated structures, the development of high-
strength, light-weight cementitious composites (HSLWCCs) is attracting attention.
The common methods used to produce them are the incorporation of air/bubbles
or the use of light-weight aggregates. However, instability of the air bubbles leads
to poor mechanical properties of the HSLWCC. Therefore, the use of light-weight
aggregates is considered to be a more promising approach [3, 4]. Although some
light-weight aggregates have good bonding to the cement matrix because of their
porous structure and pozzolanic reactivity [5–7], the developed light-weight aggre-
gate cementitious composites are still compromised by the lesser mechanical prop-
erties of the light-weight aggregates [6]. This has prompted researchers to look for
more suitable alternatives.
Hollow glass microspheres (HGMs) are a high-performance, ultra-light-weight
material consisting of a thin outer shell and an inert gas inside [8–10]. Compared
with conventional light-weight aggregates, HGMs have the advantages of superior
particle size distribution, lower density, higher strength, and sustainability [11], which
creates a new impetus to producing HSLWCCs. HGMs may have pozzolanic reaction
in the cement matrix due to their high content of SiO2 and Al2 O3 [12, 13]. It has
been shown that HGM can react with the surrounding cement matrix at low addition
amounts to form new gel particles and increase the bonding properties of HGM with
the matrix [14]. Nevertheless, the contribution of low additions of HGM to density
is not obvious, and the pozzolanic reactivity of HGMs at different additions has not
been well investigated.
Building on previous studies, the density and property changes of cementi-
tious composites produced with larger additions of HGM need to be clarified. The
pozzolanic reactivity of HGMs at different additions also needs to be discussed in
depth. Therefore, in this study, different additions of HGMs at 30, 40, 50 and 60% by
weight of cement were used to develop a HSLWCC, with a special aim of producing
a high-strength floatable cementitious composite. The variation in the density and
compressive strength of HSLWCC were investigated and the engineering proper-
ties of HSLWCC were evaluated by structural efficiency. Finally, the dispersion of
HGM in the matrix and pozzolanic reactivity were evaluated by scanning electron
microscopy (SEM).
Development of High-Strength Light-Weight Cementitious Composites … 243

2 Methods

2.1 Raw Materials

In this study, ordinary Portland cement (OPC, CEM I 52.5R) was used as the main
binder for the mixture. The HGMs were incorporated into the cement paste to reduce
the weight. The HGMs used in the mixture were ≈1–100 µm in size and had an
average particle density of ≈460 kg/m3 . They had a high compressive strength of
≈55 MPa. Figure 1 shows the SEM images and X-ray diffraction results of the
HGMs. As can be seen, the HGMs had a good non-crystalline structure and poten-
tial pozzolanic reactivity. Highly efficient polycarboxylate superplasticizer (SP) was
used to obtain suitable workability. The particle size distributions of both the OPC
and HGMs were obtained by a laser diffraction particle size analyzer, as shown in
Fig. 2. The chemical composition of the raw materials was determined using X-ray
fluorescence spectroscopy (XRF) and listed in Table 1.

2.2 Mix Proportions

The mix proportions are listed in Table 2. The amounts of HGMs added were 30,
40, 50, and 60% by weight of cement. Samples without HGMs were marked as REF,
while samples with HGMs were denoted by HSL and the replacement ratio of HGMs.
For example, the content of HGMs in HSL-30 was 30% of cement by weight. The
water-to-cement ratio was kept at 0.5. As extra HGMs were added to the mixture,
the SP content was adjusted to ensure similar workability of the different samples.
To prepare the samples with HGMs, SP was first solved in the water and the
solution was then added to the cement. The mixture was stirred at high speed for
3 min. The HGMs were added to the fresh mixture and the mixture were stirred for

Fig. 1 Scanning electron microscopy image a and X-ray diffraction b of hollow glass microspheres
244 X. Li et al.

Fig. 2 Particle size distribution of ordinary Portland cement and hollow glass microspheres
(HGMs)

Table 1 Chemical composition of materials, wt%

Type SiO2 Al2 O3 CaO MgO K2 O P2 O5 Fe2 O3 Na2 O SO3
Cement 19.39 4.0 63.61 2.4 0.72 0.13 3.82 0.59 4.66
HGM 80.3 0.85 6.31 0.25 0.05 0.05 0.04 11.5 0.25
HGM, hollow glass microsphere

Table 2 Mix proportions, wt%

Specimen Cement SF HGM SP Water
REF 1 0.2 0 – 0.5
HSL-30 1 0.2 0.3 – 0.5
HSL-40 1 0.2 0.4 0.01 0.5
HSL-50 1 0.2 0.5 0.02 0.5
HSL-60 1 0.2 0.6 0.03 0.5

another 6 min to achieve good workability. Subsequently, the mixture was poured
in to the mold on a vibrating table. After casting, samples were covered with plastic
films and stored under laboratory condition. After 24 h, the samples were demolded
and placed in a standard curing chamber until age 28 days under standard curing
conditions (temperature: 20 °C, humidity: >90%).
Development of High-Strength Light-Weight Cementitious Composites … 245

2.3 Experimental Methods

Cubes of 50 × 50 × 50 mm3 were prepared for the measurement of compressive

strength and density. The volume of the sample was tested by the water displacement
method, from which the density of the sample could be calculated. For the compres-
sion test, the samples were loaded at a rate of 1.5 kN/s. Three samples were used for
each mixture and the average value is reported. The microstructure of the different
samples was observed by SEM. The samples were soaked in the ethanol to stop the
hydration process and then dried in a vacuum chamber.

3 Results and Discussion

3.1 Density, Strength and Structural Efficiency

Figure 3a shows the density and compressive strength of the different samples. As
expected, the density of the HSLWCC decreased linearly with increasing HGM
content. With the inclusion of 30% of HGM, the density of HSL-30 decreased by
30% compared with the REF. When the HGM content increased to 60% of cement,
the density of HSL-60 reduced to 970 kg/m3 and it became floatable.
Figure 3a also shows that the inclusion of HGMs reduced the compressive strength
of the HSLWCC. However, it was interesting to discover that the decreasing rate of
compressive strength was not linear with increasing HGM content. With 30% HGMs,
the compressive strength of HSL-30 only reduced slightly by ~ 10.1% compared with
the REF samples. This might be caused by good bonding between the HGMs and the
cement matrix (see SEM results and the associated discussion in Sect. 3.2) and the
high compressive strength of HGM as mentioned in Sect. 2.1. With further increasing
of HGM content from 30 to 40%, the compressive strength of the HSLWCC reduced

Fig. 3 a Density, strength and b structural efficiency of high-strength light-weight cementitious

composite
246 X. Li et al.

significantly by 42% compared with the Ref samples. This may be due to the large
amount of HGMs disrupting the continuity of the cement paste [14]. However, the
decreasing rate of compressive strength slowed when the content of HGMs exceeded
40%. When the HGM content increased from 50 and 60%, the compressive strength
only reduced slightly. Although HSL-60 showed an ultra-low density <1000 kg/m3 ,
its compressive strength was still maintained at >30 MPa.
Structural efficiency, which is the ratio between compressive strength and density
(unit: kN·m/kg), is the main factor evaluating the lightness and strength of concrete.
A higher value of structural efficiency represents a higher specific strength (i.e., high
strength and light weight). Figure 3b shows the structural efficiency of the samples
with different additions of HGMs. As evident, the HSL-30 specimens showed the
highest structural efficiency. However, the structural efficiency decreased signifi-
cantly in HSL-40, caused by the significant reduction in compressive strength as the
density reduced linearly. When the HGM content was >40%, the changes in structural
efficiency were only slight. However, it should be noted that with high HGM content,
the structural efficiency of HSL-40, HSL-50 and HSL-60 did not reduce significantly
compared with the REF. HSL-50 showed the lowest structural efficiency among the
samples in this study, at 31.4, compared with the structural efficiency of the REF at
36.3.

3.2 SEM Analysis

Figure 4 shows the SEM images of the samples. The fracture surface of the HSL-30
sample shown in Fig. 4a indicates that the HGMs were uniformly dispersed in the
matrix. Besides, Fig. 4b also shows that the HGMs were well bonded to the matrix
and the hydration product was still partially attached to the surfaces of the HGMs
after fracture. As shown in Fig. 1b, HGMs had potential pozzolanic reactivity. The
reaction between HGMs and the cement matrix could further enhance the bonding
[15, 16]. However, it should be noted that the bonding between the HGMs and cement
matrix was still weak, as the reaction was limited, resulting in debonding between the
HGMs and the cement matrix under loading, as shown in Fig. 4c. Besides, crushing
of HGMs was also found in the samples after the compression test, as shown in
Fig. 4d. Generally, two failure modes occurred in the HSLWCC: (i) debonding of
the interface and (ii) crushing of HGMs. In the samples with lower HGM additions,
failure mode (i) was the dominant damage mode, and in the samples with high HGM
additions, damage mode (ii) dominated.
Development of High-Strength Light-Weight Cementitious Composites … 247

Fig. 4 a Fracture image of HSL-30; b HGM nucleation reaction; c morphology of HSL-60; d failure
mode of the samples

4 Conclusions

In this study, a HSLWCC was developed using HGMs as the light-weight filler.
In particular, a novel floatable cementitious composite with an apparent density of
~970 kg/m3 and compressive strength of ~31 MPa was developed.
By incorporating HGMs with a content of 30–60% by weight of cement, the
density of the developed HSLWCC ranged from 970 to 1340 kg/m3 and the compres-
sive strength ranged from 31 to 62 MPa. The compressive strength of the HSLWCC
decreased significantly when the HGM content increased from 30 to 40%. With
further increasing of HGM content, the compressive strength only reduced slightly.
The density of the HSLWCC almost decreased linearly with increasing HGM content.
The structural efficiency of the HSLWCC showed a sudden significant increase at
a HGM content of 30%, while the structural efficiency of other HSLWCC samples
was slightly lower than that of the reference sample.
Debonding of the interface and crushing of the HGM were both found at the
fracture surface. The debonding of interface dominated in mixtures with high HGM
content, whereas crushing of HGMs was usually found in the mixtures with low
HGM content.
248 X. Li et al.

References

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high performance graphene-based cement material. Constr Build Mater 335:127451
2. Zhang D, Shahin MA, Yang Y, Liu H, Cheng L (2022) Effect of microbially induced calcite
precipitation treatment on the bonding properties of steel fiber in ultra-high performance
concrete. J Build Eng 50:104132
3. Liu X, Chia KS, Zhang M-H (2011) Water absorption, permeability, and resistance to chloride-
ion penetration of lightweight aggregate concrete. Constr Build Mater 25(1):335–343
4. Tandiroglu A (2010) Temperature-dependent thermal conductivity of high strength lightweight
raw perlite aggregate concrete. Int J Thermophys 31(6):1195–1211
5. Chen R, Mo KH, Ling T-C (2022) Offsetting strength loss in concrete via ITZ enhancement:
From the perspective of utilizing new alternative aggregate. Cement Concr Compos 127:104385
6. Lu J-X, Shen P, Ali HA, Poon CS (2021) Development of high performance lightweight
concrete using ultra high performance cementitious composite and different lightweight
aggregates. Cement Concr Compos 124:104277
7. Duan W, Zhuge Y, Pham PN, Liu Y, Kitipornchai S (2022) A ternary blended binder incor-
porating alum sludge to efficiently resist alkali-silica reaction of recycled glass aggregates. J
Clean Prod 349:131415
8. Liu B, Wang H, Qin Q-H (2018) Modelling and characterization of effective thermal
conductivity of single hollow glass microsphere and its powder. Materials 11(1):133
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and filler volume fraction on the mechanical properties of hollow glass microsphere modified
polymer. Compos Part B Eng 101:53–63
10. Shahidan S, Aminuddin E, Mohd Noor K, Ramzi Hannan NIR, Saiful Bahari NA (2017)
Potential of hollow glass microsphere as cement replacement for lightweight foam concrete on
thermal insulation performance. In: MATEC web of conferences, vol 103. EDP Sciences, Les
Ulis, p 1014
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nanofibers and activated carbon powder on mechanical and dry shrinkage performance of
ultra-lightweight engineered cementitious composites. Constr Build Mater 280:122415
12. Krakowiak KJ, Nannapaneni RG, Moshiri A, Phatak T, Stefaniuk D, Sadowski L, Abdolhos-
seini Qomi MJ (2020) Engineering of high specific strength and low thermal conductivity
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ation of their influence on density, mechanical strength and water absorption. Cement Concr
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composites using hollow glass microspheres. Struct Concr J FIB 21(2):673–688
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quantification of hollow glass microspheres. Cement Concr Compos 118:103981
Development of High-Strength Light-Weight Cementitious Composites … 249

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Co-effects of Graphene Oxide and Silica
Fume on the Rheological Properties
of Cement Paste

D. Lu, Z. Sheng, B. Yan, and Z. Jiang

Abstract Polycarboxylate superplasticizer is typically used to prepare a high-

quality graphene oxide (GO) solution before mixing with cement grains. However,
even if GO is well dispersed in water, they tend to re-agglomerate in the alkaline
cement hydration environment, thus seriously decreasing the workability of the fresh
mixture. In this study, we propose a more targeted method by synthesizing GO-
coated silica fume (SF) to promote the utilization of GO in cement-based materials.
Specifically, the surface of pristine SF was modified to convert their zeta potential
(modified SF: MSF), then GO-coated SF (i.e., MSF@GO) was prepared via electro-
static adsorption of GO onto the MSF surface. The experimental results revealed that
adding 5MSF@GO hybrid (0.04% GO and 5% MSF, by weight of binder) signifi-
cantly reduced yield stress and plastic viscosity by 51.5% and 26.2%, respectively,
relative to the 0.04% GO-modified sample. These findings indicated that applica-
tion of GO-coated SF is an effective and environmentally friendly way to develop
sustainable cementitious composites.

D. Lu
Department of Civil and Environmental Engineering, The Hong Kong Polytechnic University,
Hong Kong, China
School of Civil Engineering, Harbin Institute of Technology, Harbin, China
Z. Sheng
Key Laboratory of Concrete and Prestressed Concrete Structures of Ministry of Education, School
of Civil Engineering, Southeast University, Nanjing, China
B. Yan
Wuhan Harbor Engineering Design and Research Institute Co., Ltd, Wuhan, China
Hubei Key Laboratory of New Materials and Maintenance and Reinforcement Technology for
Offshore Structures, Wuhan, China
Z. Jiang (B)
Department of Civil and Environmental Engineering, The Hong Kong University of Science and
Technology, Clear Water Bay, Hong Kong, China
e-mail: zhenliang.jiang@connect.ust.hk

© The Author(s) 2023 251

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_26
252 D. Lu et al.

Keywords Cement composites · Graphene oxide · Rheological property · Silica

fume · Surface modification

1 Introduction

Cement-based materials are the most extensively used construction materials because
of their low cost, high compressive strength, and durability [1, 2], with an estimated
yearly consumption of >30 billion tonnes [3]. However, their brittle nature limits
broad application in some structures [4]. Additionally, the cement/concrete industry is
energy intensive with a substantial environmental footprint [5]. According to previous
studies [5, 6], enhancing the microstructure of cement-based materials and improving
their mechanical strengths/durability is considered to be a candidate way to alleviate
carbon emissions. Calcium–silicate–hydrate (C–S–H) gel, the principal hydrate of
cement grains, is composed of nanocrystalline with an atomic structure similar to
that of tobermorite and/or jennyite (i.e., it is a nanoscale material) [7]. Advancements
in nanomaterials and nanotechnology have been providing great opportunities to
enhance the structure of cement composites at the nanoscale, eventually improving
the macroscale properties [7, 8].
As a typical two-dimensional nanomaterial, graphene oxide (GO) has been consid-
ered a favorable additive for improving the mechanical strengths and durability of
cement-based materials [5, 9]. However, GO tends to agglomerate in an alkaline
hydration environment when cement grains dissolve in water [10], which dramati-
cally deteriorates the workability of fresh mixtures. Note that, the fresh properties
of mixtures greatly affect the mechanical and durability properties of the hardened
composites.
Polycarboxylate superplasticizer (SP) is generally used both for preparing high-
quality GO solution and for improving the fresh properties of the cement pastes
[10]. However, obtaining a high-quality GO solution before mixing with cement
does not directly result in well-dispersed GO in the alkaline cement matrix. It has
been demonstrated [9, 11] that using silica fume (SF) can predisperse GO in the
cement matrix. These studies claim that the electrostatic repulsion between nega-
tively charged SF and GO is primarily responsible for the improved GO disper-
sion. Nevertheless, the electrostatic repulsion theory may not apply to such a system
because of the remarkably larger lateral size of GO than that of SF [11]. Inspired by
GO-coated sand, which can reduce the migration resistance in water and improve
the adsorption ability in water treatment applications [12], we developed GO-coated
SF via electrostatic adsorption of GO onto the surface of the modified SF (MSF),
aiming to better utilize GO in cementitious composites and exert the co-effects of
the two materials. Especially, this work focused on investigating the properties of
fresh cement pastes incorporating GO-coated SF, aiming to promote the application
of GO in environmentally friendly high-performance cement composites.
Co-effects of Graphene Oxide and Silica Fume on the Rheological … 253

2 Methods

2.1 Materials

Cement (P·O 42.5) and SF were used to prepare cement paste. GO was synthesized
by modified Hummers’ method, leading to a specific surface area (SSA) of ~2600
m2 /g. The chemical bonds in the GO used in this study mainly contained C–O, C =
O, C = C, and O–H. The XPS data revealed that the C/O was ~1.97.

2.2 Preparation of GO-Coated SF and Cement Paste

To improve the compatibility of GO and the cement matrix, the concept of GO-coated
SF was proposed, enabling the co-effects of GO and SF in the cement composite.
Specifically, MSF particles were obtained by treating the surface of SF with Ca(OH)2
solution. To eliminate potential chemical reactions on the SF surface, the modification
process was optimized according to previous experience [5], where the SF was added
to the Ca(OH)2 aqueous solution at a weight ratio of 1:10. After that, the SF/MSF
was mixed with the GO solution and stirred for 10 min to synthesize SF-GO or GO-
coated SF. All pastes were fabricated by mechanically stirring for 4 min. Finally, the
fresh pastes were used for rheological properties tests.

2.3 Testing Methods

The micromorphology was observed using a ZEISS electron microprobe. The surface
functionality of particles was assessed via a Nano ZS zeta potential analyzer.
SmartLab XRD with an incident beam of Cu-Ka radiation (λ = 1.54 Å) for a 2θ
scanning range of 15–65° was used to examine the crystalline phase analyses of the
powder samples. A high-resolution FEI-TEM was usd to compare the morphology
of SF-GO and GO-coated SF. A Brookfield RST-SST rheometer equipped with a
rotating vane (VT20-10) was used to perform the rheological tests. During testing,
the shear rate increased from 5 to 150 s−1 in 60 s and a corresponding decrease
in shear rate from 150 to 5 s−1 in the following 60 s. The yield stress and plastic
viscosity can be obtained as follows:

τ = τ0 + μγ + cγ 2 (1)

where τ is the shear stress (Pa), γ is the shear rate (1/s), μ is the plastic viscosity
(Pa·s), τ 0 is the yield stress (Pa), and c is the second-order coefficient (Pa·s2 ).
254 D. Lu et al.

3 Results and Discussion

3.1 Evaluation of the Surface Properties of MSF Particles

and the MSF@GO

Considering SF particles have the potential to react with Ca(OH)2 solution, as such
the modification process was optimized, aiming to ensure that any chemical reaction
on the SF surface could be discounted.
As presented in Fig. 1a, b, SF and MSF both exhibited a spherical shape with a
size ranging from 50 to 300 nm. The XRD patterns of the SF and MSF were almost
indistinguishable, without the broad peaks of C–S–H gel found in MSF samples
(Fig. 1c). As suggested in Fig. 2d, the zeta potential of the SF (–23 mV) converted
to ≈+ 3 mV (MSF), thanks to some calcium ions grafted onto the SF surface. These
findings all support that MSF maintained the surface morphology and crystalline
phase after treatment, its surface only achieving ion exchange.
As indicated in Fig. 2, a thin layer of GO was found uniformly and tightly adsorbed
onto the surface of the MSF (Fig. 3b). In contrast, several SF particles were merely
interspersed between GO layers (Fig. 3a). Such a strong MSF@GO interaction
showed great potential to exert their co-effects in cement-based materials.

Fig. 1 Properties of the silica fume (SF) particles before and after modification: scanning electron
microscopy images of a SF, b modified SF (MSF); c X-ray diffraction patterns; and d zeta potential
Co-effects of Graphene Oxide and Silica Fume on the Rheological … 255

Fig. 2 Morphology of a SF-GO and b MSF@GO; c schematic diagram of SF particles located

between GO layers (SF-GO) and GO-coated SF (MSF@GO). GO, graphene oxide; MSF, modified
SF; SF, silica fume

Fig. 3 Rheological parameters of the paste: a shear stress–shear rate curves and b viscosity–shear
rate curves. GO, graphene oxide; MSF, modified SF; SF, silica fume

3.2 Rheological Properties

The shear stress–shear rate curves shifted upwards after admixing 0.04 wt% GO,
as compared with plain paste (Fig. 3a). Simultaneously, the viscosity of the 0.04
wt% GO-modified paste also increased at both high and low shear rates (Fig. 3b).
Specifically, the yield stress and plastic viscosity of the pastes were calculated based
on an improved Bingham’s model: adding 0.04 wt% GO to the paste increased the
yield stress and plastic viscosity by 92.5% (Fig. 3b) and 88.1% (Fig. 3c), respectively.
256 D. Lu et al.

Fig. 4 a Linear regression of shear stress–shearrate curves; calculated b yield stress and c plastic
viscosity. GO, graphene oxide; MSF, modified SF; SF, silica fume

Adding SF or MSF decreased the yield stress and plastic viscosity of the pastes
(Fig. 4b, c). The mixture of 5SF-GO exhibited the highest shear stress–rate curve
(Fig. 3), with the highest yield stress (66.1 Pa) and plastic viscosity (0.54 Pa·s), which
implied that this formulation of cement composite has poor workability and is not
suitable for practical applications. Additionally, the yield stress and plastic viscosity
of the 5MSF@GO mixture decreased by 51.5% and 26.2%, respectively, relative to
the 0.04 wt% GO-modified sample (Fig. 4). The MSF@GO hybrid adsorbed GO
onto the MSF, thereby making the spherical particles easily migrate. In addition,
the negatively charged GO coated onto the surface of the MSF, which provided
electrostatic repulsion among particles, enabled better dispersion of GO (and MSF)
and released entrapped water to turn into free water.

4 Conclusions

In this study we developed a high-performance cement composite incorporating GO-

coated SF. The surface properties of GO-coated SF, the rheological properties, and
compressive strength of composites were studied.
The surface of pristine SF particles was modified to convert their zeta potential;
while the surface morphology and chemical composition remained unchanged, the
SF surface merely achieved ion exchange. The positively charged MSF allowed
negatively charged GO to adsorb on its surface via electrostatic adsorption. The
Co-effects of Graphene Oxide and Silica Fume on the Rheological … 257

strong bonding showed great potential for exerting a cooperative improvement effect
in cement composites.
Different from the traditional direct introduction of GO or SF into cement paste,
which seriously increased the yield stress and plastic viscosity of the mixtures.
Adding 0.04 wt% GO together with 5 wt% MSF decreased the yield stress and
plastic viscosity by 51.5% and 26.2%, respectively. The improved fresh properties
provide guarantees for the transportation and construction of the cement composites.

References

1. Lu D, Ma LP, Zhong J, Tong J, Liu Z, Ren W, Cheng HM (2023) Growing nanocrystalline

graphene on aggregates for conductive and strong smart cement composites. ACS Nano
2. Lu D, Zhong J (2022) Carbon-based nanomaterials engineered cement composites: a review. J
Infrastruct Preserv Resil 3(1)
3. Monteiro PJM, Miller SA, Horvath A (2017) Towards sustainable concrete. Nat Mater
16(7):698–699
4. Lu D, Shi X, Zhong J (2022) Understanding the role of unzipped carbon nanotubes in cement
pastes. Cem Concr Compos 126
5. Lu D, Shi X, Zhong J (2022) Interfacial nano-engineering by graphene oxide to enable better
utilization of silica fume in cementitious composite. J Clean Prod
6. Pan Z, He L, Qiu L, Korayem AH, Li G, Zhu JW, Collins F, Li D, Duan WH, Wang MC (2015)
Mechanical properties and microstructure of a graphene oxide–cement composite. Cem Concr
Compos 58:140–147
7. Singh NB, Kalra M, Saxena SK (2017) Nanoscience of cement and concrete. Mater Today:
Proc 4(4):5478–5487
8. Lu D, Shi X, Zhong J (2022) Interfacial bonding between graphene oxide coated carbon
nanotube fiber and cement paste matrix. Cem Concr Compos 134
9. Lin J, Shamsaei E, Basquiroto de Souza F, Sagoe-Crentsil K, Duan WH (2020) Dispersion of
graphene oxide–silica nanohybrids in alkaline environment for improving ordinary Portland
cement composites. Cem Concr Compos 106
10. Zhao L, Guo X, Liu Y, Ge C, Chen Z, Guo L, Shu X, Liu J (2018) Investigation of dispersion
behavior of GO modified by different water reducing agents in cement pore solution. Carbon
127:255–269
11. Lu Z, Hou D, Hanif A, Hao W, Li Z, Sun G (2018) Comparative evaluation on the dispersion
and stability of graphene oxide in water and cement pore solution by incorporating silica fume.
Cem Concr Compos 94:33–42
12. Hou W, Zhang Y, Liu T, Lu H, He L (2015) Graphene oxide coated quartz sand as a high
performance adsorption material in the application of water treatment. RSC Adv 5(11):8037–
8043
258 D. Lu et al.

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Automated 3D-Printer Maintenance
and Part Removal by Robotic Arms

K. Andrews, K. Granland, Z. Chen, Y. Tang, and C. Chen

Abstract 3D printing by means of fused filament fabrication involves extruding and

depositing melted material in layers to produce a 3D part. Current 3D printing requires
manual intervention from a human operator between prints, leading to inefficiency.
The focus of this study was facilitating the automation of the additive manufacturing
process. Based on suggestions for future works in this field, this study extended on
automated 3D-part removal systems by implementing additional operations to auto-
mate the production process. The proposed system uses robotic arms and grippers
to operate and maintain 3D printers; specifically, the removal of 3D-printed parts,
the cleaning of printer beds, the application of glue to the printer beds to assist with
print adhesion, and the monitoring of bed levelness. The importance of this contri-
bution is the improved efficiency of 3D-printing production, allowing for continuous
3D-printer operation and decreasing the requirement for human interaction and moni-
toring in the production process. The system is demonstrated using a 7 degrees of
freedom KUKA robotic arm and ROBOTIQ gripper to autonomously operate and
maintain an Ender 3 V2 printer. Sensor data and information from the 3D printers was
used to determine the required operation or function to be performed by the robotic
system. Tasks were performed by automated movement sequences of the robotic
arm and gripper using supplied data. System status was recorded for monitoring
and alerting human operators when intervention was required. The implementation
of these functions using an automated robotic system allows 3D-printing produc-
tion to operate continuously for longer periods, increasing production efficiency as
downtime and human involvement for maintenance between prints is minimized.

Keywords 3D printing · Automated robotic system · Smart manufacturing

K. Andrews · K. Granland · Z. Chen · Y. Tang · C. Chen (B)

Digital Twin Laboratory, Monash Smart Manufacturing Hub, Monash University, Clayton, VIC,
Australia
e-mail: chao.chen@monash.edu

© The Author(s) 2023 259

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_27
260 K. Andrews et al.

1 Introduction

Monash Smart Manufacturing Hub/Digital Twin Lab (DTL) incorporates automated

processes and data collection to efficiently manufacture, store and monitor 3D-printed
parts. Data and information from the physical systems are used to create a digital twin
that allows for all relevant information to be monitored, displayed and controlled via
a computer. 3D printing is performed by melting and extruding material in layers
to form a 3D part. 3D printers require external interaction for the removal of parts,
and other maintenance tasks need to be regularly performed to allow for continuous
printer operation. In this study, we propose the use of a KUKA robotic arm and
ROBOTIQ 2F-85 gripper to automate the 3D-printing process of Ender 3 V2 printers.
The system uses information regarding printer and 3D-part status from the printers
and outputs data on task completion status for use by the DTL. This integration of
physical processes and information sharing will allow the system to be incorporated
into the Monash Smart Manufacturing Hub.
Related works have explored the application of robotic arms to remove prints by
exchanging flexible printer beds after each print [1, 2] or by using specialized end-
effector mechanisms to directly pick up complex 3D-printed parts [3]. Consumer
products [4, 5] allow for continuous printing with conveyor-belt beds or attached
pushing mechanisms [6, 7] to slide parts from the end of the print bed. Here, we
propose to build on functionalities from these implementations to create an auto-
mated system that can address the operation and maintenance requirements that
arise from 3D-printing production. Although related works have focused on devel-
oping standalone automation features such as part removal, we propose an integrated
system that can perform a range of functions so that the 3D-printing workflow can
be fully automated.
Our objective was to demonstrate the implementation of an automated system
using a robotic arm and gripper to enable the workflow shown in Fig. 1 to remove 3D-
printed parts from the printers, apply glue to print beds for improved layer adhesion,
clean the print beds and check bed levelness. These functionalities were implemented
by programming the robotic arm and gripper to perform movement sequences and
interact with the 3D printer to achieve continuous 3D-printer operation.

Fig. 1 Automated printer maintenance and operation workflow

Automated 3D-Printer Maintenance and Part Removal by Robotic Arms 261

2 Methods

2.1 Literature Review

3D printing has the advantage of requiring minimal setup compared with other manu-
facturing techniques, enabling parts to be printed on demand quickly and cheaply.
The ability to continuously print parts is an important aspect of efficient 3D printing,
and requires maintenance and operation tasks to be performed regularly on the 3D
printers.
Removing finished 3D prints by use of robotic arms has been addressed [1, 2]
with a method consisting of a detection algorithm for failed prints and a part removal
function consisting of flexible magnetic beds that are removed and replaced. The
advantage of using flexible magnetic beds is that complex parts can be removed
easily because the robotic arm does not interact directly with the printed parts.
Becker [3] implemented a system capable of analyzing complex 3D-printed
parts and determining a satisfactory part removal process using a robotic arm and
customized end-effector/gripper for part removal. A CAD-based implementation [8]
used grasping and motion planning simulation to determine valid end-effector paths
for removing 3D-printed objects, which allowed model geometry to be used to deter-
mine optimal gripping points. Vision-based implementation [9], uses a depth camera
and reinforcement learning methods to pick and place objects in a simulation envi-
ronment. Such approaches provide solutions to automated 3D-part removal that can
adapt to various part geometries. Becker [3] found that this saved time and decreased
the requirement for human involvement in the part removal process, while noting that
future works should include transitioning to a robotic arm system on a mobile base
to enable interaction with other machines and increasing the flexibility of automated
tasks that can be performed.
Aroca et al. [10] implemented a 2-degree of freedom manipulator and monitoring
system to remove 3D-printed parts to enable continuous 3D printing. Future works
are proposed for the use of a robotic arm to also apply glue to the printer bed for
improved print adhesion.
Consumer products [4, 5] allow for continuous printing with a conveyor-belt setup,
which is advantageous for prints that require a long z-axis because the part can be
moved along the belt to be extended, and parts are pushed off the end of the bed,
so additional mechanisms are required for the handling of parts after the printing
stage in an automated process. Numerous third-party products and mechanisms [6,
7, 11] are also available that sweep across the print bed to push parts forward and
off the print bed. However, these do not check if parts are unstuck properly because
it is a ‘blind’ process and does not consider if the bed is sufficiently clean. Also,
additional mechanisms are required for handling of parts after the printing stage
as they are pushed to fall off the edge of the printer into a pile. Another aspect
of enabling continuous 3D printing is ensuring workflow efficiency. Jim and Lees
[12] demonstrate how task sequencing efficiency improvements of a robot can be
implemented in the automation of 3D printing and post processing. By optimizing the
262 K. Andrews et al.

efficiency of handling parts and moving through a sequence of functions a workflow

can be created for producing parts that requires minimal human interaction. However,
they noted that further work is required to consider when 3D printers or other post-
processing machines are broken or require maintenance.
These works formed the basis for the functionalities to be implemented in this
study. As highlighted, the additional automation of 3D-printer maintenance tasks and
other supporting operations can further improve continuous 3D printed part produc-
tion and efficiency [10, 12]. We addressed the gap in this field by demonstrating
numerous automated tasks that are required for continuous 3D printing. The auto-
mated application of glue to the printer bed and cleaning of the bed after prints are
removed allows for continual 3D printing and decreasing the need for human input
and monitoring. The physical checking of bed levelness can be used to alert a human
operator when maintenance tasks are required before printing begins. Automated part
removal can minimize printing downtime so that continuous 3D-printing production
can be better achieved.

2.2 Automated System

The automated system set up is shown in Fig. 2 with the KUKA robotic arm and
gripper on a mobile base positioned in front of the workbench. An Ender printer is
mounted with 3D-printed brackets to the workbench. The tool holder mounted next
to the printer contains a glue stick and sponge.

Fig. 2 Automated system setup

Automated 3D-Printer Maintenance and Part Removal by Robotic Arms 263

Fig. 3 The 3D printer and tool holder CAD environment

In order to automate the maintenance and operation tasks when 3D printing, print
completion status, print size and print location information is required. This data is
provided as variables in the KUKA robot’s code and used as parameters that dictate
robotic arm and gripper movements. It is proposed that the printer data is sent from
the DTL to the robotic system in future works.
The movement sequences for the KUKA robotic arm were programmed in Java
using Sunrise Workbench. The gripper end-effector was moved with translation and
rotation commands and combined with gripper position commands to create complex
movement sequences. Figure 3 shows the layout of the 3D printer and tool holder
modeled in CAD. This is used to determine the path for the robot.
Coordinate frames consisting of gripper position and rotation values were saved
for frequently used positions. The movement of the gripper and arm was also
controlled by torque sensing on the robotic arm joints and force sensing in the
gripper. This feedback was used to dictate the start and end of certain movement
sequences.

3 Results and Discussion

3.1 Robot–Printer Calibration

In order to achieve precise movements and interactions between the robotic arm and
3D printer, calibration was required so that all arm movements were performed rela-
tive to the printer. The gripper was calibrated using two calibration points mounted
to the top of the printer frame. As shown in Fig. 4, calibration was performed by
initially positioning the gripper approximately in front of the first calibration point
264 K. Andrews et al.

Fig. 4 Calibration between the robotic arm and 3D printer

then translating the gripper in x and z directions until a force was detected; the posi-
tions where the force was detected were saved, providing a known point for the robot
to reference. The second calibration point was used to measure the printer’s rotation
relative to the base of the robot. By comparing the difference in positions between
both calibration points the required gripper rotation of the gripper was calculated.
The gripper was rotated to be square with the printer so that all arm movements were
linear relative to the printer. This calibration process allowed positions on the printer
to be programmed relative to the known reference frame at the calibration points for
accurate arm movements when performing tasks. This process could be substituted
for a computer vision system that detects the position of the gripper relative to the
printer in future. A demonstration video is available at https://github.com/Kai-and
rews/-Automated-3D-printer-maintenance-and-part-removal-by-robotic-arms.

3.2 Bed Levelness Monitoring

The levelness of the printer bed was monitored by an automated sequence moving
the robotic gripper across each corner of the printer bed and lowering the gripper onto
the bed while recording the relative heights where a force threshold was detected, as
shown in Fig. 5. The height differences are displayed on the KUKA console to alert
human operators as to which side of the printer bed requires adjusting to achieve a
level printer bed. In future the bed levelness data will be sent to the DTL to alert
operators of maintenance requirements if the bed is not sufficiently level. The printer
bed was manually set to different states of levelness and detected using the gripper.
As shown in Table 1, the robotic arm and gripper could detect all bed levelness
configurations accurately. Please see the demonstration video.
Automated 3D-Printer Maintenance and Part Removal by Robotic Arms 265

Fig. 5 Automated bed level detection implementation

Table 1 Bed level detection position values

Bed setup Front left (mm) Front right (mm) Back left (mm) Back right (mm)
Level 60.49 60.64 60.77 60.76
Front higher than back 60.51 60.57 63.42 63.51
Left higher than right 63.33 60.55 63.83 60.81

3.3 Application of Glue for Improved Part Adhesion

Automated application of glue to the printer bed was achieved using a glue stick held
in a 3D-printed mount consisting of a round hole and slot cut-out for the glue stick
to rest on. As shown in Fig. 6, the glue stick had an attached flat section that slotted
into the holder to prevent the glue stick from rotating as the robotic arm rotated the
glue stick knob to extend the glue. A 3D-printed attachment was mounted on the
glue stick, providing a suitable surface for the gripper to hold the glue stick when in
use. The glue stick was raised from the holder and lowered onto the print bed. As
depicted in Fig. 7, glue was applied to the print bed by moving the glue stick over the
area where a part was to be printed. Force was monitored to ensure glue was applied
evenly across the bed by moving the arm down to maintain 8 N of vertical force to
ensure contact as glue was used. The area that glue was applied to was determined
using position and size information of the part that was to be printed. After glue
application, the glue stick was returned to the holder. As shown in Fig. 7 the arm was
able to apply glue evenly, fully covering the print area and in the correct position
where the part was to be printed. Please see the demonstration video.

3.4 Part Removal

The automated removal of 3D-printed parts used software inputs that described the
center point position of the part and its base length and width. The printer bed was
266 K. Andrews et al.

Fig. 6 Automated glue extension process

Fig. 7 Automated glue application and glue coverage

allowed to sufficiently cool before the robotic gripper was positioned above the part
and opened. As shown in Fig. 8, the gripper was then lowered around the part until
contact with the printer bed was detected and the gripper closed using force sensing so
that the base of the part was securely held. The robotic arm then performed a sequence
of rotations sideways, forward and backwards as shown in Fig. 9 while torque and
position monitoring were used to determine if the part was stuck (indicated if torque
was in the joints above a threshold) to prevent damage to the part or printer bed.
The part was lifted from the printer bed and placed in a designated location next
to the printer. This method allowed for parts with simple geometry and <85 mm
(the grippers open width) to be removed successfully from the printer bed. Future
improvements for the automated removal of more complex printed parts include
using a more specialized end-effector and gripping techniques customized for each
part. Please see the demonstration video.

3.5 Bed Cleaning

The cleaning of the printer bed was performed using a sponge with a handle
containing a cleaning solution, which rested on a 3D-printed holder. As shown in
Fig. 10, an attachment with flat sides was mounted to the handle to provide a suitable
gripping point for the robotic gripper. The sponge was picked up by the robotic arm
Automated 3D-Printer Maintenance and Part Removal by Robotic Arms 267

Fig. 8 Automated part removal demonstration

Fig. 9 3D-printed part removal

and moved over the print bed, then lowered until a vertical force of 7 N was detected,
indicating that the sponge was fully in contact with the printer bed. The sponge was
then moved forwards and backwards repetitively to remove dried glue and plastic
build-up. Cleaning solution was applied to the sponge by moving the sponge up and
down against the printer bed to squeeze the cleaning solution from the handle and
sponge. The sponge was returned to the holder after printer bed cleaning. As shown
in Fig. 11, the automated cleaning of the printer bed was able to successfully remove
glue and residue build-up. This process could be further improved by implementing
an automated scraping process using a scraping tool to remove larger amounts of
physical contamination stuck to the printer bed. Please see the demonstration video.

Fig. 10 Automated printer bed cleaning process

268 K. Andrews et al.

Fig. 11 Comparison between before and after automated bed cleaning

Table 2 Execution times for

Function Completion time (s) Success rate
automated functionalities
Arm calibration 26 10/10
Bed level detection 20 10/10
Glue application 43 10/10
Part removal 26 10/10
Bed cleaning 64 10/10
Total 179 100%

3.6 Overall System Discussion

As shown in Table 2, the robotic arm and gripper repeatedly performed all automated
functionalities consistently, completing 100% of test runs. Successful execution was
defined as the robotic arm and gripper being able to correctly determine bed levelness,
apply sufficient glue over the entire required area for printing, successfully remove
the 3D-printed part and place it in a designated location and sufficiently clean the
print bed so that another print can take place. The execution time for each automated
process demonstrates the ability of the system to efficiently perform the required
operation and maintenance tasks for 3D printing with a total execution time of 2 min
59 s, which can be improved in future with further velocity increases to the robotic
arm movements.

4 Conclusions

In this paper, automation of the additive manufacturing process using a robotic arm
and gripper is proposed and described. An automation sequence was developed to
facilitate bed levelness detection with the ability to record data for monitoring by
human operators, glue application to assist in print adhesion covering 100% of the
Automated 3D-Printer Maintenance and Part Removal by Robotic Arms 269

print area, 3D-printed part removal and printer bed cleaning sufficiently to allow
for another print to occur. The contribution of this research is the demonstration of
an automated robotic system that performs required 3D-printer operation and main-
tenance tasks for continuous unmanned 3D-printing production. Recommendations
for improvements and future work include creating custom tools for glue application
and bed cleaning that allow for improved workflow using the automated robotic arm
and gripper. Calibration between the robotic gripper and 3D printer can be improved
using a computer vision system. The automated removal of parts can be optimized
using adaptive or specialized gripper mechanisms for smart gripping of more complex
3D-printed parts. The data outputs on system status such as bed levelness or tasks
completion can also be integrated into the DTL. We have demonstrated the ability
of a robotic arm and gripper to enable continuous unmanned 3D printing through
automation of key processes.

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Commons license, unless indicated otherwise in a credit line to the material. If material is not
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the copyright holder.
Machine Vision-Based Scanning Strategy
for Defect Detection in Post-Additive
Manufacturing

S. Zhang, Z. Chen, K. Granland, Y. Tang, and C. Chen

Abstract The surge in 3D printer availability, and its applications over the past
decade as an alternative to industry-standard subtractive manufacturing, has revealed
a lack of post-manufacturing quality control. Developers have looked towards auto-
mated machine learning (ML) and machine-vision algorithms, which can be effec-
tive in developing such additive manufacturing (AM) technologies for industry-wide
adoption. Currently, most research has explored in-situ monitoring methods, which
aim to detect printing errors during manufacturing. A significant limitation is the
single, fixed monitoring angle and low resolution, which fail to identify small or
hidden defects due to part geometry. Therefore, we investigated a novel ex-situ
scanning strategy that combines the advantages of robotics and machine vision to
address the limitations; specifically, the viability of image-recognition algorithms
in the context of post-fabrication defect detection, and how such algorithms can be
integrated into current infrastructure by automatically classifying surface faults in
printed parts. A state-of-the-art and widely accepted ML-based vision model, YOLO,
was adapted and trained by scanning for prescribed defect categories in a sample of
simple parts to identify the strengths of this method over in-situ monitoring. An
automated scanning algorithm that uses a KUKA robotic arm and high-definition
camera is proposed and its performance was assessed according to the percentage of
accurate defect predictions, in comparison with a typical in-situ model.

Keywords 3D printing · Additive manufacturing · Defect detection · Machine

vision · Neural networks

S. Zhang · Z. Chen · K. Granland · Y. Tang · C. Chen (B)

Digital Twin Laboratory, Monash Smart Manufacturing Hub, Monash University, Clayton,
VIC 3800, Australia
e-mail: chao.chen@monash.edu

© The Author(s) 2023 271

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_28
272 S. Zhang et al.

1 Introduction

1.1 Machine Learning in Additive Manufacturing

Machine learning (ML), a sub-branch of artificial intelligence (AI), has seen steady
adoption and popularity in applications that require adaptive decision-making for
their algorithms. Such applications include but are not limited to, data modeling in
mathematics, decision making and logic structures in systems automation, and the
focus of this study, defect detection in additive manufacturing (AM) [1]. The strength
of ML in real-world applications is that it allows the software to intelligently interpret
data from connected hardware and telemetry such as sensors, cameras, and radars.
AM, known also as 3D printing or rapid prototyping (RP) [2], is based on the prin-
ciple of sequentially layering material filament according to a computer-aided design
(CAD) model. This is the opposite to subtractive manufacturing, currently the preva-
lent fabrication method that involves reducing a stock of material down to shape by
cutting, milling, lathing etc. Research at the intersection of these two technologies has
become popular in the past 5 years, with the focus being directed towards improving
part quality, minimizing operational and material costs, and optimizing the fabrica-
tion process [3]. A key step in achieving these outcomes is ensuring components can
be reliably produced, known as printability [4]. This is what ML integration aims to
do, using both printer and camera data to automatically predict and diagnose sources
of error.

1.2 Applications of AM Defect Detection

With the rapid adoption and prevalence of AM, a method of ensuring consistency
and defect detection in 3D-printed parts is, however, yet to be universally adopted
[5]. This is a major challenge because it is not only highly dependent on the printing
parameters of the model, but also the machine operator’s expertise and the reliability
of the printer itself. Primary among the failure modes in 3D printing is filament
entanglement, also known as spaghettification, warping, and surface defects after
the part has been fabricated. Ideally such errors are mitigated during manufacturing,
and indeed endeavors have been made with the use of live print camera-feed (in-situ)
integration [6], but there is still no doubt that the value lies in informing smarter part
designs or operational methods from failure diagnosis data.
ML in 3D-printing applications to date has emphasized process monitoring, and
there has been comparatively little emphasis on post-fabrication quality assurance.
Recognizing this significant inconsistency in AM technologies, many adopters have
begun integrating image-recognition ML into existing models or designing new
models with this technology pre-installed. A popular method of using ML to classify
these defect patterns is with a convolutional neural network (CNN) [7]. CNNs are a
class of artificial neural network and are structured similar to neurons in the brain,
Machine Vision-Based Scanning Strategy for Defect Detection … 273

with a connectivity pattern that allows each node in the network to respond to signals
only from other nodes that are relevant. Together with an automated part-scanning
algorithm, an ‘ex-situ’ defect detection method has the potential to be integrated in
place of the traditional in-situ model.
This was the motivation behind our study—how to leverage the power of
programmable robotics, and continually improve a ML algorithm to automatically
detect and classify defects in 3D-printed parts. Addressing this fundamental draw-
back in AM would allow future research pathways to consider the integration of not
only the ML framework in classifying faults during the manufacturing process itself,
but also to use the data to self-diagnose and improve the manufacturing process for
future iterations of the model.

2 Literature Review

Significant progress has been made in introducing AM technologies to the consumer

market over the past decade. The rapidly increasing userbase of 3D-printing tech-
nology, the growth and expansion of global manufacturing demand, and the push for
efficiency and automation have increased the urgency for engineers and researchers
to recognize and consolidate a solution for mitigating defects in printed parts. Our
literature review confirmed that defect detection in AM is a problem that goes
beyond small-scale production, so we present specific and general solutions, and
conclude that a fundamental method of reducing 3D-printing defects is needed for
the continued development and integration of AM worldwide.

2.1 Defining the Problem

It has been well recognized that defects and failure caused by 3D-printing processes
are a critical barrier to the widespread adoption of AM and that it deserves serious
attention. Recently in 2020, Wang et al. [8] comprehensively reviewed the state-
of-the-art of ML applications in a variety of domains to study the main use-cases
in the research and development of AM. Those authors state that although current
ML in AM research is intensively concentrated on print parameter optimization
and in-process monitoring, there is an expectation that ML research efforts will be
directed to more rational manufacturing plans and automated feedback systems for
AM. This is a vital step in pushing ‘smart’ AM into the near future. Their article
highlighted the lack of reports on ML usage in AM with regard to determination
of microstructure, material property prediction and topology optimization, and they
conclude that further research into these areas is crucial in determining printability,
the design of ML algorithms, and optimal part designs with respect to materials.
A similar state-of-the-art review performed by Oleff et al. [5] focused on the
challenge of industrializing AM by analyzing current monitoring trends to identify
274 S. Zhang et al.

the key weaknesses. These authors found that although the amount of research into
anomaly detection evaluation was mostly for all in-situ process monitoring tech-
niques, the emphasis was decidedly on part geometries and surface properties in
terms of over- and underfill. These parameters refer to an ‘infill’ percentage, which
simply dictates the density of the printed part. Oleff et al. noted that measurements
of surface roughness and quality, as well as mechanical properties such as tensile
strength, were addressed by only two of the 221 relevant publications from their
research. They conclude that inspecting such material characteristics was simply not
considered in any of the publications, and thus represents a significant area of further
investigation.
Researchers have begun to study the implications of implementing defect detec-
tion in AM. For example, Chen et al. [3] examined the future of surface defect
detection methods by comparing the key AM part inspection technologies through
testing of various “traditional” detection methods, such as infrared imaging and Eddy
current testing, as well as ML-based techniques such as CNN image recognition and
auto-encode networks. Auto-encode networks proved to be effective at learning fault
features, rather than classifying them, but required consistent input and output data
dimensions. It was highlighted that ML defect detection is directly based on the
characteristics of the AM field, and as such is the most sustainable program moving
into the future. The authors conclude that, although more effective than traditional
inspection technologies, NN-based methods of defect detection ultimately are deeply
data-driven and establishing a universal model applicable at scale requires further
study into each of their respective advantages and disadvantages.
Yao’s findings were replicated in China by Qi et al. [9], who evaluated the effec-
tiveness of different NN structures with the intention of optimizing the performance
of AM parts. Qi et al. document the limitations in 3D-printing performance when
applying numerical and analytical models to AM. Their paper reports that not only
is a ML approach to AM valid due to its ability to perform complex pattern recogni-
tion without solving physical models, but also that NN are effective in CAD model
design, in-situ modeling, and quality evaluation. The validity and potential of linking
AM and NN technologies is brought to light, whereby they can be effectively inte-
grated from design phases to post-treatment; however, the authors conclude that key
challenges remain in the area of ML data collection and AM quality control.

2.2 Searching for Solutions

Petsiuk and Pearce [6] took a hybrid analytical approach to developing a 3D-printing
defect detection method in their study aimed at supporting intelligent error-correction
in AM. In their paper they present a comparison model, whereby an in-situ printing
process was photographed layer-by-layer and compared with idealized reference
images rendered by a physics engine, “Blender”. They found that a similarity compar-
ison did not introduce significant operational delays but rather demanded time for
its virtual environment and rendering processes. A notable strength of this approach
Machine Vision-Based Scanning Strategy for Defect Detection … 275

is that similarity and failure thresholds can be fine-tuned, providing both flexibility
and varying sensitivity to defects during manufacturing. Emphasis is placed by the
two authors on their model’s ability to scale exponentially with the number of parts
manufactured, needing only to render a base image set per part, as well as the model’s
independence of training data, presenting a significant reduction in resources required
to implement the method into AM under unfavorable printing conditions, such as
miscalibrated printer components, contamination or leveling discrepancies, which
are prevalent control issues in mainstream AM markets.
Similarly, Paraskevoudis et al. [10] outlined a method of assessing the quality of in-
situ 3D-printing using an AI-based computer vision system. Through analyzing live
video of the process by utilizing a deep CNN, a larger variant of a CNN, a primary
mode of AM defect was determined to be “stringing”, which is caused by excess
extruded material from the printer nozzle forming irregular protrusions on the printed
surface, often causing issues when dimension tolerances are critical. The deep CNN
model was demonstrably effective at identifying stringing from the experimental
data, but proved to be unreliable when applied on external data acquired from web-
based sources. This identified a critical need for continuous model improvement
through training on “new” data. Those authors comment on the further applications
of this form of error-detection, stating that, provided a sustainable detection model is
found for other key AM failure modes, this approach may be developed to adjusting
the printing process itself, whether by correcting its parameters or by terminating the
process itself. Such a feature would reduce the skill ceiling needed to operate such
machinery, meaning fewer engineers and more technicians in practice.
These international findings were replicated in Singapore in a 2021 study by
Goh et al. [11], who explored and summarized the various types of ML tech-
niques, alongside their current use in various AM aspects while elaborating on the
current challenges being faced. Through the perspective of in-situ monitoring of
3D-printing processes, Goh et al. found the high computational cost and large-scale
data acquisition for ML training to be significant challenges in practice. They also
found that CNNs better capture spatial features and are hence ideal for 2D image
and 3D model applications. The authors similarly stress the importance of large
datasets for achieving high detection accuracy, and state that the realization of predic-
tive modeling in digital twins for AM ultimately depends on the ML algorithm’s
classification accuracy, its input data quality, and its multi-task learning capabilities.
In most real-world 3D-printing applications, longevity AM as a reliable method
of part production hinges on its ability to diagnose issues with print quality. A study
conducted by Meng et al. [12] reviewed parameter optimization and discrepancy
detection in the AM field by comparing the performance of common ML algorithms.
They documented an iterative training method for an ML model wherein data from
printing parameters, in-situ images and telemetry, microstructural defects and rough-
ness, and the part’s geometric deviation and mechanical strength were used to make
predicative surrogate models to assist in-process optimization. This methodology
is known as “active learning”, whereby input–output pairs of data are formed and
used to train ML models without querying new labeling data. Practically speaking,
this significantly alleviates the cost, time, and human labor of conducting dedicated
276 S. Zhang et al.

experiments with input labeling. The authors emphasize the gap in research with
regard to active-learning ML algorithms in AM applications and conclude by stating
that such an algorithm would be highly efficient in cases where a dataset is yet to be
acquired.
Regardless of the defect analysis technique used, it has become clear that the best
outcomes are achieved with a form of ML NN that is both accurate and scalable and
able to adapt its behavior according to previous performance. Han et al. [7] considered
the direct application of ML in defect classification in AM to assess its viability in the
context of real-world image datasets. Through a process of localized segmentation of
defects across image frames taken from in-situ print monitoring, a steady-state CNN
model was established from 103 verification data sets. This model then accurately
predicted defects from 101 “new” input images of 3D-printed parts. In practice,
the authors’ model outperformed several established detection frameworks such as
“Faster RCNN” and “SSD ResNet” in terms of average precision, but lacked in terms
of detection speed. The consistency of the segmentation model can be attributed to
just that—the process of image instance segmentation effectively distinguishes the
differences and boundaries between similar surface defects, which is an approach
that can be used to help develop a similar strategy on a smaller scale.

2.3 Preparing for Tomorrow, Today

Today’s global manufacturing industry continues to increase in size and pursue more
time- and cost-effective means of production in AM. Unfortunately, research into
“smart” AM, whereby sources of print failure are determined and used to inform
better printing performance, has so far been both general and sparse. The literature
reviewed by us reflected that void in knowledge. Particularly needed is a way to cate-
gorize the defects caused by erroneous printing parameters and to develop effective
defect recognition in AM to address them.

3 Methods

3.1 Methodological Approach

Our goal was to evaluate the efficacy of using ML to automatically classify 3D-
printing defects. We aimed to both describe the characteristic strengths and weak-
nesses of using such a method and gain a more in-depth understanding about its
feasibility in the context of current 3D-printing applications.
We anticipated that a combination of quantitative and qualitative data would be
required to achieve this aim; namely, photographic images and their derivatives such
Machine Vision-Based Scanning Strategy for Defect Detection … 277

as those described by Petsiuk and Pearce [6], and Han et al. [7], to examine, decom-
pose and categorize them as input data for the CNN model. The datasets would
comprise these two data types, as the images would be parsed by a computer model
and then verified by a human.
In order to maintain a control environment for testing the ML model’s behavior, all
preliminary data used to construct and train were primary data. From the experiments
carried out by Han et al. [7] in similar CNN models, it was apparent that such a
model may be very sensitive to slight deviations in input data, and so secondary
data were omitted from our study but remains as a clear extension to model testing
in later build iterations. Furthermore, as an act of simplifying the scope of testing,
the main geometries chosen for the 3D-printed parts to be scanned were square-
or rectangular-based in nature. By limiting the sample space to this particular set of
shapes, the algorithm for rotating and capturing the surfaces of the part was drastically
simplified. This limitation, however, did not affect the validity of what was being
tested, but rather left open the door to research into converting this method into a
universal model.
Finally, as a direct comparison with the ex-situ model, was created to both aid in
accelerating the rate of data collection and act as a benchmark for performance and
accuracy.

3.2 Data Collection Methods

Data were collected after establishing the CNN model in YOLOv5, whereby samples
of 3D parts printed locally in the Monash Smart Manufacturing Hub Staging Lab
served as a proportion of the total part sample space. The main selection criteria
for these samples were their defect type and locality relative to the part’s geometry,
and were selected to assist in faster pattern recognition training and create a fair
comparison between the in-situ and ex-situ models.
The relevant tools, equipment, and materials used to gather data were a Creality
Ender 3 v2 3D printer, PLA printer filament (colored and white PLA +; 1.75 mm
diameter), JAI HD-camera and Windows-based operating system (PyTorch IDE and
LabelImg software).
Images were captured in a controlled environment to ensure consistency not only
within the scope of the experiment itself, but also to the degree that could be replicated
by another researcher or engineer (Fig. 1). This environment comprised a fixed,
white-balanced background calibrated against a control in equal lighting conditions
(brightness and color-temperature).
We anticipated encountering issues with object visibility when the color of the part
being examined was identical to the reference background. In such cases, a unique
background color was to be fixed or alternate backgrounds exchanged. Another
potential issue was the base dataset not being sufficiently large. As mentioned by
Chen et al. [3] and Goh et al. [11], a major limitation in mass-integration of NN-based
models in AM is the sheer size of training dataset required. It is necessary to consider
278 S. Zhang et al.

Fig. 1 Preliminary data

collection setup: KUKA
arm, gripper, JAI camera rig

the effect of this when evaluating the performance of ML in this application. It was
only feasible for us to attain < 200 samples of data for use in training and validation.

3.2.1 KUKA Robotic Arm Algorithm

The arm was programmed to grab, move, and rotate test pieces to allow the mounted
JAI camera to capture multiple isometric views of the part (Figs. 2, 3, 4 and 5).
Specifically, two external isometric images of the part were captured. Based on the
initial hypothesis, this method reduced the number of images needed to train the
CNN for a particular defect, and thus reduced the overall duration of scanning. The
advantage of having a programmable robotic arm is the ability to maneuvre the piece
to expose every surface in just a few images, each captured at an isometric angle
containing multiple faces.
Conditions surrounding the test rig were kept constant in terms of lighting (cool
white light of 5300 K) and background (solid background sheet behind and below
the sample) to limit the effect of extraneous variables.

Fig. 2 Example testing part:

white hollow cube
measuring 80 × 80 × 80 mm
Machine Vision-Based Scanning Strategy for Defect Detection … 279

Fig. 3 Robotic arm in first

position: working part is
located and lifted

Fig. 4 Robotic arm in

second position: working
part is rotated to first
isometric angle relative to
hanging JAI camera

Fig. 5 Robotic arm in third

position: working part is
rotated to second isometric
angle relative to hanging JAI
camera
280 S. Zhang et al.

Further stages of development are: calibration pieces (homing, reference plane,

improved consistency in both programming and in result collection); hanging camera
mounting at isometric angle; and surrounding part background.

3.2.2 YOLO Image Classification

Using the isometric images gathered with our scanning method, the YOLOv5 model
was trained under human supervision, which involved digitally labeling the input
data with the LabelImg software (Fig. 6). These annotations are the mechanisms
through which the CNN learns to recognize defect features.
An equivalent in-situ model provided image data that were also used as training
data. Further improvements to training methodology are using a 1:9 ratio of true
negative to true positive training data because 10% of image training data would not
contain the corresponding defects and supplementing primary data with open-source
secondary datasets. As a CNN model’s performance improves proportionally to the
volume of training input, it is expected that this will improve the overall performance.
Difficulties included larger forms of warping not being easily distinguished from
curved features of the part’s geometry, which restricted the effectiveness of the
preliminary model to only similar cuboidal shapes. To mitigate this during labeling,
larger examples of these features were bisected for better localization (Fig. 7).

Fig. 6 Example of labeling

of primary defect modes
using LabelImg: layer shift,
stringing and warping
Machine Vision-Based Scanning Strategy for Defect Detection … 281

Fig. 7 Bisection method of

larger examples of defects.
This particular example
shows dividing a large
warping defect spanning the
entire width of the part into
two separate instances

3.3 Methods of Analysis

In both the quantitative and qualitative data analyses, outliers and anomalies were
factored into the assessment of the model’s functionality and reliability, but will be
omitted from computational methods such as being used as training input.
The main method of quantitative analysis was the passing of image data through
a CNN and recording the corresponding output. This network functions as a filtering
device, using a network of calibrated nodes that respond to their associated shape or
pattern. The nodes make up the decision-making mechanism of the CNN and can be
adjusted during experimentation.
Qualitative analysis will comprise identification, verification and classification
against the following main AM print failure categories [10]:
– spaghettification/stringing
– layer shift
– warping.
These categories can then be linked with associated mechanical errors or defects:
– under-extrusion
– over-extrusion
– nozzle blob
– poor initial layer
– poor bridging overhang
– premature detachment of print
– no extrusion
– skirt issues.
Human verification of these defect features will be used to train the CNN.
282 S. Zhang et al.

3.4 Justifications

An immediate limitation of a camera-based setup is detecting printing defects that

are internal to the structure of the printed part. If a 3D-printed component needs
to pass a quality check under the method outlined here, there would be no way of
verifying the integrity of the inside surfaces and whether they fulfill certain failure
mode criteria. An example of this would be a hollow cube, in which there is an
interior surface that contributes to the physical properties of the part such as density,
rigidity etc. Although it is possible to reveal print defects using higher energy such
as X-ray [10], it would ultimately be more reliable and cost-efficient to simply detect
such major print errors during the manufacturing process itself, such as was explored
by Vosniakos et al. [2], Petsiuk and Pearce [6] and Oleff et al. [5].

4 Results and Discussion

4.1 Preliminary Results and Discussion

4.1.1 Deliverables

Our experimentation aimed to deliver a simple, working demonstration of analyzing

a 3D-printed part for surface defects. The process achieved the following: (a) part
retrieval from a dedicated storage location using a KUKA robotic arm, (b) an auto-
mated algorithm for part scanning using a high-definition JAI camera, (c) classifica-
tion of surface defects if present and (d) percentage accuracy rating in comparison
with an equivalent in-situ model.
At the time of writing, initial findings and results have only been taken from
training data and experiments with the in-situ testing rig.
YOLOv5 worked consistently, but is yet to prove a reliable ML solution,
noticeably because of the sheer volume of training input required.
The gathering of data will be a significant priority in the future stages of this study,
with measures already taken to divide the printing load across the ex-situ and in-situ
models. Further, an additional two Ender 3 v.2 rigs will be established external to
the laboratory for simultaneous data acquisition.
An initial accuracy of 18.9% from the accompanying in-situ rig was due to
insufficient training data with the model.
The preliminary KUKA robotic arm scanning algorithm based on isometric
images was effective only for square-based prisms in a sample of existing test print
pieces.
Machine Vision-Based Scanning Strategy for Defect Detection … 283

4.2 Areas of Future Work

Although this study does tackle a prominent drawback with AM technology, it is

considerably limited in scope, particularly for adapting this methodology to parts of
any size or shape.

4.3 Study Limitations

– The KUKA robotic arm programming is currently limited to cuboidal and

rectangular-shaped parts. This is only a limitation of the scanning algorithm itself
however, and once addressed by an adaptable part maneuvering strategy will
present no significant challenge to training the CNN.
– The ability for this model to detect and classify minute defects or other small
tolerancing issues may depend greatly on the resolution of the attached camera
equipment. A solution for this is simply to upgrade to a higher resolution camera
befitting for the tolerance requirements.
– Gathering large samples of verified training data is an unavoidable hurdle in
assuring a high rate of accuracy for ML-based models such as YOLOv5. This
will be a significant contributor to the cost and time investment for this model to
be adopted at scale.
– Although defects can be quite reliably identified through the method outlined in
this study, an ideal method of representing this data as a useful diagnosis for users
was ultimately not investigated due to limited working constraints.
– The severity of defects was not considered. Although the type of defect may be
successfully discerned, neither its degree nor magnitude can be quantified.
– The in-situ experiment running parallel to this method captured footage on
different image-capturing hardware. Although this did not affect the premise of
this study to a great extent, as in-situ monitoring remains limited by its fixed angle,
nevertheless a closer comparison should be made using identical equipment.

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the copyright holder.
Electrical and Sulfate-Sensing Properties
of Alkali-Activated Nanocomposites

Maliheh Davoodabadi, Marco Liebscher, Massimo Sgarzi,

Leif Riemenschneider, Daniel Wolf, Silke Hampel, Gianaurelio Cuniberti,
and Viktor Mechtcherine

Abstract We investigated the formation of the conductive network of carbon

nanotubes (CNTs) in alkali-activated nanocomposites for sulfate-sensing applica-
tions. The matrix was a one-part blend of fly ash and ground granulated blast-furnace
slag, activated by sodium silicate and water. Sodium dodecylbenzenesulfonate was
used as the surfactant for dispersion of the CNTs in the aqueous media. The nanocom-
posites were investigated by a laboratory-developed setup to study the electrical and
sensing properties of the alkali-activated material. The electrical properties (i.e.,
conductivity) were calculated and assessed to discover the percolation threshold
of the nanocomposites. Furthermore, the sensing behavior of nanocomposites was
studied upon sulfate (SO2− 4 ) exposure by introduction of sulfuric acid ((H2 SO4 ))
and magnesium sulfate (MgSO4 ). The sensors were able to preliminarily exhibit a
signal difference based on the introduced media (H2 SO4 &MgSO4 ), CNT content

M. Davoodabadi · M. Liebscher (B) · V. Mechtcherine

Institute of Construction Materials, Faculty of Civil Engineering, Technische Universität Dresden,
Dresden, Germany
e-mail: marco.liebscher@tu-dresden.de
M. Davoodabadi · M. Sgarzi (B) · L. Riemenschneider · G. Cuniberti
Institute for Materials Science and Max Bergmann Centre of Biomaterials, Dresden University of
Technology, Dresden, Germany
e-mail: massimo.sgarzi@unive.it
Dresden Center for Nanoanalysis (DCN), Dresden University of Technology, Dresden, Germany
M. Sgarzi
Department of Molecular Sciences and Nanosystems, Ca’ Foscari University of Venice, Venice,
Italy
D. Wolf · S. Hampel
Leibniz Institute for Solid State and Materials Research, Dresden, Germany
G. Cuniberti
Center for Advancing Electronics Dresden (CfAED), Dresden University of Technology, Dresden,
Germany
Dresden Center for Computational Materials Science (DCMS), Dresden University of
Technology, Dresden, Germany

© The Author(s) 2023 285

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_29
286 M. Davoodabadi et al.

and H2 SO4 volumetric quantity. The results of this research demonstrated a sensing
potential of CNT alkali-activated nanocomposites and can be applied in the concrete
structural health monitoring.

Keywords Aggressive ion sensing · Alkali-activated nanocomposites · Carbon

nanotubes · Structural health monitoring

1 Introduction

Aggressive ions such as sulfates (SO2−4 ) are naturally available in the surrounding
aquatic media of concrete infrastructures. Moreover, these anionic species can be
liberated from the underground environments or internal sources by moisture ingress
(chemical origin) or can be produced through the metabolism of microorganisms
(biogenic origin). The main sources of SO2− 4 anions are sulfuric acid (H2 SO4 ),
magnesium sulfate (MgSO4 ), and sodium sulfate (Na2 SO4 ), which cause severe
degradation of concrete material over long-term exposure [1–4]. An innovative idea
regarding prevention of concrete material deterioration is the development of a sensor
to distinguish SO2−
4 species coming from identical [5] or non-identical sources or
having different concentrations/quantities.
Therefore, in the present study, carbon nanotube (CNT) alkali-activated
sensors are proposed for SO2− 4 species sensing and discriminating released from
H2 SO4 &MgSO4 , for the first time to the best of our knowledge. The sensors were
fabricated from a sodium-based fly ash ground granulated blast-furnace slag (GGBS)
alkali-activated material and CNTs. As the first step, the percolation threshold of the
sensors was determined by measuring the electrical resistance (converted to conduc-
tivity) of CNT alkali-activated nanocomposites, incorporating different content of
CNTs. To study the SO2−4 sensing potential and transitional behavior of the nanocom-
posites from insulating to conducting mode, the sensors were fabricated to incor-
porate different CNT concentrations. The assessments were carried out by intro-
ducing H2 SO4 &MgSO4 with identical concentration. Finally, the sensors were fabri-
cated with the percolated CNT concentration and evaluated by introducing different
volumetric quantities of H2 SO4 .

2 Methods

2.1 Materials

The applied CNTs were TUBALL™, supplied by OCSIAL Europe and their prop-
erties can be found in Davoodabadi et al. [5, 6]. The surfactant was a technical
grade of sodium dodecylbenzenesulfonate (SDBS) produced by Merck KGaA. The
Electrical and Sulfate-Sensing Properties of Alkali-Activated … 287

utilized precursors were fly ash (Steag Power Minerals GmbH) and GGBS (Opterra
GmbH); their properties are described in Davoodabadi et al. [7]. Sodium disilicate
powder (Sikalon; Wöllner GmbH) was used to activate and geopolymerize the blend.
Sulfuric acid and magnesium sulfate for the sensing measurements were diluted to
0.1 M from a stock solution of 98% ACS reagent sulfuric acid (Merck KGaA), and
anhydrous magnesium sulfate powder (Merck KGaA), respectively.

2.2 Methods

The nanofluids were prepared by ultrasonication of the CNTs and SDBS in ultrapure
water according to the procedures in Davoodabadi et al. [5–7]. The CNT concentra-
tion range spanned from 0.010 to 1.000 wt% of oxide mass and SDBS was added with
the same mass as CNTs. The nanocomposites were fabricated with the formulation
in Davoodabadi et al. [5–7]. The mixed slurries were cast into plastic molds with slot
dimensions of 60 × 10 × 10 mm3 for 24 h. After demolding, the nanocomposites
were cured in the chemical laboratory ambient conditions. Thereafter, the nanocom-
posites were heat treated (at 105 °C for 24 h) to eliminate any negative impact of
water on the CNTs’ conductive network.

2.3 Characterizations

A programmed setup was used for the electrical properties and sensing measure-
ments composed of a KEYSIGHT B2912A precision source/measure unit (SMU),
a computer, and Grafana time-series database [5]. The received data in the form of
the resistance (R) in Ω were converted to resistivity (ρ = R.A.L−1 ) in Ω.m and
conductivity (σ = ρ−1 ) in S.m−1 by applying the sensors’ cross-sectional area (A)
and length (L). For further analysis, relative resistance (RR = 100.(R1 − R0 ).R0−1 )
was used for sensor evaluations. A GeminiSEM 500 (Carl Zeiss QEC GmbH) was
used for scanning electron microscopy imaging of the specimens’ cross section.
The FEI Tecnai F30 (ThermoFisher Scientific) was used to conduct high resolution
transmission electron microscopy (HRTEM imaging) of the samples.
288 M. Davoodabadi et al.

3 Results and Discussion

3.1 Percolation Threshold

In the available literature, the sensing properties of geopolymers exploited the elec-
trolytic (ionic) conductance of the composites because of their ion-rich structure and
pore system [8–12]. However, investigated alkali-activated composites in this study
are insulators and exhibit insufficient conductive (electronic) properties to act as a
sensor. The measured inherent resistances of these composites were unsteady, in the
range of mega ohm (≥10 MΩ). Because measurements are conducted on heat-treated
specimens, the electrolytic conductivity of the ion-rich framework is assumed to be
negligible. Sufficient inclusion of CNTs causes the resistance range of nanocom-
posites to descend abruptly from mega ohm (highly insulator character) to hundreds
of ohms (highly conductive behavior). This proves that the electronic conductive
network of CNTs has been generated, and the nanocomposites are percolated [13].
The correlation curve of the nanocomposites’ conductivity and CNTs’ concentra-
tion is depicted in Fig. 1. The observed increasing trend of conductivity corresponded
to the required quantity of CNTs for establishing a functional conductive network,
mostly in tube-contacting mode rather than the tunneling or hopping mode [13]. This
conductive network of CNTs provides the alkali-activated matrix with a percolating
transition zone between 0.070 wt% and 0.200 wt% as indicated by the inset in Fig. 1.
With respect to percolation theory, such curves can be fitted with power regres-
sion models [14–17]. In addition to the investigated nanocomposites, CNT Portland
cement-based nanocomposites have a relatively similar percolation trend, but their
documented thresholds exhibit, naturally, non-conclusive values. Some of the docu-
mented percolating transition zones of CNT Portland cement-based nanocomposites
are shown in Table 1 for comparison.

Fig. 1 Percolation diagram

of CNT alkali-activated
nanocomposites
(concentrations (wt%)
between 0 and 0.25 are
0.010, 0.025, 0.050, 0.075, &
0.100)
Electrical and Sulfate-Sensing Properties of Alkali-Activated … 289

Table 1 Reported
Reference Liu et al. [24] Danoglidis Hong et al. [26]
percolating transition zones
et al. [25]
of CNT cementitious
nanocomposites Range 0.80–1.60 vol% 0.10–0.15 0.30–0.60 wt%
wt%

The most influential parameters on the percolation of nanocomposites are the

CNTs’ structure (i.e., chirality, aspect ratio, waviness, and the number of shells) and
CNTs’ tunneling resistance/distance. The fabrication methodology (applied surfac-
tant, ultrasonication energy, mixing, and curing), which determines the 3D orienta-
tion, configuration, interconnection of CNTs, and geometry of agglomerates, further
affects the percolating character of the nanocomposite [15, 16, 18–21]. Considering
the percolation threshold as an outset, the preference of composite researchers is to
maintain the nano-additives’ concentration as low as possible (i.e., low or ultra-
low percolation thresholds), because high concentrations of nano-additives have
a destructive impact on the microstructure and consequently on the mechanical
properties of the nanocomposite [14, 22, 23] (Table 1).
Based on these arguments, to achieve a firm conductive network of CNTs, in
which the nanocomposite is not so overloaded as to be destroyed by the additives
and not too under-loaded to be an insulator, a concentration of 0.1 wt% was consid-
ered for the sensor fabrication. In Sect. 3.2. Sulfate discrimination, the selection of
CNT concentration of 0.1 wt% as the percolation threshold is further validated. In
comparison, a wide concentration range of 0.05–2 wt% was reported in the literature
for sensing and smart applications of the Portland cement-based nanocomposites
[27–34].

3.2 Sulfate Discrimination

The CNT alkali-activated nanocomposites sensing responses upon exposure to 90 μL

(30 μL in each cycle) of 0.1 M H2 S O 4 & MgS O 4 with CNT concentration ranging
from 0.0 to 1.0 wt% are illustrated in Fig. 2. The response of the sensors upon exposure
to different sulfate containing regimens was explored and compared. A range of CNT
concentrations were considered to confirm the influence of the generating conductive
percolated network of CNTs on the sensing behavior of the nanocomposites with
respect to percolation analysis.
The nanocomposites with incorporation of 0.0 and 0.010 wt% exhibited fully
insulating behavior (Fig. 2a, b). The responses were highly noisy and stochastic as
expected from the ultra-low conductivity of these nanocomposites; the approximate
magnitude was 3E-5 S·m–1 as shown in Fig. 1. With CNT incorporation of 0.025
wt% and 0.050 wt%, the conductivity rose sharply to almost 3E-2 and 1 S·m–1 ,
respectively, and the responses were slightly functional but still not accurate (Fig. 2c,
d). Considering the onset of the percolating area, at CNT concentration of 0.070 wt%
290 M. Davoodabadi et al.

a 100
SWCNT Conc. 0.000 wt.%
H2SO4
b SWCNT Conc. 0.010 wt.%
H2SO4
c SWCNT Conc. 0.025 wt.%
H2SO4
10 0
MgSO4 MgSO4 MgSO4
80

Relative Resistance (%)

Relative Resistance (%)
Relative Resistance (%)

0 -5
60

40 -10 -10

20
-20 -15

0
-30 -20
-20

-40 -40 -25

0 60 120 180 240 300 360 420 480 540 600 660 0 60 120 180 240 300 360 420 480 540 600 660 0 60 120 180 240 300 360 420 480 540 600 660
Time (s) Time (s) Time (s)

d SWCNT Conc. 0.050 wt.%

e SWCNT Conc. 0.075 wt.% f SWCNT Conc. 0.100 wt.%
H2SO4 H2SO4 H2SO4
4 4
2 MgSO4 MgSO4 MgSO4
Relative Resistance (%)

Relative Resistance (%)

Relative Resistance (%)

3 3
1

2 2

0
1 1

-1
0 0

0 60 120 180 240 300 360 420 480 540 600 660
0 60 120 180 240 300 360 420 480 540 600 660 0 60 120 180 240 300 360 420 480 540 600 660
Time (s)
Time (s) Time (s)

g SWCNT Conc. 0.250 wt.%

h SWCNT Conc. 0.500 wt.% i SWCNT Conc. 0.750 wt.%
H2SO4 H2SO4 H2SO4
4 4 4
MgSO4 MgSO4 MgSO4
Relative Resistance (%)

Relative Resistance (%)

Relative Resistance (%)

3 3 3

2 2 2

1 1 1

0 0 0

0 60 120 180 240 300 360 420 480 540 600 660 0 60 120 180 240 300 360 420 480 540 600 660 0 60 120 180 240 300 360 420 480 540 600 660
Time (s) Time (s) Time (s)

Fig. 2 The sensing responses of CNT alkali-activated nanocomposites exposed to 90 μL (30 μL

in each cycle) of 0.1 M H2 S O 4 &MgS O 4 ; CNT concentrations (wt%) in the nanocomposites are
a 0.000; b 0.010; c 0.025; d 0.050; e 0.075; f 0.100; g 0.250; h 0.500; i 0.750

the functional response of the sensors began at 0.075 wt% with sensor conductivity
of ≈6 S·m–1 (Fig. 2e). From this point onwards, the CNT signals showed a regular
configuration, which corresponded to the normal behavior of pristine p-type CNTs
(Fig. 2e–j). The H2 SO4 & MgSO4 discrimination mechanism in the percolated area
was mostly based on the difference in the signal shape and magnitude.
The correlations of maximum relative resistance and CNT concentration are
plotted in Fig. 3. The main results to note were (i) the positive relationship of the rela-
tive resistance and CNT concentration, and (ii) the higher sensitivity of the nanocom-
posites towards H2 SO4 exposure than MgSO4 . Result (ii) means the sensors can
discriminate SO2− 4 species introduced from two different sources (H2 SO4 & MgSO4 ),
by means of relative resistance as demonstrated in Fig. 3, and by signal configuration
and shape as illustrated in Fig. 2. The CNTs’ signal differentiation was particularly
recognizable in the percolating transition zone of the nanocomposites (between 0.075
and 0.250 wt%) shown in Fig. 2e–g.
Electrical and Sulfate-Sensing Properties of Alkali-Activated … 291

H2SO4
MgSO4

Maximum Relative Resistance (%)

4

0.075
0.075 0.1 0.25
0.100 0.250 0.5
0.500 0.750
0.75 1.000
1

SWCNT Concentration (wt.%)

Fig. 3 Maximum relative resistance variations of CNT alkali-activated nanocomposites based on

CNT concentrations upon exposure to 90 μL of 0.1 M H2 S O 4 &MgS O 4

3.3 Quantity Differentiation

The H2 SO4 quantity differentiation potential of the sensors is shown in Fig. 4. For
this purpose, the nanocomposites were fabricated with CNT inclusion of 0.1 wt%
based on the percolation and sensing analyses. The same concentration regimen of
H2 SO4 (i.e., 0.1 M) was introduced in volumetric quantities of 90, 180, 270 μL (30,
60, and 90 μL, respectively in each cycle) to the sensors to evaluate the responses.
The increments of relative resistance were approximately + 140% and + 60% with
the volume increasing from 90 μL to 180 μL and afterwards to 270 μL, exhibiting
a linear behavior.

3.4 CNTs’ Conductive Network

The CNTs’ network distribution and expansion are shown in Fig. 5. The addition
of SDBS as surfactant for the dispersion of CNTs, entrained air bubbles into the
alkali-activated microstructure. In addition, SBDS probably had a negative impact
on the activating reactions and created a weaker microstructure. Nevertheless, the
alkali-activated microstructure exhibited sufficient strength (compressive strength
of 40 ± 2.48 MPa for 28-day nanocomposite) and mechanical performance. The
main mechanism was the reinforcing effect of the CNTs because of their distribu-
tion (Fig. 5a), and crack covering and bridging abilities (Fig. 5b). Furthermore, this
mode of interaction by CNTs has a secondary function, which is the formation of a
conductive network (Fig. 5c). This network endows the nanocomposites with high
292 M. Davoodabadi et al.

Fig. 4 Quantity differentiation of CNT alkali-activated sensors (CNT conc. 0.1 wt%). a Exposure
to different volumes of 0.1 M H2 S O 4 ; b regression of relative resistance vs. volumetric quantity

electrical conductivity and consequently a sensing ability as explained in previously.

The percolated conductive network of CNTs can be seen in the HRTEM images
of Fig. 6. The network comprised overlapping CNT agglomerations, which formed
directional pathways.

4 Conclusions

The present conceptual study has proposed a structural sensor for assessing the SO2−
4
sensing potential of CNT alkali-activated nanocomposites. The investigated discrim-
ination criteria were CNT concentration, SO2− 4 bearing media (H2 SO4 vs.MgSO4 ,
and analyte volumetric quantity H2 SO4 . The obtained results can be summarized as
follows.
• The percolating zone of the nanocomposites was between 0.07 and 0.20 wt% of
CNT content.
• The sensors can be fabricated by incorporation of 0.1 wt% of CNT into the
alkali-activated matrix material based on the percolation threshold study of the
nanocomposites.
• The sensors exhibit differentiation behavior by variation of shape and magnitude
of the obtained relative resistance.
• There was a linear correlation between relative resistance and CNT concentration
when the sensors were exposed to 0.1 M H2 SO4 &MgSO4 . Similarly, the rela-
tionship between the relative resistance of the sensors and volumetric quantity of
0.1 M H2 SO4 was linear.
• The sensors exhibited a higher magnitude of relative resistance when exposed to
0.1 M H2 SO4 compared with 0.1 M MgSO4 .
Electrical and Sulfate-Sensing Properties of Alkali-Activated … 293

Fig. 5 The reinforcing and conductive network of CNTs in the microstructure of the alkali-activated
matrix: a CNT distribution, b crack covering and bridging by CNTs, c CNTs’ network expansion

• The percolated sensors presented a response curvilinear shape upon H2 SO4

exposure and a rectangular shape upon MgSO4 exposure.
294 M. Davoodabadi et al.

Fig. 6 Percolated conductive network of CNTs in the nanostructure of the alkali-activated matrix
at the atomic scale of HRTEM. The percolated network of overlapped CNTs is shown as yellow
pathways

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Advances in Characterization
of Carbonation Behavior in Slag-Based
Concrete Using Nanotomography

B. Mehdizadeh, K. Vessalas, B. Ben, A. Castel, S. Deilami, and H. Asadi

Abstract Exposure of concrete to the atmosphere causes absorption of CO2 and

carbonation via a chemical reaction between the CO2 and calcium hydroxide and
calcium-silicate-hydrate reaction products inside the concrete. A greater under-
standing of carbonation behavior and its micro- and nanoscale impacts is needed
to predict and model concrete durability, cracking potential and steel depassivation
behaviors. New and sophisticated techniques have emerged to analyze the microstruc-
tural behavior of concrete subjected to carbonation. High-resolution full-field X-ray
imaging is providing new insights to nanoscale behavior. Full-field nano-images
provide significant insight into 3D structural identification and mapping. Nanoto-
mographic modeling of an accelerated carbonated test specimen can also provide a
3D view of the pore structure that resides inside slag-based concrete. This is crit-
ical for better understanding of the capillary porosity and pore solution behaviors
of concrete in situ. We investigated the analysis of durability properties, including
the carbonation behavior of slag-based concrete, by evaluating microstructural and
nanotomographic identification techniques.

Keywords Accelerated carbonated test · Nanotomography · Slag-based concrete ·

X-ray imaging

1 Introduction

The possibility of achieving high strength and durability has made concrete a well-
established structural material for use in roads, tunnels and bridge infrastructure [1–
4]. Various studies have proposed alternative materials (supplementary cementitious

B. Mehdizadeh (B) · K. Vessalas · A. Castel

School of Civil and Environmental Engineering Technology, University of Technology Sydney,
NSW, Australia
e-mail: bahareh.mehdizadehmiyandehi@student.uts.edu.au
B. Ben · S. Deilami · H. Asadi
Advanced Technical Services Infrastructure and Place, Transport for NSW, NSW, Australia

© The Author(s) 2023 297

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_30
298 B. Mehdizadeh et al.

materials (SCMs)) to partially replace cement to reduce the CO2 emissions caused
by its production [5–8]. Using SCMs such as fly ash, slag, silica fume, etc. as partial
replacement of cement is effective in making concrete more cost-effective, and high
performing in terms of strength and durability requirements [9–16].
The carbonation of concrete caused by the presence of CO2 in the atmosphere
(≈380 ppm) acts as an environmental load and contributes significantly to the
deterioration of concrete structures [17], because it causes the alkalinity of the
concrete to decrease (pH 8). If the pH decreases below 12, the risk of corrosion
is increased for any fixed steel components present within the concrete elements
[18, 19]. Carbonation increases the porosity of the concrete, resulting in reduced
compressive strength and impermeability in the carbonated zone of the concrete.
The results of several experimental studies have revealed that the carbonation of
slag-based concrete is highly dependent on the water/cement ratio, cement replace-
ment ratio, curing method, and in-situ environmental conditions [20–24]. Because
the carbonation process is long term in its manifestation, many researchers have used
accelerated carbonation tests by adding pressurized CO2 or increased the tempera-
ture to increase CO2 diffusivity in the pore solution to shorten the experimental time
[25–28].
Depending on the cement replacement ratio and the type of SCM, the properties
of SCM-based concrete change over time as a result of both the hydration process
and the carbonation behavior [29–31]. After carbonation, the total capillary porosity
volume of the slag-based concrete increases, which has a negative impact on the
durability of the concrete by creating a larger pore structure and thus increasing the
permeability coefficients for the concrete [32–34].
To control this behavior, nondestructive techniques such as micro- and nanoto-
mography can be used to monitor changes in the pore structure of slag-based
concretes [35]. Han et al. demonstrated how these modern technologies are useful
for analyzing the depth of concrete carbonation [36].
Recent advancements in micro- and nanoscale techniques give insight into fore-
casting and simulating concrete durability, cracking potential, and steel depassiva-
tion behaviors. We present the main advances in these techniques in investigating
carbonated concrete behavior. Although several studies were reviewed there is still
limited information published about the use of tomographic techniques for better
understanding of the carbonation behavior of slag-based concretes.

2 Tomography Techniques

Tomography, in the broadest sense, is any technique that uses sectional views as an
intermediary stage before reassembling a three-dimensional (3D) object. This char-
acterization technique is useful for identifying the richness of the microstructure in
three dimensions rather than only presenting two-dimensional projections. 3D micro-
and nanoscale views are required to fully comprehend and monitor the behavior
of the concrete. Materials scientists have used X-ray tomographic techniques for
Advances in Characterization of Carbonation Behavior in Slag-Based … 299

decades to discover the behavior of 3D microstructures [36–40]. FIB/SEM tomog-

raphy (focused-ion beam scanning electron microscopy), electron tomography, X-ray
micro-computed tomography (micro-CT), and X-ray nano-computed tomography
(nano-CT) are nondestructive 3D imaging techniques that are useful for investigating
the interior structure of a wide range of materials [34, 35, 38–41].
Tomographic techniques can give new insights into concrete deterioration by
providing precise information on the changed layers of concrete affected by carbon-
ation, corrosion, leaching or sulfate attack [42]. Carbonation as a durability concern
in concrete is strongly influenced by microstructural and pore network characteris-
tics. The porosity parameters obtained from the tomographic data can successfully
indicate the internal pore network geometry and microstructural features [37]. X-
ray nano- and micro-CT techniques are useful for monitoring slag-based concrete
behavior subjected to carbonation [43, 44].

2.1 X-ray Micro-CT Imaging

X-ray micro-CT is a radiographic imaging technique that generates a series of

cross-sectional images to identify the internal structure of materials without causing
damage to the specimen [45–47]. In general, micro-CT provides a series of recon-
structed images represented by a pixel, either 8- or 16-bit. Micro-CT processing
transforms the 8 or 16-bit images into binary or segmented images, which are useful
for examining the features of the specimen. A 3D microstructural image can be
created by stacking 2D segmented images to analyze volumetric, multi-directional,
and other advanced sample features. The pore structure of the concrete can affect
the qualities of the concrete, including strength, durability, and permeability. Using
micro-CT images can assist in discovering more information about these properties
[48–51]. Although several scales are used, such as macro-, micro- and nano-CT, the
spatial resolution of micro-CT can be used for microscopic CT scanning [45]. The
micro-scale CT technique can scan a range from 1 to 10 µm in size, which is also a
good range for investigating the size of capillary pores inside the cement paste [45,
52, 53]. For covering the whole range of cement paste pore sizes, nano-CT images
can further scan a range of 10 µm to < 10 nm [48, 54]. In order to obtain high-quality
nano-CT data it requires meticulous sample preparation. Measuring 3D morphology
at < 10 nm can give unique information on transportation characteristics and the
mechanical and durability behavior of the concrete [55].
300 B. Mehdizadeh et al.

2.2 Micro-CT and Nano-CT to Investigate Carbonated

Concrete

During the hydration process, the cement paste interacts with CO2 . Carbonated
cement paste contributes to steel reinforcement corrosion, causing major and long-
term durability issues for concrete structures. Several studies have noted that as
carbonation develops, porosity reduces due to calcite formation filling the pore
microstructure [56–59]. Phenolphthalein is a well-known technique for determining
carbonation depth in concrete, but it necessitates the destruction of the sample, and
the findings vary based on the sampling location [60, 61]. In contrast, X-ray micro-
and nano-CT are nondestructive techniques that can identify and analyze the depth
of carbonation and the microstructure of the concrete during the carbonation process
[62–67]. In addition, they can continually monitor the evolution of the carbonated
area of the same sample at different ages [68].
Some researchers have used micro-CT to determine the microstructural develop-
ment of the cement paste during the carbonation process, particularly the distribu-
tion of porosity and the effective pore width [69, 70]. In both studies, the results
of average porosities revealed that the porosity reduces with additional carbonation
time, which infers that calcite (calcium-bearing phase) forms during the carbona-
tion process, and that the porosity distribution may validate the pore microstructural
change (e.g., reduction in porosity caused by carbonation). Because the porous struc-
ture of concrete extends from the nano- to the macroscopic scale [71], X-ray micro-
and nano-CT with good resolution are in high demand for characterizing a wide range
of behavior [74]. Particle size and shape, interfacial topology, particle structure, pore
structure, carbonation depth, and morphology of distinct solid phases in concrete
have all been studied by these methods [72–76].
X-ray micro-CT results show that microcracks form from the surface to the inside
of the cement paste after carbonation. Furthermore, the carbonated area increases in
depth with increasing carbonation time. Moreover, cracks form during the carbona-
tion process and reduce the density [77]. A new generation of laboratory-based nano-
CT with high resolution [45] can provide 3D images for measuring different prop-
erties of the concrete such as durability, cracking potential, and steel depassivation
behavior. Dimensional and transitional stability is necessary for generating quality
data by any instrument involved in sub-micron imaging. It has been demonstrated
that nano-CT with high-resolution scanning is a well-established and mature method
[78] for gaining insight of the porous and hierarchical structure of the concrete at
the sub-micron scale [55]. The ability to perform nano-CT scanning under ambient
conditions keeps test samples in their natural form under normal and accelerated
conditions [55, 71, 79–83]. Han et al. [83] reported that by using nano-CT, the
solid-phase composition, pore structure, damage degradation, and nano-mechanical
characteristics of the concrete can be measured at different accelerated carbonation
ages. Their results confirmed that without any prior drying preparation, X-ray CT
is a suitable technique for obtaining 3D images of concrete to assess the degree of
microstructural damage.
Advances in Characterization of Carbonation Behavior in Slag-Based … 301

2.3 Microstructural Characteristics and Properties

of Slag-Based Concrete During Carbonation

SCMs play a critical role in controlling and improving the mechanical and durability
properties of the concrete [14–16]. Han et al. [83] described how they used micro-CT
to track the progression of carbonation-induced fractures and how the carbonation
depth increased with exposure duration for concrete containing a slag content up to
70% of the total binder content. Figure 1 shows several cracks produced during the
carbonation process, which demonstrates how micro-CT can be utilized to categorize
carbonation behavior over time [81, 83]. The micro-CT results revealed that the width
and length of microcracks significantly affects the carbonation behavior of concrete
[83]. More CO2 penetration causes an increase in crack length. When the fracture
width is < 10 µm at 1 year, CO2 diffusivity around the crack is nearly equal to that in
the surrounding concrete. However, with fracture width > 10 µm, the CO2 diffusivity
in the concrete increases. When the crack width is > 100 µm, the CO2 diffusivity is
somewhat further increased [80, 81, 83, 84].
In another study, Han et al. [67] analyzed the carbonation depth of the cement
paste with different slag addition from 0 to 70% under 0–14 days of accelerated
carbonation testing (after 3 months of curing). Because of slag’s cementitious and
pozzolanic properties, the hydration reaction of the slag-based cement will improve
the pore microstructure. Large pores will eventually transition into smaller pores with
the pozzolanic reaction, which significantly reduces the CO2 diffusion coefficient and
therefore the rate of carbonation will also decrease. However, slag-based concrete
generates a large amount of calcium hydroxide during the hydration process, and
calcium hydroxide is an important chemical component in carbonation. The results
from Han et al. demonstrated that the ideal slag addition to the binder to mitigate
carbonation is < 50%. Figure 2 shows the specimen’s carbonation front and depth

0 day 3 days 7 days 14 days

50% Slag

70% Slag

Fig. 1 Cross-sectional images of the carbonation front for 50% and 70% slag with increasing
accelerated carbonation testing time, 3D voxel size 0.086 mm3 [67, 81]
302 B. Mehdizadeh et al.

0 day 3 days 7 days 14 days

50% Slag

70% Slag

Fig. 2 Views of the carbonation depth with 50% and 70% slag during 14 days of accelerated
carbonation testing, 3D voxel size 0.086 mm3 [67]

of penetration with 50% and 70% slag during 14 days of accelerated carbonation.
The carbonation front can be identified by micro-CT, which is seen to increase for
50% slag addition with the carbonation zone steadily expanding with increasing
curing time. However, for 70% slag addition, the specimens are totally carbonated in
only 7 days. These findings demonstrate that the ideal slag addition, which mitigates
carbonation, is < 50%. Figure 3 shows that the carbonation depth estimated from
micro-CT appears to be the same as with the phenolphthalein method during 14 days
of accelerated carbonation without the addition of slag. These results prove that
micro-CT is a reliable and appropriate technique for characterizing the carbonation
depth of the concrete [67]. Table 1 is a summary of micro-and nano-CT techniques
and the test conditions of different samples for identifying carbonation behavior.

Fig. 3 Comparable results

for micro-CT and
phenolphthalein method for
determining carbonation
depth, 3D voxel size 0.086
mm3 [67]
Advances in Characterization of Carbonation Behavior in Slag-Based … 303

Table 1 Use of micro- and nano-CT with accelerated carbonation conditions

Type of test Sample or Test condition Water/
resolution cement
ratio
Ji-Su 6 weeks of Resolution Temperature 20 ± 2 °C, relative 0.5
Kim accelerated 65 µm humidity 60 ± 5% and 5% CO2
[45] carbonation test Sample: 390 × concentration
650 × 520 µm
Dong 4 weeks of Resolution Temperature 20 ± 2° 0.5
Cui [70] accelerated test 60 µm C, relative humidity 70 ± 5% and
Sample: 9 × 40 3% CO2 concentration
× 40 mm
Han 14 weeks of 0.066 × 0.066 × Temperature 20 ± 2° 0.53
Jiande accelerated 0.066 mm C, relative humidity 70 ± 3% and
[67] carbonation tes 20% CO2 concentration
Han 12, 13, 14 and 0.066 × 0.066 × Temperature 20 ± 2 °C, relative 0.53
Jiande 16 weeks of 0.066 mm humidity 75 ± 3%, CO2
[81] accelerated concentration 20 ± 3%
carbonation test
Keshu 5 and 6 weeks of Pixel size of Temperature 20 ± 2 °C, relative 0.53
Wan the accelerated 60 µm humidity of 70% ± 5%, CO2
[85] carbonation test concentration 20%

3 Conclusions

Tomographic techniques for assessing carbonation behavior in concrete are still in

the early stage of development, but advances made in the previous decade have
been significant. As has been demonstrated, micro-and nano-CT can efficiently and
nondestructively investigate micro- and nanostructural behaviors. These techniques
are suitable for analyzing micro- and nanoscale topologies and morphologies, and
studying the porosity network. The advantage and benefit of using micro- and nano-
CT 3D imaging to assess the pore network of the concrete is the volumetric insight
into the interactions between different phases and pores. The techniques may be
useful for characterizing the carbonation behavior of slag-based concretes, with r
providing greater insight into accelerated carbonation testing and the impact on the
pore structure. However, the need to use small-sized samples necessitates meticulous
sample preparation.

Acknowledgements The authors gratefully acknowledge the University of Technology Sydney

and Transport for NSW for financial support for this study.
304 B. Mehdizadeh et al.

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Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Application of Surface-Modified
Nanosilica for Performance
Enhancement of Asphalt Pavement

D. He, H. H. Chan, Z. H. Xiao, T. Wu, L. M. Leung, M. Sham, B. Chen,

S. F. S. Lee, and C. K. K. Kwan

Abstract The fatigue performance of asphalt binder can be significantly increased

by incorporating nanosilica, but due to the poor dispersion of nanosilica in asphalt,
high mechanical energy is required to achieve good homogeneity. The viscosity of
asphalt is nearly double when 3% nanosilica is added. Consequently, the mixing
temperature of asphalt with aggregate is dozens of degrees higher to compensate for
its high viscosity. To solve this problem, modification of nanosilica was introduced
and the effects of its performance on asphalt are reported here. Several types of
nanosilica at various dosages were modified and characterized by Fourier-transform
infrared spectroscopy (FT-IR) and thermogravimetric analysis (TGA). The properties
of asphalt, such as viscosity, rutting, fatigue, softening point, penetration and ductility
were investigated. The viscosity of asphalt with 3% modified nanosilica was 80%
lower than that of the asphalt with unmodified nanosilica and a 20 °C reduction in
temperature for mixing the asphalt binder with aggregate was achieved. Moreover, the
rutting and fatigue performance of asphalt with modified nanosilica were significantly
enhanced. The asphalt with modified nanosilica not only had lower viscosity but also
has similar performance to asphalt with unmodified nanosilica. The asphalt with
modified nanosilica had excellent performance and thus good practical application
potential.

Keywords Asphalt · Fatigue · Surface-modified nanosilica · Viscosity

D. He (B) · H. H. Chan · Z. H. Xiao · T. Wu · L. M. Leung · M. Sham · B. Chen

Nano and Advanced Materials Institute Limited, Kowloon, Hong Kong
e-mail: danhe@nami.org.hk
S. F. S. Lee
Hong Kong Wah for Development Limited, Hong Kong, Hong Kong
C. K. K. Kwan
On Fat Lung Innovative Resources Ltd., Hong Kong, Hong Kong

© The Author(s) 2023 309

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_31
310 D. He et al.

Fig. 1 Mechanism of surface-modification of nanosilica by silane coupling agents

Fig. 2 Viscosity of different 1.2

silica–asphalt mixes at 135 Nano-silica, 135 oC
and 160 °C Nano-silica, 160 oC
1.0
Treated-nano-silica, 135 oC
Viscosity (Pa S)

Treated-nano-silica, 160 oC
0.8

0.6

0.4

0.2

0.0 0.5 1.0 1.5 2.0 2.5 3.0

Weight ratio (%)

Figure 1 shows the mechanism of surface-modification of nanosilica by silane

coupling agents [1, 2] and Fig. 2 shows that the viscosity of asphalt was higher when
either nanosilica or treated nanosilica is added [3]. The viscosity also increased with
the content of silica added, regardless of its type. However, the viscosity of asphalt
with modified nanosilica was comparable to that of virgin asphalt, which verified the
purpose of surface modification.
The increase in viscosity caused by the nanosilica was more significant than with
treated nanosilica [4, 5]. Table 1 lists the rheological properties of virgin asphalt
(60/70), asphalt with nanosilica and asphalt with treated nanosilica. According to
its performance grade (PG), virgin asphalt (60/70) can be classified as PG 64. The
rutting parameter (G*/sin δ, at original) increased by 12% with the addition of 1% of
Treated nanosilica and the PG of asphalt binder with 1% Treated nanosilica increased
by one degree (PG 64 to PG 70). The fatigue parameter (G* × sin δ) of asphalt binder
with 1% Treated nanosilica decreased by 7% and fatigue performance was enhanced.
Compared with virgin asphalt, the asphalt with modified nanosilica not only showed
a significant increase in rheological properties but also had a similar viscosity [6, 7].
Application of Surface-Modified Nanosilica for Performance … 311

Table 1 Influence of treated nanosilica on the rheological properties of asphalt binder

Viscositya Rutting parameter Fatigue Softening point
parameter
(135°C, Pa s) G*/sin δa PGa G*/sin δb G*·sin δc (°C)
(kPa) (°C) (kPa) (kPa)

Virgin asphalt 0.44 0.89 64 3.15 3997.1 47.6

0.5% 0.55 1.08 70 3.44 3854.3 48.3
nanosilica
1% nanosilica 0.59 1.1 70 3.79 2628.3 48.55
0.5% Treated 0.45 0.99 ≈70 3.24 3906.5 48.3
nanosilica
1% Treated 0.46 1 70 3.3 3705 48.55
nanosilica
a Original binder; PG, performance grading, G*/sin δ ≥ 1.0 kPa
b Rolling Thin-Film Oven
c Pressure aging vessel (PAV-25 °C) residue

References

1. Karnati SR, Oldham D, Fini EH, Zhang L (2020) Application of surface-modified silica nanopar-
ticles with dual silane coupling agents in bitumen for performance enhancement. Constr Build
Mater 244:118324
2. Li F, Li H, Lai X, Wu W, Zeng X (2014) Study on the graft modification of γ-
mercaptopropyltrimethoxysilane on the surface of nano-silica. Inorg Chem Ind 2014:1006–4990
3. Yao H, You Z, Li L, Lee CH, Wingard D, Yap YK, Shi X, Goh SW (2013) Rheological properties
and chemical bonding of asphalt modified with nanosilica. J Mater Civ Eng 25:1619–1630
4. Bhat FS, Mir MS (2021) Rheological investigation of asphalt binder modified with nanosilica.
Int J Pavement Res Technol 14:276–287
5. Abed AH, Oudah AM (2018) Rheological properties of modified asphalt binder with nanosilica
and SBS. IOP Conf Ser: Mater Sci Eng 433:012031
6. Xiantao Q, Siyue Z, Xiang H, Yi J (2018) High temperature properties of high viscosity asphalt
based on rheological methods. Constr Build Mater 2018(186):476–483
7. Zhang J, Yang F, Pei J, Xu S, An F (2015) Viscosity-temperature characteristics of warm mix
asphalt binder with Sasobit. Constr Build Mater 2015(78):34–39
312 D. He et al.

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Effect of Different Additives
on the Compressive Strength of Very
High-Volume Fly Ash Cement
Composites

R. Roychand, J. Li, M. Saberian, S. Kilmartin-Lynch, M. M. Ul Islam,

M. Maghfouri, and F. Chen

Abstract The cement industry is responsible for about 5–7% of global greenhouse
gas emissions and with the rapid rise in global warming, it is imperative to produce
an ecofriendly alternative to Portland cement. Fly ash (FA) is an abundantly avail-
able and least utilized industrial byproduct with good pozzolanic properties that can
help reduce the carbon footprint of cement composites. We investigated replacing
80% of the cement content with different blends of FA, nanosilica (NS) and silica
fume (SF). Hydrated lime and a set accelerator were used to increase the pozzolanic
reactivity of the blended cement composites. The portlandite released with 20%
cement content was insufficient for the pozzolanic reaction of the blended cement
composites containing FA and SF, requiring externally added hydrated lime. The
addition of a set accelerator significantly increased the pozzolanic reaction and the
resultant compressive strength, and these increased with the increasing content of
the set accelerator. The replacement of SF with NS led to a remarkable increase in
the pozzolanic reaction. The corresponding compressive strength of FA mixed with
cement composites increased with increasing percentage composition of NS.

Keywords Cement · Compressive strength · Fly ash · Nanosilica · Set accelerator

R. Roychand (B) · J. Li (B) · M. Saberian · S. Kilmartin-Lynch · M. M. U. Islam

School of Engineering, RMIT University, Melbourne, VIC, Australia
e-mail: rajeev.roychand@rmit.edu.au
J. Li
e-mail: jie.li@rmit.edu.au
M. Maghfouri
Department of Civil Engineering, University of Ottawa, Ottawa, ON, Canada
F. Chen
Arup, Melbourne, VIC, Australia

© The Author(s) 2023 313

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_32
314 R. Roychand et al.

1 Introduction

An increasingly important priority worldwide is to transform waste into a valuable

resource for other applications, supporting the circular economy. The cement and
concrete industry is one of the major sectors that is actively contributing toward
a closed loop circular economy by incorporating a range of industrial byproducts
and waste materials [1–3], such as, fly ash (FA) [4, 5], slag [6, 7], waste tire rubber
[8–12], recycled concrete aggregate [13, 14], organic waste [15] and medical waste
[16–19]. However, the production of cement accounts for around 5–7% of global
greenhouse gas emissions [20, 21]. Considering the estimated 4 billion tonnes of
annual cement production by 2030, a remarkable volume of CO2 will be generated
worldwide [22, 23]. In this regard, extensive research is currently being undertaken
to reduce CO2 emissions by using industrial byproducts such as FA [24–27], slag [28,
29], silica fume (SF) [30, 31], and metakaolin [32] as supplementary cementitious
materials. FA is a byproduct of coal-fired thermal power plants, and is considered
an environmental pollutant and requires considerable financial and environmental
input for its safe disposal. However, due to its pozzolanic properties, it can be used to
substitute cement in concrete production, which not only improves the mechanical
and durability properties of the concrete composite but also contributes to cutting
down its carbon footprint.
Most of the worldwide FA production is categorized as ASTM Class F low
calcium FA [33, 34], which has low reactivity [35]. Researchers have explored
various methods of improving the pozzolanic reactivity of the FA, such as reducing
the particle size of FA, adding hydrated lime (HL), sodium hydroxide, lime water,
sodium silicate [12], a set accelerator (SA) [13], nano-alumina, nanosilica (NS),
nanocalcium carbonate, nanoferrous oxide, nanozinc oxide and nanotitanium oxide.
Although an alkaline activator is always required for the pozzolanic reaction of
FA, out of all other additives, NS [36] could be the highest performing additive for
significant enhancement of the mechanical properties of high-volume FA (HVFA)
cement composites. The high reactivity of NS is attributed to its high surface area
[37] and highly amorphous structure, which reacts with available lime to produce
additional strength-forming calcium–silicate–hydrate gel counterbalancing the low
early reactivity of FA. However, the key challenge in using NS in HVFA cement
composites is its high cost, which hinders its application in large-scale commercial
projects. An alternative to NS is SF, though with relatively lower reactivity because
of its micro-sized particles compared with the nano-sized particles of NS, which
delivers a considerable difference in their respective reactive surface areas and their
corresponding reactivities. Therefore, we carried out a comparative analysis of the
use of different micro- and nano-sized additives in HVFA cement composites at 80%
replacement of the cement content.
Effect of Different Additives on the Compressive Strength of Very … 315

Table 1 Chemical composition of ordinary Portland cement, fly ash, silica fume, nanosilica, and
hydrated lime
Oxides OPC FA SF NS HL
(%) (%) (%) (%) (%)
SiO2 19.8 72.5 88.2 99.5 1.4
CaO 62.7 0.5 1.2 – 74.3
Al2 O3 4.8 22.6 1.3 – 0.7
Fe2 O3 3.3 1.2 2.1 – 0.3
SO3 2.8 0.3 1.7 – –
MgO 2.2 0.3 1.6 – 0.5
Na2 O 0.2 0.1 0.1 – 0.2
K2 O 0.4 0.2 0.2 – 0.1
TiO2 – 1.3 – – –

2 Methods

The materials adopted in this experimental program were ordinary Portland cement
(OPC), FA, SF, NS, HL, SA, superplasticizer (SP) and potable water. The sand-to-
cement ratio was kept at 2.4. The chemical compositions of the different materials
are shown in Table 1, and the mineralogical compositions are presented in Fig. 1.
The mix designs studied in this experimental program are provided in Table 2. Three
replicates of 50 mm cube mortar samples were prepared for each mix design for 7- and
28 day compressive strength tests. The samples were cured as per AS1012.8.1:2014
until the time of testing.

3 Results and Discussion

Figure 2 shows the compressive strength values of different mix designs at 7 and
28 days of curing. It can be seen that by replacing 80% OPC with 65% FA and
15% SF in mix M1, there was a considerable reduction in compressive strength.
FA has low reactivity that negatively influences strength development [38], thereby
significantly reducing the 7- and 28 day compressive strength results of mix M1.
Moreover, class F FA has negligible lime content, and the calcium hydroxide released
by the hydration reaction of 20% OPC content was insufficient for accelerating the
pozzolanic reaction. In mix M2, an additional 5% HL was added to the mix design
as a percentage of the total cementitious material (CM). The increase in the alkaline
activator accelerated the pozzolanic reaction, thereby bringing about 76.4% and
108.5% rise in the 7- and 28 day compressive strength results, respectively, compared
with M1.
316 R. Roychand et al.

Fig. 1 Mineralogical composition of ordinary Portland cement, fly ash, silica fume, nanosilica,
and hydrated lime

In mix M3, sodium thiocyanate and calcium nitrate-based SA was used at 12.5 mL/
kg of CM. The addition of SA to mix M3 brought about 68% and 10% improvement
in 7- and 28 day compressive strength, respectively, compared with M2. In mix M4,
the quantity of the SA was doubled to 25 mL/kg of CM, which led to a considerable
increase in the pozzolanic reaction, thereby increasing the 7- and 28 day compressive
strength results by 11.7% and 14.1%, respectively, compared with M3. In mix M5,
the water/cement ratio of the mix design was reduced from 0.3 to 0.25. Because the
water–cement ratio has an inverse relationship with compressive strength, the 7- and
28 day strengths of M5 were increased by 21.4% and 7.8%, respectively.
In mix M6, SF was replaced by NS, which has a considerably higher surface area
than SF and accelerates the pozzolanic reaction of NS-modified mix designs. The NS
content was 4%, and the FA content was increased to 71% to keep the total cement
replacement level at 80%, with the remaining 5% being the HL content. Because NS
has a high surface area, replacing SF with NS significantly increased the SP require-
ment. Mix M6 showed a significant increase (i.e., 54.8%) in its 7 day compressive
strength. At 28 days, though relatively small, it showed a 6% improvement in its
compressive strength compared with M5. On increasing the NS content to 5% in
Effect of Different Additives on the Compressive Strength of Very … 317

Table 2 Mix designs

Design mix Percentage S/CM W/CM SA SP
Cementitious material (CM) (By weight) (mL/kg of
CM)
OPC FA SF NS HL
C 100 – – – – 2.4 0.30 – 4
M1 20 65 15 – – 2.4 0.30 – 8
M2 20 60 15 – 5 2.4 0.30 – 8
M3 20 60 15 – 5 2.4 0.30 12.5 8
M4 20 60 15 – 5 2.4 0.30 25.0 8
M5 20 60 25 – 5 2.4 0.25 25.0 12
M6 20 71 – 4 5 2.4 0.25 25.0 20
M7 20 70 – 5 5 2.4 0.25 25.0 25
M8 20 69 – 6 5 2.4 0.25 25.0 30
M9 20 67 – 8 5 2.4 0.25 25.0 40
FA, fly ash; HL, hydrated lime; OPC; ordinary Portland cement; NS, nanosilica; S, sand; SF, silica
fume; SP, superplasticizer (mL/kg of CM); W, water

Fig. 2 Compressive strength results

mix M7, the 7- and 28 day compressive strength showed a further increase of 9.4%
and 5.3%, respectively, compared with M6. In mix M8, the NS content was increased
to 6%, which brought about 19.3% and 2.7% increase in the respective 7- and 28 day
compressive strength results, which can be attributed to the increase in pozzolanic
activity due to the addition of highly reactive NS. Mix M9 increased the NS content to
8% and reduced the FA content to 67%. The SP content was considerably increased to
keep similar workability. The 7- and 28 day compressive strength increased by 19%
318 R. Roychand et al.

and 9.2%, respectively, reflecting the positive effect of nano-sized highly reactive
amorphous silica on the pozzolanic activity of blended cement composites.

4 Conclusions and Recommendations for Future Work

(C1) The cementitious blend of HVFA and SF, replacing 80% of OPC content,
required the inclusion of HL to increase the pozzolanic reaction of both the
SF and FA.
(C2) The addition of a SA increases the pozzolanic reaction of FA and SF, which
increases with increasing SA content. This increase in the pozzolanic reaction
is also reflected in the increase in compressive strength.
(C3) The replacement of micro-sized SF with nano-sized highly amorphous NS
significantly increases the pozzolanic reaction of the mix, resulting in signif-
icant increases in the 7- and 28 day compressive strength results as the NS
content increased. However, the SP requirement increased significantly to
maintain workability.
(C4) NS-blended very HVFA cement composites replacing 80% of the cement
content that can give 28 day compressive strength results on par with those of
a control mix containing 100% cement content.
(R1) Mix 9 had the highest results for improving the 7- and 28 day compressive
strengths. It should be explored further to look for any potential of further
enhancement of its mechanical properties.
(R2) Further work needs to be carried out on workability (slump), setting time,
durability and techno-economic analysis to carry this work forward.

Acknowledgements The authors gratefully acknowledge the support of the RMIT X-Ray Facility
and Arup.

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rubber and polypropylene fiber. Constr Build Mater 225:983–996
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aggregate incorporating recycled tyre waste. Constr Build Mater 264:120266
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forms: a waste utilisation approach to support circular economy. J Build Eng 38:102199
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Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Spalling Resistance of Hybrid
Polyethylene and Steel Fiber-Reinforced
High-Strength Engineered Cementitious
Composite

S. Rawat, Y. X. Zhang, and C. K. Lee

Abstract We analyzed the effect of elevated temperatures on the integrity of high-

strength engineered cementitious composite (ECC) made with a hybrid combination
of polyethylene (PE) and steel fibers. The 50 mm cube specimens were subjected
to temperature ranging from 200 to 800 °C at three different heating rates: 1, 5,
and 10 °C/min. Five different types of mixes with varying content of supplementary
cementitious materials and fibers were evaluated. No spalling was observed at 1–
5 °C/min heating rate and <400 °C. However, at a heating rate of 10 °C/min for
temperature 600–800 °C, all ECC specimens with a PE fiber volume of 1.25 and 1%
steel fiber spalled explosively. Moreover, cementitious matrix with silica fume was
more prone to spalling at 800 °C and the use of slag or quaternary blend of slag and
dolomite at an optimum content was effective in maintaining the integrity of the ECC
specimens even at very high heating rates. Thus, the type of cementitious matrix is
equally important to consider, as well as fiber type and content, while analyzing the
spalling resistance of ECC.

Keywords Fibers · Elevated temperature · Heating rate · Spalling

1 Introduction

High-strength concrete may undergo severe spalling when exposed to elevated

temperatures. The use of polypropylene (PP) fibers in concrete is the most widely
recognized method of preventing explosive spalling from both the economic and tech-
nical perspective. Though existing literature validates the suitability of PP fibers in

S. Rawat (B) · Y. X. Zhang · C. K. Lee

School of Engineering and Information Technology, The University of New South Wales,
Canberra, ACT, Australia
e-mail: s.rawat@student.unsw.edu.au
Y. X. Zhang
School of Engineering, Design and Built Environment, Western Sydney University, Penrith,
NSW, Australia

© The Author(s) 2023 321

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_33
322 S. Rawat et al.

spalling prevention, their sole addition may not be very effective in post-fire strength
retention. In addition, the use of PP fibers in concrete may not lead to any signifi-
cant improvement in the tensile performance to prevent durability-related damage.
Engineered cementitious composite (ECC) is an alternate material type with superior
tensile performance at normal temperatures and has also been found promising in
elevated temperature scenarios. Existing studies of the fire performance of ECC have
considered individual or combined roles of steel, PP, and polyvinyl alcohol (PVA)
fibers [1]. The melting of PP or PVA fibers prevents the build-up of vapor pressure and,
simultaneously, steel fibers can provide the required resistance to prevent mechan-
ical decay in the higher temperature range. However, PVA fibers are not suitable
for the development of ultrahigh-performance engineered cementitious composites
due to their relatively low tensile strength and hydrophilic nature [2]. Another type
of low-melting fiber, polyethylene (PE) fibers, have higher strength and modulus of
elasticity than PP or PVA fibers and cementitious composites with PE fibers have
exhibited superior ductility, strength, and energy absorption capacity.
There has been limited work on the fire performance of PE fiber-reinforced ECC
and therefore, this study is a further step towards understanding the effect of elevated
temperatures on the integrity or spalling resistance of high-strength ECC made with
a hybrid combination of PE and steel fibers.

2 Methods

Five different types of mixes were considered to simultaneously study the effect of
fiber content and type of supplementary cementitious material (SCM) on spalling
resistance. The constituents of the main mix (Mix 1) used in the study were first
optimized using the Grey Taguchi method and the Taguchi method with Utility
Concept to obtain optimum compressive and tensile performance. More details about
the optimization methodology and other mix parameters can be found Rawat et al.
[3]. In addition, four other mixes were considered to study the effect of heating rate
on mixes with varying ratios of SCM (slag, dolomite, and silica fume) and steel
and PE fiber content as shown in Table 1. The specimens were exposed to different
temperature ranges and heating rates to analyze their effects on spalling resistance
and residual compressive strength.
Fiber content is expressed as volume fraction of the mix, whereas all other
constituents’ ratios are expressed as weight proportion of the cement content. FA,
fly ash; HRWR, high range water reducer.
Spalling Resistance of Hybrid Polyethylene and Steel Fiber-Reinforced … 323

Table 1 Mix proportions used in the present study

Mix ID Binder Sand Water HRWR PE (%) Steel (%)
Cement Slag FA Dolomite
1 1 0.94 0.19 0.38 0.91 0.50 0.05 1.50 0.75
2 1 0.83 0.17 0.22 0.81 0.44 0.04 1.50 0.75
3 1 0.94 0.19 0.38 0.91 0.50 0.05 1.25 1
4 1 1.00 0.00 0.00 0.73 0.40 0.04 1.50 0.75
Cement Silica FA Dolomite Sand Water HRWR PE (%) Steel (%)
fume
5 1 0.11 0.00 0.00 0.44 0.22 0.02 1.50 0.75

3 Brief Discussion

Figure 1 shows the normalized compressive strength of mixes 1 and 5 after exposure
to different temperatures and heating rates considered in the study. In general, the
residual compressive strength decreased with increase in heating rate for all types of
mixes. Nevertheless, the effect was not significant, and the difference diminished at
higher heating rates. Both Mix 1 and 5 had a different cementitious matrix but the
same fiber content. However, all specimens of Mix 5 spalled at 800 °C, indicating
that the role of thermally better performing SCMs as adopted in Mix 1 may help in
improving the spalling resistance if the fiber content is sufficient.
In addition, all the specimens of Mix 3 spalled at 600 and 800 °C at 10 °C/min.
Figure 2 shows the specimens of Mix 3 spalled inside the furnace, which further
confirmed the efficiency of PE fibers in mitigating spalling in ECC specimens. Mix
3 had a lower content of PE fiber, which may not be sufficient to mitigate the sudden
increase in vapor pressure due to the very high heating rate.

1
1.2
Normalised residual Compressive Strength

Normalised residual Compressive Strength

0.9 Mix 1 Mix 5

0.8 1

0.7
0.8
0.6
0.5 0.6
1°C/min 1°C/min
0.4
5°C/min 0.4 5°C/min
0.3
10°C/min 10°C/min
0.2
0.2
0.1
0 0
0 200 400 600 800 0 200 400 600 800
Temperature (°C) Temperature (°C)

Fig. 1 Normalized residual compressive strength of Mix 1 and Mix 5 at different temperatures and
heating rates
324 S. Rawat et al.

Fig. 2 Explosive spalling in Mix 3 specimens at 800 °C and 10 °C/min heating rate

4 Conclusions

We investigated the effect of heating rate on the spalling resistance of hybrid fiber-
reinforced ECC and found that a mix with 1.5% PE and 0.75% steel fibers may be
effective in preventing spalling even at very high heating rates. Moreover, this spalling
resistance was dependent on the type of SCM used. We also observed that the use of
thermally better performing SCM such as slag and dolomite can greatly improve both
spalling resistance and residual compressive strength. However, a more systematic
study including the role of specimen size is needed to obtain a deeper understanding
of the effect of PE fiber content and the mix proportions on the spalling resistance
of ECC.

References

1. Rawat S, Lee CK, Zhang YX (2021) Performance of fibre-reinforced cementitious composites

at elevated temperatures: a review. Constr Build Mater 292
2. Liu JC, Tan KH (2018) Fire resistance of ultra-high performance strain hardening cementitious
composite: residual mechanical properties and spalling resistance. Cem Concr Compos 89
3. Rawat S, Zhang YX, Lee CK (2022) Multi-response optimization of hybrid fibre engineered
cementitious composite using Grey–Taguchi method and utility concept. Constr Build Mater
319
Spalling Resistance of Hybrid Polyethylene and Steel Fiber-Reinforced … 325

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Roads Issues and the Social License
to Operate

W. Young and M. Shackleton

Abstract Because it is agreed that the transport system needs the support and
engagement of the public in its development and implementation, we explored the
concept of the “social license to operate the road system” (SLORS) to show how
it can assist in road policy and network implementation. We looked at 18 policy
issues and identified their placement in a SLORS framework. The policy issues were
categorized into 5 zones: User Advocacy Zone, Support Zone, Equilibrium Zone,
Tolerance Zone and Opposition Zone. The relevant zone provides information to
policy makers on the public’s view of the policy. For instance: in the Advocacy Zone
issues such as Driver behavior should improve, Roads must be safe for all users, the
Physical quality of the road and their surface should improve, and Road travel should
be more environmentally sustainable in the future are supported—and should be rela-
tively easy to implement. In the Opposition zone People paying a toll or road charge
for each trip, Private companies having a large role in planning and management of
roads, Increased congestion and Increased traffic on roads in the future receive less
support and thus will require considerable effort marketing their implementation.
SLORS may assist in pointing the policy maker in the best direction to get the policy
supported.

Keywords Policy · Roads · Social license · SLORS (Social License to Operate

the Road System)

W. Young (B)
Department of Civil Engineering, Monash University, Clayton, VIC, Australia
e-mail: bill.young@monash.edu
M. Shackleton
National Transport Research Organisation, Port Melbourne, VIC, Australia

© The Author(s) 2023 327

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_34
328 W. Young and M. Shackleton

1 Introduction

It is generally considered that the transport system needs the support and engage-
ment of the public in its development and implementation. A social model approach
to developing roads policy looks beyond the transport sector, beyond governments
and beyond the road community to build wider acceptance of transport solutions.
Here we describe at an approach for gaining insight into how the community views
policies related to the performance of the road transport system: the “social license to
operate the road system” (SLORS). In exploring the concept of the SLORS, we show
how it can assist in the implementation of road policy and network considerations.
We initially examined the (SLO) approach and how it adds to our understanding
of the introduction of new products. The policy issues were then categorized into
5 zones: User Advocacy Zone, Support Zone, Equilibrium Zone, Tolerance Zone
and Opposition Zone. Interpretation of the SLORS in a policy sense is outlined
graphically.

2 Research into the Public’s Acceptance of Road Transport

Systems

At its base, road infrastructure operations require a coordinated, efficient and well-
informed planning process, triple bottom line assessment and strategic asset manage-
ment system. Because it is generally thought that an acceptable road system must
meet the needs of the community, their view of the transport policy is an impor-
tant input into these processes and, in particular, their implementation. Community
consultation and people’s behaviour are the main methods of collecting this informa-
tion. As transport systems become more complex and invasive the general community
expresses more concern about its impacts, and many transport projects and govern-
ment decisions are questioned. In some cases transport projects have been stopped,
delayed or not started. Public acceptance has been suggested as an important factor
for the successful realization of transport plans, projects and policies, and a number
of approaches have been used to quantify the performance and customer satisfaction
with transport infrastructure [1]. In particular, BITRE looked at “…how customer
preference might be better incorporated to improve the long term efficiency and oper-
ation of Australia’s infrastructure asset” [1]. A key component of communication
with the community in the roads area is transparency and a need to quantify their
views. We add to the above approaches by exploring the quantification of SLORS.
In particular, how can SLORS be used to assist in the implementation and operation
of particular policies and the introduction of new products.
Roads Issues and the Social License to Operate 329

3 Consideration of SLORS

We examined how to include the public in the development of road policy decisions
as part of the planning and policy development processes. The views of the commu-
nity about the future of roads comprise an essential input into each of the planning
processes because they are the system’s end-users. This process can be assisted by the
quantification of a SLORS. At the policy development level there could be varying
levels of acceptance of particular issues by the community.
The survey methodology and data used in this study were collected from a series of
cross-sectional questionnaires and focused group open-ended surveys over a period
of 3 years. Industry and respondents were asked about the major issues and these
were developed into a series of formal questionnaires. The questionnaire sample
was collected using social media and a panel. The policy issues were developed
over a 3-year period and changed as new issues were raised and old issues refined.
Respondents were asked what “should” take place and what they think “will” take
place. The “will” and “should” questions formed the basis of the quantified views
using a 5-point semantic scales. Here, we only look at the stage 3 data.
There were 18 policy issues considered in stage 3 (Fig. 1) and these formed the
basis for the exploration of the SLORS. Figure 2 presents the SLORS framework for
consideration of these issues in a policy sense. The SLORS can be measured in terms
of how the public perceives the acceptability or not of particular transport policies.
That is, whether a policy issue will and should take place in road operation. The “will”
provide an indication of what the respondent thinks the particular issue will be like in
30 years; “should” indicates what they think should occur. These 2 perceptions form
a grid showing the implications of the SLORS: what should happen is the vertical axis
and what will happen is the horizontal axis. The SLORS measure is the difference
between the “should” and “will” ratings. Other relationships between the “should”
and “will” ratings, such as ratios, logarithms etc. are shown by the graph, and will be
explored further in the future. The vertical distance between the “should” rating and
the line of equality (≈45° line) between the “will” and “should” ratings pictorially
represents the SLORS, which is the discrepancy between what the respondent thinks
should take place and what will take place. This discrepancy, depending on its positive
and negative value, will indicate the level of support, tolerance and opposition to the
particular issue. More specifically, the line of equality (Should–Will) is the level of
acceptance the community thinks these issues will take place and that they should
take place at the same level at a particular point in time. Above the line of equality is
where the community thinks that these policy issues should take place and that they
will take place at a lower level, which is a level of advocacy and support for these
issues. Below the line of equality is opposition to particular issues. These measures
of the issues can be subdivided into 5 zones (Fig. 2): User Advocacy Zone, Support
Zone, Equilibrium Zone, Tolerance Zone and Opposition Zone.
330 W. Young and M. Shackleton

Key to data point numbers (summarized descriptions)

Advocacy Zone
1 Driver behavior 2 Roads much safer for all users
3 Physical quality of roads improves 5 Road travel more environmentally
sustainable
Support Zone
7 Public transport more common 9 Technology improves the level of
mode service of roads
11 Cars/trucks separated from cyclists
and pedestrians
Equilibrium Zone
8 Local roads priority to 13 Increase in number & capacity of
pedestrians/cyclists roads
14 Road and rail tunnels more 16 Smart road infrastructure
common in the future becomes necessary
Tolerance Zone
4 Roads and their use remain the 6 Car, truck and bus travel all
same driverless
15 Parking on major roads not be
permitted
Opposition Zone
10 People pay a toll or road charge for 12 Private companies have a larger
each trip role
17 Increasing congestion on our roads 18 Increasing traffic on our roads in
in the future the future

Fig. 1 Issues presented in a SLORS framework

Roads Issues and the Social License to Operate 331

Fig. 2 Some implications of the relativity of average “should” and “will” ratings in a SLORS
framework

4 The Data

The data we used was a subset of data collected for a broad study of the future of
roads. Stage 3 data includes only data where all “should” and “will” ratings were
given, and were collected between 24/2/20 and 24/4/20. The 18 policy issues (Fig. 1)
were included in the questionnaire. The respondents were asked to answer how likely
they thought that each statement described WILL occur (Fig. 1) and to what extent
they agreed that what the statement described SHOULD occur (Fig. 1). These ratings
formed the base for the SLORS (Fig. 1) and are discussed below.

5 Quantification of the SLORS

5.1 Respondents’ View of What SHOULD Happen to Roads

The measure of the community’s support for particular policies comes in many parts.
One is what they think should happen and was measured using a 5-point Likert scale:
Strongly Agree (5), Agree (4), Neutral (3), Disagree (2) and Strongly Disagree (1).
Figure 1 presents the mean for “should” rating for each policy issue. The quantitative
results of the data showed that most things were changes that people thought “should”
happen, as measured by a mean equal to and above an average score of 3.00. The
average rating ranged from a high of 4.22 (separation of bicyclists and pedestrians
332 W. Young and M. Shackleton

from cars and trucks) to a low of 2.34 (paying tolls and road charge)s. Overall, the
“should” ratings and the percentage of people disagreeing with the issue provided a
good indication of the community support for particular policy issues. This is one
measure of how the community views the road system and should be taken into
account when considering particular policies.

5.2 Respondents’ View of What WILL Happen to Roads

A complementary measure to what “should” happen is the community’s view of

what “will” happen and was also measured on a 5-point Likert scale: Very Likely
(5), Likely (4), Neutral (3), Unlikely (2) and Very Unlikely (1). The ranking of average
ratings for the “will” scores indicated the preferred products. The quantitative results
of the data showed that most things are changes that people think will happen. The
average “will” ratings ranged from a high of 4.26 (increasing traffic) to a low of 2.37
(improving driver behaviour). The respondents’ indication of what will happen is not
an indication of the SLORS but taken together with the support levels for the policy
issues provides a strong indication of the difference between what people want to
happen and what they think will happen. This will be considered in the next section.

5.3 The SLORS

The “should” ratings provide one measure of the community’s view of road trans-
port policy, but do not, however, provide an indication of the community’s dissat-
isfaction with the policies, because they understand some things will happen. The
difference between the “should” and “will” ratings provides a measure of the level
of discrepancy or dissatisfaction that people have with the transport policy. More
specifically, the “will” ratings (Fig. 1) indicate what the respondents think the road
system will be like in the future. The combination of this measure with the “should”
rating (Fig. 1) gives the magnitude of dissatisfaction with what should and will take
place in 30 years; that is, the SLORS. As shown in Fig. 1, the “should” minus the
“will” difference for each respondent is averaged over the entire population. A posi-
tive ranking indicates that the desirability of a measure exceeds the likelihood of it
happening, which is the support region in the SLORS diagram (Fig. 1). A negative
indicates things that will happen but should not: The opposed region.
The SLORS diagram (Fig. 1) shows the mean “should” and “will” ratings for the
data set plotted against one another, with the 45° line of equality between “should”
and “will”. Because the SLORS is estimated by subtracting “will” from “should”,
attributes with a positive rating are above the line and show where perceived desir-
ability exceeded perceived likelihood of eventuating. For instance, it shows that
although the respondents think that driver behaviour should improve (1.44), roads
should be safer (1.15), and the physical quality of the roads should improve (0.99).
Roads Issues and the Social License to Operate 333

These are unlikely to happen despite net support. Those attributes with a negative
rating are below the line of equality and show where the perceived likelihood of
happening exceeds perceived desirability. Increased congestion (−1.50), increased
traffic (−1.34), road charges (−1.18) and the role of private companies in planning
(−0.86) fall into this category and have less support. They are likely to happen but
people think they should not happen. Increased action would need to be put in place
to achieve these goals.

5.4 Interpreting the SLORS

There are two major differences in this application of SLO from previous social
license to operate applications in roads.
1. Road users have views on a wide range of policy issues from planning through
to constructed infrastructure and even the behavior of users. It does not look only
at 1 project as do other applications of SLO.
2. Other than in project-specific studies, users are seldom asked for their views on
what they think should happen and at the same time what they think will really
happen on roads. Generally, studies look at only what should happen, which is
only half of the picture.
These contributions by users are potential keys to solving a number of imple-
mentation agency issues before they occur. Strategic initiatives and policy changes
may meet significant stakeholder opposition, effectively preventing implementation
of something that makes engineering and/or economic sense. An example is road-
pricing and the concept that users should pay for the network capacity they use, and,
critically, only the network capacity that they use. The net result is that custodians of
the road network often face a choice—do nothing or risk a public backlash. The net
result is long lead-in times for projects and changes, giving rise to perceptions that
change is slow. Thus, an important part of change is the marketing of the changes to
stakeholders, including road users.
The methodology described here may allow an agency that is considering a basket
of policy options or initiatives to gain some insight as to the phasing strategy. They
can begin with options with a high SLORS to start getting some benefits of change,
while change requiring a significant shift in attitudes can be shepherded over longer
periods to win road users over, or gain their license to bring about the envisaged
change.
Thus, points below the equality line in Fig. 2 represent issues where “will”
is greater than “should”; that is, factors where support is less than the perceived
inevitability or where social license is deficient. Points above the equality line repre-
sent issues where the social license is positive; road users want the change more than
they perceive it to be likely to eventuate. Points on the line represent issues where
support and perceived eventuation are in balance.
334 W. Young and M. Shackleton

This gives some idea of what the ‘quick wins’ might be and where significant
effort may need to be put into winning the support of road users. This approach can
be incorporated into the SLORS framework (Fig. 3). The idea of road users taking
on an advocacy role (User Advocacy Zone) against hold outs for changes the agency
wants to make would have strong appeal to the agency. It avoids the agency being
accused of forcing their change through and reduces the expenditure of resources
that the agency needs to effect the change. To take advantage of this, it is necessary
to know at what point support tips into advocacy on the road network.
Similarly, it would be helpful to know—for those factors in the social license
deficiency zone (Opposition Zone)—which are the issues where it is possible to
change road user perceptions and gain social license in a reasonable amount of time,
and which are candidates for really long-term efforts, or for which a rethink may be
needed.
To illustrate how a finer gradation of support/opposition is introduced, 2 lines have
been added to the SLORS framework in Fig. 3 to represent these tipping points: 120%
(Support Zone) and 80% Opposition Zone) of the line of equality value (Equilibrium
Zone = mix of tolerance and support).
By adopting such an approach to assessing social license, an agency would then
have the following strategies available:
1. Supply users with materials and publicity for items in the Advocacy Zone
2. Embark on minor ‘marketing’ of items in the Support Zone to reduce resistance
3. Focus efforts on items in the Tolerance and Equilibrium zones in order to gain
some degree of social license for them
4. Rethink the desirability of items in the Opposition Zone, form coalitions with
others trying to achieve the same measures or make plans for a long process of
persuasion.
Figure 3 shows the 18 policy issues identified in their SLORS zones, for the tipping
points described above. Two of the Opposition Zone factors are in fact outcomes—
increased traffic and congestion. Therefore, if an agency hoped to do nothing to
reduce traffic or congestion that “do nothing” approach would be resisted. Conversely
though, any actions taken to ameliorate outcomes in the Opposition Zone can be
assumed to have the “Support”, or enjoy the “Advocacy”, of road users.
Put another way, the “should”, “will” and SLORS ratings provide guidance on
the acceptance or not of particular policy issues by the community. These need to
be put into an overall policy context. Figure 4 shows the full SLORS framework
and the ratings for the 18 policy issues, illustrating the policy tipping points. Issues
can be divided into zones: Advocacy support (Alignment between desirability and
likelihood); Issues where people show some Tolerance; and Issues where there is
Opposition.
Advocacy support can be found in the areas of improvement: driver behaviour
(1.44), road safety improvements for all road users (1.15), improvements in the
quality of the road surface (0.99) and road transport being environmentally sustain-
able (0.83). Support is likely for policies related to: physical separation of active
transport from cars and trucks (0.61), use of technology to improve level of service
Roads Issues and the Social License to Operate 335

Key to data point numbers (summarized descriptions)

Highly desirable, Highly likely
9 Technology improves the level of 1 Cars/trucks separated from cyclists
service of roads 1 & pedestrians
13 Increase in number & capacity of 1 Road and rail tunnels more
roads 4 common in the future
15 Parking on major roads not 1 Smart road infrastructure becomes
permitted 6 necessary
Highly desirable, less likely
1 Driver behavior improves 2 Roads much safer for all users
3 Physical quality of roads 5 Road travel more environmentally
improves sustainable
7 Public transport more common 8 Local roads priority to
mode pedestrians/cyclists
Less desirable, less likely
4 Roads and their use remain the 6 Car, truck and bus travel all
same driverless
Less desirable, highly likely
10 People pay a toll or road charge 12 Private companies have a larger
for each trip role
17 Increasing congestion on our 18 Increasing traffic on our roads in
roads in the future the future

Fig. 3 Alternate zone definitions for the SLORS framework

336 W. Young and M. Shackleton

Fig. 4 A comprehensive SLORS framework

(0.49) and public transport being a more common mode choice (0.47). A mix of
tolerance and support may be obtained for: priority given to: more tunnels for road
and rail (0.12), active travel in local and shopping roads (0.05), more smart infras-
tructure (0.05) and increased number and capacity of roads (0.03). For these factors,
the views of desirability and likelihood are similar in magnitude. The factors where
the respondents may show some tolerance are: banning parking on major roads
(−0.14), roads and their use will stay the same (−0.33), and automation of the road
transport network (−0.47). Those factors where there is likely to be opposition are
private sector being involved in planning (−0.86), paying for the use of the network
(−1.18), increased traffic in the future (−1.34), and increased congestion (−1.50).
Under the SLOR zones posited, Australian road agencies can rely on user
advocacy for actions and strategies to:
1. Improve driver behavior, make roads safer, better physical quality and make use
of roads more environmentally sustainable.
2. Reduce traffic and congestion, by virtue of “increased traffic” and “increased
congestion” being in the “opposition” zone as posited.
3. Support strategies to improve the attractiveness of public transport as a choice,
increase deployment of technology (ITS/VMS) to improve levels of service on
roads and to physically separate motorized traffic from cyclists and pedestrians.
4. Work on public acceptance of an all-autonomous motorized vehicle fleets, a ban
on parking along major roads and any plans to allow road usage to stay the same.
5. Reconsider pricing as a strategy, or significantly change how it may be applied,
or simply educate the road users on an intelligent reframing of what “pay per
use” means.
Roads Issues and the Social License to Operate 337

6 Conclusions

We explored an approach to gaining insight into how the community views the
performance of the road system: the concept of a SLORS and shows how it can assist
in the implementation of policy and network considerations. The measurement of
the SLORS was described and 18 policy issues were rated for whether they “should”
and “will” be implemented. The SLORS is the difference between the “should” and
“will” ratings. The policy issues were categorized into 5 zones: User Advocacy Zone,
Support Zone, Equilibrium Zone, Tolerance Zone and Opposition Zone. The Zone
into which an issue falls provides information to policy makers on the public’s view
of the policy. For instance: in the Advocacy Zone issues such as driver behaviour
should improve, Roads must be safe for all users, the Physical quality of the road
and their surface should improve, and Road travel should be more environmentally
sustainable in the future are in the user Advocacy Zone. In the Opposition Zone
People paying a toll or road charge, Private companies having a large role in planning
and management, Increased congestion and Increased traffic on roads in the future
receive less support. This information may assist in pointing policy makers in the
best direction to gain community support.

Reference

1. Australian Government. Bureau of Infrastructure, Transport and Regional Economics (2017)

Measuring infrastructure asset performance and customer satisfaction: a review of existing
frameworks. Report 147. Department of Infrastructure and Regional Development, Canberra

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
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An Intelligent Multi-objective Design
Optimization Method
for Nanographite-Based Electrically
Conductive Cementitious Composites

W. Dong, Y. Huang, B. Lehane, and G. Ma

Abstract Nanographite (NG) is a promising conductive filler for producing effec-

tive electrically conductive cementitious composites for use in structural health moni-
toring methods. Since the acceptable mechanical strength and electrical resistivity
are both required, the design of NG-based cementitious composite (NGCC) is a
complicated multi-objective optimization problem. This study proposes a data-driven
method to address this multi-objective design optimization (MODO) issue for NGCC
using machine learning (ML) techniques and non-dominated sorting genetic algo-
rithm (NSGA-II). Prediction models of the uniaxial compressive strength (UCS) and
electrical resistivity (ER) of NGCC are established by Bayesian-tuned XGBoost with
prepared datasets. Results show that they have excellent performance in predicting
both properties with high R2 (0.95 and 0.92, 0.99 and 0.98) and low mean absolute
error (1.24 and 3.44, 0.15 and 0.22). The influence of critical features on NGCC’s
properties are quantified by ML theories, which help determine the variables to
be optimized and define their constraints for the MODO. The MODO program is
developed on the basis of NSGA-II. It optimizes NGCC’s properties of UCS and
ER simultaneously, and successfully achieves a set of Pareto solutions, which can
facilitate appropriate parameters selections for the NGCC design.

Keywords Machine learning · Multi-objective design · NSGA-II · NGCC ·

Optimization

W. Dong · B. Lehane
Department of Civil, Environmental & Mining Engineering, University of Western Australia,
Perth, WA, Australia
Y. Huang (B) · G. Ma
Department of Civil & Transportation Engineering, Hebei University of Technology, Tianjin,
China
e-mail: yimiao.huang@hebut.edu.cn

© The Author(s) 2023 339

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_35
340 W. Dong et al.

1 Introduction

Nanographite (NG) is a promising conductive filler for producing highly effec-

tive electrically conductive cementitious composites. As reported by [1], the NG-
based cementitious composite (NGCC) has efficient strength and conductivity as a
self-sensing construction material for use in non-destructive structural health moni-
toring (NDSHM). There are generally three forms of NG studied in NGCC, namely
graphene nanoplatelet (GNP), graphene oxide (GO), and reduced GO (rGO). Apart
from the form of NG used, as well as its physical properties, it has been confirmed
that other experimental factors, such as the dispersion of NG and curing condition,
are linked to the strength and conductivity variations of NGCC [2, 3]. Significant
experimental research has been conducted to determine the optimal values for those
factors during the production of NGCC [4, 5].
The purpose of design optimization of NGCC is to approach its high strength and
low electrical resistivity (ER) simultaneously, which can seem difficult to realize
by experimental methods. Once, an experiment could only consider one factor with
limited design values for variable-controlling analysis, making the design experi-
ence lack universality. Former experimental studies often discussed mechanical and
electrical objectives separately. Nevertheless, they have provided numerous test data.
We present our study that aimed to address the multi-objective design optimization
(MODO) problem of NGCC using advanced data mining techniques.
The study proposed a comprehensive data-driven computing and analyzing
method, which integrates algorithms of machine learning (ML) and non-dominated
sorting genetic algorithm (NSGA-II). First, the uniaxial compressive strength (UCS)
and ER of NGCC were modelled by Bayesian-optimized XGBoost with compiled
experimental datasets. During the modelling process, the Weight and SHAP (SHapley
Additive exPlanations) theories were used to interpret the ML models and determine
variables that were critical for optimization. Furthermore, the established models
worked as objective functions in NSGA-II to develop the MODO program. The
feasibility and accuracy of the proposed MODO method are discussed and proved
with a case study. Using this method, researchers can quickly find the optimal designs
of NGCC that satisfy their application demands.

2 Methods

2.1 Establishment of Calculation Models

We modelled the UCS and ER of NGCC through a complex ML modelling frame-

work, as shown in Fig. 1, to ensure the established models were accurate and reli-
able. The framework followed four steps. Original experimental datasets of UCS and
ER were constructed after an extensive literature review. They were then processed
with proper feature engineering strategies before estimator training. In Step 3 of
An Intelligent Multi-objective Design Optimization Method … 341

Fig. 1 Machine learning modeling framework

modelling, we applied four classic ML algorithms, namely support vector regressor

(SVR), random forest regressor (RFR), XGBoost, and back-propagation neural
network (BPNN), to the processed datasets. Three searching methods, comprising
grid searching (GS), particle swarm optimization (PSO), and Bayesian optimization
(BO), were used to tune the hyperparameters of the algorithms for achieving their
best performance. At this stage, estimators based on four algorithms were compared
with each other to determine the most suitable models for UCS and ER. Finally,
we used SHAP theory to interpret the established models to identify critical input
variables and quantify their influence.

2.2 Development of the MODO Program

Figure 2 shows the workflow of the developed MODO program. The structure of the
program was based on the mechanism of NSGA-II. Individuals in the generated group
were vectors consisting of variables to be optimized. Lower and upper boundaries
should be defined for each variable. Other variables that were also considered in both
datasets were fixed. Afterwards, the generated group concatenated fixed variables to
form the input datasets for UCS and ER calculations using the established models.
Pipelines were tools packing feature engineering methods and making the format of
input datasets consistent. Individuals were ranked based on the UCS and ER results
342 W. Dong et al.

Fig. 2 Flowchart of the multi-objective design optimization (MODO) program

according to the Pareto rule. Individuals in the high Pareto ranks and with larger
crowding distances were kept for creating the next generation group. At the end of
the program, the final group was the Pareto set formed with optimal design solutions.

3 Results and Discussion

Bayesian-optimized XGBoost models proved the most suitable for the UCS and ER of
NGCC according to the comparative results, and their hyperparameter combinations
are given in Table 1. Models had minimal gaps between training and testing subsets
at the end of the training, indicating no over- or under-fitting issues. Mean absolute
error (MAE) and determination coefficient (R2 ) were two indexes for assessing the
accuracy of the established models. Small MAE values (1.24 and 3.44, 0.15 and
0.22) and high R2 scores (0.95 and 0.92, 0.99 and 0.98) of the two models yielded
satisfactory and reliable prediction abilities for the UCS and ER of NGCC. Figure 3
shows the feature importance ranking based on the SHAP interpretation results of
the two models. It can be seen that the UCS and ER prediction models share almost
the same feature importance ranking, where the mixing amount of NG (GC) was a
An Intelligent Multi-objective Design Optimization Method … 343

Table 1 Hyperparameter
Hyperparameter UCS model ER model
tuning results
Estimator numbers 381 128
Learning rate 0.19 0.25
Gamma 0.31 0.74
Maximum tree depth 13 6
Subsample size 0.40 0.64

Fig. 3 Feature importance ranking: a uniaxial compressive strength; b electrical resistivity

dominant variable in both properties. The high influence could also be observed in
other features: NG’s physical properties of thickness and diameter (GT and GD), the
water dosage and curing age.
The developed MODO program proved feasible in a case study. In the case study,
GNP was selected and the curing age was set at 28 days while the other six variables
were optimized. The optimization process is recorded in Fig. 4 and the program
successfully converged to the final Pareto set through iteration. As listed in Table
2, optimization results indicated that all the given design solutions qualified for the
NDSHM application with acceptable strength and conductivity. Higher UCS led to
higher ER (lower conductivity). Additionally, optimal values could be found for some
variables. For example, the ideal thickness of GNP was ≈6.36 nm. The water/cement
ratio was 0.32 in most solutions. The ultrasonication process for GNP dispersion was
better if it lasted for 30 min. The differences in output results between solutions were
mostly due to the changes in the dosage of sand and GNP, as well as the GD. UCS
and ER of NGCC became smaller by adding more GNP.
344 W. Dong et al.

Fig. 4 Optimization process of the multi-objective design optimization (MODO) program of

nanographite-based cementitious composite. (ER, electrical resistivity; UCS, uniaxial compressive
strength)

Table 2 Multi-objective design optimization (MODO) design solutions for nanographite-based

cementitious composite
S/C GD GT GC W/C US UCS ER
0.05 21.44 5.99 0.05 0.31 0.57 92.16 8.97
0.48 21.81 6.64 0.45 0.3 0.51 90.58 4.66
0.08 22.11 6.61 0.55 0.32 0.51 90.17 4.38
1.49 21.53 6.51 0.45 0.31 0.52 89.16 3.30
1.49 24.44 6.51 0.51 0.32 0.52 88.66 3.10
2.95 24.77 6.54 0.57 0.34 0.56 87.28 2.92
2.21 23.63 6.67 0.63 0.33 0.56 86.49 2.68
0.58 21.35 6.62 2.15 0.31 0.51 86.22 0.27
0.09 21.99 5.93 4.62 0.31 0.52 85.65 0.23
0.57 21.71 5.99 4.81 0.32 0.5 84.76 0.22
0.58 21.44 5.99 7.27 0.32 0.56 82.42 0.20
(ER, electrical resistivity; GC, content of NG; GD, diameter of NG, GT, thickness of NG; NG,
nanographite; S/C, sand/cement ratio; UCS, uniaxial compressive strength; US, ultrasonication; W/
C, water/cement ratio)
An Intelligent Multi-objective Design Optimization Method … 345

4 Conclusions

Robust and reliable calculation models were established for the UCS and ER of
NGCC by BO-tuned XGBoost. The SHAP interpretation results of the established
models identified significant influential factors, including NG dosage and other vari-
ables. Moreover, we explained their quantitative influence on the properties of NGCC
by analyzing their SHAP value distributions. NSGA-II was combined with the estab-
lished models as objective functions to develop the MODO program of NGCC. The
program proved feasible through a case study in which it successfully obtained the
Pareto set of design solutions. The given solutions determined the optimal values for
some variables.

Acknowledgements We appreciate the support of The University of Western Australia through the
“Scholarship for International Research Fees and Ad Hoc Postgraduate Scholarship”. The research
work was financially supported by the ARC Research Hub for Nanoscience-Based Construction
Material Manufacturing (IH150100006). Experimental data used in machine learning modelling in
this study were retrieved from published literature and all data sources are acknowledged.

References

1. Qureshi TS, Panesar DK (2019) A comparison of graphene oxide, reduced graphene oxide and
pure graphene: early age properties of cement composites. In: Proceedings of the International
Conference on Sustainable Materials, Systems and Structures (SMSS2019) New Generation of
Construction Materials. vol 1
2. Wotring EE (2015) Dispersion of graphene nanoplatelets in water with surfactant and rein-
forcement of mortar with graphene nanoplatelets. University of Illinois at Urbana-Champaign,
Master’s Theses
3. Liu J, Fu J, Yang Y, Gu C (2019) Study on dispersion, mechanical and microstructure properties
of cement paste incorporating graphene sheets. Constr Build Mater 199
4. Sun S, Han B, Jiang S, Yu X, Wang Y, Li H, Ou J (2017) Nano graphite platelets-enabled
piezoresistive cementitious composites for structural health monitoring. Constr Build Mater
136
5. Li X, Liu YM, Li WG, Li CY, Sanjayan JG, Duan WH, Li Z (2017) Effects of graphene oxide
agglomerates on workability, hydration, microstructure and compressive strength of cement
paste. Constr Build Mater 145
346 W. Dong et al.

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Machine Learning-Aided Nonlinear
Dynamic Analysis of Engineering
Structures

Y. Feng, Q. Wang, D. Wu, and W. Gao

Abstract A machine learning (ML) technique was used to assist in the dynamic
analysis of mixed geometric and material nonlinearities of real-life engineering struc-
tures. Various types of inputs of system properties were considered in the 3D dynamic
geometric elastoplastic analysis, giving a series of realistic nonlinear descriptions of
complex, large deformation structural behaviors. To resolve the numerical challenges
of solving the mixed nonlinear problems, a newly established ML technique using
a new cluster-based extended support vector regression (X-SVR) algorithm was
applied. With this technique, a surrogate model can be built at each time step in the
Newmark time integration process, which can then be used to predict the deflection,
force and stress of the relevant structural performance at different loading time stages.
To demonstrate the accuracy and efficiency of the proposed framework, practical
engineering applications with linear and nonlinear properties are fully demonstrated,
and the nonlinear behavior of the structure under predicted working conditions in
the future was predicted and verified in numerical studies.

Keywords Engineering structures · Machine learning · Nonlinear dynamic

analysis

Y. Feng (B) · Q. Wang · W. Gao

Centre for Infrastructure Engineering and Safety, School of Civil and Environmental Engineering,
The University of New South Wales, Sydney, NSW, Australia
e-mail: yuan.feng1@unsw.edu.au
D. Wu
School of Civil and Environmental Engineering, University of Technology Sydney, Sydney,
NSW, Australia

© The Author(s) 2023 347

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_36
348 Y. Feng et al.

1 Introduction

The nonlinear response of a practical structure is affected by various factors such

as system properties, operational coefficients, loads and environment, as well as
uncertainties in collected information and estimation models. Liu et al. investi-
gated the random mean values of elastoplastic responses of structure using a proba-
bilistic partial differentiation approach [1]. Feng et al. introduced a stochastic elasto-
plastic analysis of two-dimensional engineering structures with the aid of sampling-
based machine learning algorithm [2]. In our study, the uncertain parameters were
studied simultaneously within the nonlinear dynamical framework with the help of
an advanced machine learning (ML) technique [3]. By using the ML algorithm, an
explicit regression function can be obtained to represent the relationship between
the uncertain inputs and the nonlinear responses. Subsequently, frequent, and fast
nonlinear prognosis can be conducted to assess the nonlinear behavior of engineering
structures during the dynamic loading process.
Here, a brief introduction to the proposed ML-aided framework for nonlinear
dynamics of engineering structure is given. The two main components of the
approach are briefly introduced. First, the deterministic solution to geometric–elasto-
plastic dynamics is presented. Then, the novel ML technique named the “extended
support regression” is introduced. To demonstrate the accuracy and applicability
of the proposed framework, an illustrative numerical case is incorporated to build
the proposed framework and demonstrate the nonlinear response for the concerned
structure.

2 Methods

2.1 Solution to Geometric–Elastoplastic Dynamics

For structural systems with both material and geometric nonlinearities, the plastic
strain and second-order Green–Lagrange terms must be considered in the incremental
strain–displacement relations as:

Δε = Δεe + Δε p + Δεg = (B + Bg )Δu (1)

where Δεe , Δε p and Δεg denote the elastic, plastic and high-order strain increments;
B and Bg denote the material and geometric nonlinear strain–displacement matrixes
of the deformation. In the dynamics framework without the effects of external load
and damping, the equilibrium function of a nonlinear structural domain by using the
principle of virtual work can be represented as:
{ { {
δεT σd V − δuT τd S + δuT ρ üd V = 0 (2)
V S V
Machine Learning-Aided Nonlinear Dynamic Analysis of Engineering … 349

where ü denotes the virtual acceleration vector; ρ denotes the material density and
τ denotes the surface traction along the boundary. By following the incremental
strategy, and substituting Eq. (1) into the above virtual field, it can be re-expressed
as:
({ { ) { {
BT Dep Bd V + BgT DBg d V Δu = ΦT τ d S − (B + Bg )T σ d V
V V
{
S V

− ΦT ρΦd V ü = 0 (3)
V

Consequently, the global governing equation of the dynamic geometric-material

hybrid nonlinearities problem can be represented as:

(Kep + Kg )ΔU = Fex − Rin − M Ü (4)

By using the Newmark time integration technique, the solution to the above
equation can be acquired through:
(
(a0 M + Kep
t
+ Kgt )ΔUt+1 = Fex
t+1
− Rin
t
+ (a0 Ut + a1 U̇t + a2 Üt )M − a0 MUt
/[ ] / /
a0 = 1 β(Δt)2 , a1 = 1 (βΔt), a2 = 1 (2β) − 1
(5)
/ /
where Δt denotes the incremental time step; β = 1 4 and κ = 1 2 are set for all
the calculations based on the trapezoidal rule [4].

2.2 Extended Support Vector Regression

After thousands of simulations of deterministic nonlinear dynamic analysis based

on the finite element model, the collected structural response was considered as the
training dataset for the ML technique to recognize the relation between outputs and
input parameters.
The adopted extended support regression was the ML technique proposed by Feng
et al. [5]. It is based on the theory of classical support vector machine (SVM) and the
doubly regularized SVM (DrSVM), and it offers notable practicability in handling
nonlinear prediction problems. For detailed definitions of each parameter in Eq. (6),
see Feng et al. [6].

fˆN (x) = (pk − qk )T k̂(x) − êkT Ĝk υ∗k (6)

where pk , qk denote two positive kernelized parameters; k̂(x) denotes the kernel
matrix; êk , Ĝk denote two matrix vectors; υ∗k denotes the obtained solution.
350 Y. Feng et al.

2.3 Modelling

2.3.1 Geometric–Elastoplastic Dynamics of Transmission Tower

As a numerical example, the nondeterministic nonlinear behaviors of a 3D transmis-

sion tower structure were investigated, and the geometry is presented in Fig. 1. In
this case, the four supports were fixed and the magnitude of a time-varying pressure
P(t) = 8e10 × (1 − cos(20t)) Pa, was applied at the tip of the tower. By establishing
the surrogate system, the nonlinear responses of the tower were predicted at various
time steps. The uncertainty information of the system is listed in Table 1.
By substituting the new input information into the X-SVR model, the predicted
contour plots of nonlinear deflections of the tower structure at three time steps are

Fig. 1 Transmission tower geometry and adopted mesh condition

Table 1 Variational input data of transmission tower

Variational input Distribution type Mean Standard deviation Interval
E (GPa) Lognormal 210 10.5 –
υ Uniform – – [0.28, 0.32]
ρ (kg/m3 ) Normal 7380 370
σY (GPa) Beta 0.45 0.0225 –
H Lognormal 15 0.75 –
Machine Learning-Aided Nonlinear Dynamic Analysis of Engineering … 351

(a) t = 0.25 s, X-SVR (b) t = 0.25 s, Num (c) t = 0.50 s, X-SVR (d) t = 0.50 s, Num

Fig. 2 X-SVR predicted and numerical simulated nonlinear deflection of a transmission tower at
different times

shown in Fig. 2, as well as the deterministic numerical simulation results as verifi-

cation. In Fig. 2, the predicted nonlinear displacements from X-SVR correlate well
with the numerical simulation results.

3 Conclusions

In the numerical modelling example, the nonlinear response of the transmission tower
predicted by the proposed method was compared with the numerical simulation result
and satisfactory alignment was identified from the comparison. Consequently, the
new nonlinear dynamic analysis framework aided by the ML technique for engi-
neering structures was verified as effective for practical engineering problems. We
believe that the proposed framework can improve the accuracy and efficiency of the
relevant structural nonlinear dynamic’s evaluation process.

Acknowledgements The numerical computations were undertaken with the assistance of resources
and services from the National Computational Infrastructure (NCI), which is supported by the
Australian Government.

References

1. Liu N, Tang WH, Zhou J (2002) Reliability of elasto-plastic structure using finite element
method. Acta Mech Sin 18:66–81
2. Feng Y, Gao W, Wu D, Tin-Loi F (2019) Machine learning aided stochastic elastoplastic analysis.
Comput Methods Appl Mech Eng 357:112576
352 Y. Feng et al.

3. Feng Y, Wang Q, Wu D, Gao W, Tin-Loi F (2020) Stochastic nonlocal damage analysis by a

machine learning approach. Comput Methods Appl Mech Eng 372:113371
4. Newmark NM (1959) A method of computation for structural dynamics. J Eng Mech 85:67–94
5. Feng Y, Wang Q, Wu D, Luo Z, Chen X, Zhang T (2021) Machine learning aided phase field
method for fracture mechanics. Int J Eng Sci 169:103587
6. Feng J, Liu L, Wu D, Li G, Beer M, Gao W (2019) Dynamic reliability analysis using the
extended support vector regression (X-SVR). Mech Syst Signal Process 126:368–391

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Insights into the Size Effect
of the Dynamic Characteristics
of the Perovskite Solar Cell

Q. Li, D. Wu, and W. Gao

Abstract Driven by government policy and incentives, solar power production has
soared in the past decade and become a mainstay during the worldwide clean-power
transition process. Among the various next-generation photovoltaic technologies,
perovskite solar cells (PSCs) are the most important emerging area of research due
to their outstanding power conversion efficiency and affordable scale-up operation.
We adopted the nonlocal strain gradient theory and the first-order shear deforma-
tion plate theory to investigate the size-dependent free vibration behavior of PSCs.
The size-dependency in the nanostructure of the PSCs was captured by coupling
the nonlocal and strain gradient parameters. In accordance with the Hamilton prin-
ciple, the governing equations set was derived. Subsequently, the Galerkin proce-
dure was applied to address the dynamic characteristics analysis of PSCs with
simply supported and clamped edges. Compared with the size-insensitive traditional
continuum plate model, the current multiscale framework revealed a size effect on the
free vibration of the PSC. Moreover, some parametric experiments were conducted
to explore the impacts of scale length parameter, nonlocal parameter, and boundary
conditions on the natural frequency of the PSC.

Keywords Free vibration · Nonlocal strain gradient theory · Perovskite solar

cell · Size effect

Q. Li · W. Gao (B)
Centre for Infrastructure Engineering and Safety (CIES), School of Civil and Environmental
Engineering, The University of New South Wales, Sydney, NSW, Australia
e-mail: w.gao@unsw.edu.au
D. Wu (B)
School of Civil and Environmental Engineering, University of Technology Sydney, Sydney,
NSW, Australia
e-mail: di.wu-1@uts.edu.au

© The Author(s) 2023 353

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_37
354 Q. Li et al.

1 Introduction

The perovskite solar cell (PSC) has garned tremendous attention from the scien-
tific community over the past few years due to outstanding optical and electronic
properties [1]. This new-generation photovoltaic device rose to prominence in 2012
with an energy conversion efficiency of 9.7% and then rapidly achieved a new certi-
fied record with 25.7% in 2022 [2]. Although PSCs have proven their competitive
power conversion efficiencies and the prospect of further improved performance,
their structural response during their operational lifetime still lacks investigation [3].
The study of the free vibration of PSC lays a solid foundation for optimizing the
structure, because it is a crucial part of analyzing the dynamic response to various
loading scenarios.
For mainstream solar power generation, technologies cannot be deployed in the
field without accurately estimating their structural response. Despite the significance
in real-life applications, there are few studies on the size-dependent dynamic char-
acteristics of PSCs among the broad spectrum of available scientific reports that
emphasize the interest and importance of this topic. Thus, we attempted to fill this
gap by using nonlocal strain gradient theory (NSGT) to reveal the size effect of the
free vibration behavior of the PSC with both simply supported and clamped boundary
conditions.

2 Methods

2.1 Formulations

Grounded on NSGT [4], the internal energy density potential incorporating nonlo-
cality and higher-order strain gradient tensor were formulated as:

  1    
 
U εi j , εi j , α0 , εi j,m , εi j,m , α1 = εi j Ci jkl α0 x − x  , e0 a εkl dV 
2
 V
(1)
1    
+ l 2 εi j,m Ci jkl α1 x − x  , e1 a εkl dV 
2 V

where εi j and εi j are the strain tensors at the arbitrary point x and its neighboring
point x  in V , respectively; Ci jkl denotes the elastic coefficients of the classical
elasticity; α0 and α1 indicate the nonlocal attenuation function and the additional
kernel function, correspondingly, which are related to the nonlocal parameters e0 a
and e1 a, and the distance between the considered points x and x  in V; e0 and e1 refer
to the material constants; a is the internal characteristic length, l is a material length
scale parameter which describes the higher-order strain gradient field.
The constitutive equations in NSGT herein can be described as:
Insights into the Size Effect of the Dynamic Characteristics … 355

 
σ = α0 x  , x, e0 a Ci jkl εkl,m

dV  (2)
V

 
σ (1) = l02 α1 x  , x, e1 a Ci jkl ∇εkl,m

dV  (3)
V

where ∇ represents the Laplacian operator. The total stress tensor predicted by the
NSGT can be then expressed as:

t = σ − ∇σ (1) (4)

Based on the assumption that the same nonlocal attenuation functions and
parameters are taken, namely, e1 a = e0 a, then a general constitutive relation yields:
   
1 − (e0 a)2 ∇ 2 ti j = 1 − l 2 ∇ 2 Ci jkl εkl (5)

Integrating by parts, then collecting the coefficients of δu 0 , δv0 , δw0 , δθx , and δθ y
to zero, the governing equations of the size-dependent PSC can be obtained. Then,
corresponding
 to the investigated
 boundary conditions, the generalized displace-
ments u 0 , v0 , w0 , θx , θ y can be expanded in double trigonometric series. With the
aid of the Galerkin method, the governing equations can be derived in the following
form:

K − ω2 M = 0 (6)

By solving the above eigenvalue problem, the natural frequency of the nanostruc-
tures ωnl can be resolved from the smallest eigenvalue.

3 Numerical Results

In NSGT the size-dependency of the nanostructure is characterized as nonlocal

parameters e0 a and material length scale parameters l0 . It should be noted that the
current NSGT can degenerate into other lower-order continuum models by adjusting
the related coefficients. Specifically, a pure NET model, l0 = 0 is taken to elimi-
nate the strain gradient term. In terms of the system not including stress nonlocality,
e0 a = 0. For the classical continuum model, herein the first-order shear deformation
plate theory (FSDT), without considering the nonlocality and higher-order strain
gradient, e0 a = l0 = 0.
The PSC is fabricated with a layer stack sequence of FTO/TiO2 /Perovskite/Spiro-
MeOTAD/Au. The total thickness, length, and width of a PSC are h = 1587.5nm
and L a = L b = 0.2mm, respectively. The related material properties can be found
in [5]. The dimensionless parameters are introduced as e0 a = eL0aa , l0 = Ll0a .
356 Q. Li et al.

a b

Fig. 1 Size effect on the natural frequency of a PSC with simply supported edges (SSSS); b PSC
with clamped restraint (CCCC). NSGT, nonlocal strain gradient theory; PSC, perovskite solar cell

The natural frequency of the simply supported and clamped PSCs with various
dimensionless nonlocal and material length parameters is shown in Fig. 1. It can be
seen that, for both boundary conditions, the increase of the two size effect param-
eters results in distinct variations in the natural frequency of the PSC. With the
attendance of the nonlocal parameter, the natural frequency of the PSC demonstrates
a declined trend. However, the material length scale parameter tends to enhance ωnl
in the considered range. One possible explanation is that the nonlocal parameter
addresses the effect of stiffness-softening, whereas the material length scale param-
eter addresses the effect of stiffness-hardening. In particular, the effects induced by
the size-dependent coefficients become more prominent when their values approach
the geometric size of the structure. Moreover, under two boundary restraints, as
predicted, the clamped PSC possessed greater natural frequency than the supported
plate due to the additional constraints.

4 Conclusions

In the present study, the size-dependency on the dynamic characteristics behavior

of the PSC wasinvestigated based on NSGT. By introducing the nonlocal parameter
and the material length scale parameter, the size-tendency of the nanostructure was
captured in a complicated mix of stiffness-softening and stiffness-hardening. It was
revealed that both scale parameters play crucial roles in the free vibration of the PSC.
Moreover, it should be noted that the size-dependent model showed its characteristics
as the scale parameters approached the dimensions of the structure.

Acknowledgements This research was supported by an Australian Government Research Training

Program Scholarship and Australian Research Council projects IH150100006, IH200100010, and
DP210101353.
Insights into the Size Effect of the Dynamic Characteristics … 357

References

1. Nayak PK, Mahesh S, Snaith HJ, Cahen D (2019) Photovoltaic solar cell technologies: analysing
the state of the art. Nat Rev Mater 4:269–285. https://doi.org/10.1038/s41578-019-0097-0
2. Perovskite Solar Cells|Department of Energy n.d. https://www.energy.gov/eere/solar/perovskite-
solar-cells. Accessed October 26, 2020
3. Nair S, Patel SB, Gohel JV (2020) Recent trends in efficiency-stability improvement in perovskite
solar cells. Mater Today Energy 17:100449. https://doi.org/10.1016/j.mtener.2020.100449
4. Lim CW, Zhang G, Reddy JN (2015) A higher-order nonlocal elasticity and strain gradient
theory and its applications in wave propagation. J Mech Phys Solids 78:298–313. https://doi.
org/10.1016/j.jmps.2015.02.001
5. Li Q, Tian Y, Wu D, Gao W, Yu Y, Chen X et al (2021) The nonlinear dynamic buckling behaviour
of imperfect solar cells subjected to impact load. Thin-Walled Struct 169:108317. https://doi.
org/10.1016/J.TWS.2021.108317

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Non-probabilistic Informed Structural
Health Assessment with Virtual
Modelling Technique

Q. Wang, Y. Feng, D. Wu, and W. Gao

Abstract In real-life engineering, non-probabilistic structural information is very

common in many and varied disciplines. This class of information is characterized by
incompleteness and imprecision, such as interval, fuzzy sets, etc. Non-probabilistic
structural information can be reflected in the structural performance and cause it
to fluctuate within a specific range, instead of being deterministic. Thus, without
appropriate consideration of non-probabilistic information, serious or even disastrous
accidents may occur. Therefore, fully estimating the structural health status using
non-probabilistic information, especially detecting the lower and upper bounds of the
concerned structural response, is extremely significant in uncertainty-sensitive fields.
To conquer this challenge, a virtual modeling technique underpinning a structural
health assessment framework is introduced. The twin extended support vector regres-
sion (T-X-SVR) approach is embedded for virtual model construction. Continuous,
differentiable expression of the established virtual model allows the optimal solutions
for each interval analysis to be easily achieved. Information update is another inherent
feature, which enables structural health assessment to be implemented with updated
conditions without rebuilding the virtual model. To demonstrate the applicability of
the proposed virtual modeling technique underpinned structural health assessment
framework, the non-probabilistic informed elastoplastic nonlocal damage analysis
was investigated for engineering structures.

Q. Wang (B) · Y. Feng · W. Gao

Centre for Infrastructure Engineering and Safety, School of Civil and Environmental Engineering,
The University of New South Wales, Sydney, NSW, Australia
e-mail: qihan.wang@unsw.edu.au
D. Wu
School of Civil and Environmental Engineering, University of Technology Sydney, Sydney,
NSW, Australia

© The Author(s) 2023 359

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_38
360 Q. Wang et al.

1 Introduction

In the practical engineering field, structural health assessments attract more attention
to meet the current higher requirement for smart engineering. Mass information is
obtained via various devices (e.g., actuators, monitors, sensors, etc.), but the problem
is transferring from acquiring the data to information extraction or mining of the data.
Engineering structures may contain, experience or confront multifarious informa-
tion from different sources. Non-probabilistic information is one method to represent
information with the features of incompleteness and imprecision, such as interval,
fuzzy sets, etc. [1, 2]. Moreover, from industry practice, accumulated evidence has
repeatedly revealed that stochastic- or probabilistic-based information quantification
can have the dilemma that the probability distribution characteristics are challenging
to be credibly determined, because of insufficient amount or poor quality of the
experimental data. Thus, non-probabilistic structural information has widespread
applicability in real-life industries [3].
By considering the non-probabilistic characteristics within the information of
system inputs, the structural response correspondingly presents non-probabilistic
features (e.g., interval, fuzzy, or imprecise). Without loss of generality, non-
probabilistic information is herein considered as a fuzzy parameter. A structural
health assessment involving fuzzy information was conducted by seeking the fuzzy-
valued bounds or membership functions of the structural response. Through a
level-cut strategy, the fuzzy problem was transformed into a series of optimization
algorithms on the interval realizations.
However, in practical engineering applications, the relationship between system
information and the quantity of interests is normally underpinned, sophisticated,
and implicit. Directly implementing the optimization algorithms on this constitutive
relationship is extremely challenging. As a single deterministic calculation could
already be very time consuming, non-probabilistic uncertainty quantification with
large simulations to search for the extremes would become computationally infea-
sible. Thus, an alternative strategy to tackle the non-probabilistic informed structural
health assessment is proposed based on a supervised machine learning technique,
namely twin extended support vector regression (T-X-SVR) [4]. Supreme math-
ematical features of T-X-SVR allow the feasibility of optimal solutions on given
intervals being effectively and efficiently obtained in the established virtual model.
Furthermore, the virtual model-aided health assessment has an inherent advantage
of information updates without the need to reconstruct the model.
Non-probabilistic Informed Structural Health Assessment with Virtual … 361

2 Methods

2.1 Non-probabilistic Information

The non-probabilistic information herein is considered as the fuzzy variable. A fuzzy

variable ξ F is a gradual weighting of a vector space ϒ with the membership functions
{μξ F (x) : ϒ → [0, 1], ∀x ∈ K}. This membership function can be written as a set
of ordered pairs,

{(x, μ(x))|x ∈ ϒ ∧ μ(x) ∈ [0, 1]} (1)

The set of all fuzzy sets ϒ is denoted by ⎡(ϒ). For numerical implementations,
it is necessary to introduce α− levels. For a fuzzy variable ξ F and α ∈ [0, 1], α−
level cut can be written as,

ξαF := {x ∈ ϒ : μξαF ≥ α} (2)

It is significant to note that for each given degree of truth (or membership level)
α, the problem converts to an interval form. Each interval problem is conducted by
seeking the lower bound (LB) and upper bound (UB) of the concerned structural
response, which can be formulated as follows,

optimi ze : X (ξαF )
ξαF
(3)
F
s.t., ξαF ∈ [ξ Fα , ξ α ]

2.2 Virtual Model Construction

To provide a more robust effective and efficient manner to tackle this engineering-
simulated problem, a supervised machine learning technique, namely T-X-SVR [4]
was used for the virtual model construction. The established virtual model alterna-
tively depicts the implicit constitutive relationship between the system inputs and
the concerned structural responses. T-X-SVR aims to minimize the gap between
two bounds of the virtual model and the datasets from two different directions. The
optimal solution for weights and bias can be effectively obtained by solving two
quadratic programming problems (QPPs).
362 Q. Wang et al.

2.3 Virtual Model Aided Non-probabilistic Uncertainty

Quantification

After the virtual model construction, a series of optimization programming prob-

lems for fuzzy informed engineering problems can be conducted on the virtual
model instead of the original constitutive relationship. There are two main bene-
fits in utilizing the virtual model as an alternative: (1) A mathematical function for
each model possesses cheap computational costs; (2) On any specific interval of
the inputs, the optimal solutions for the estimation of extremes can be effectively
obtained through the derivation methods.

2.4 Numerical Investigation

A jet engine was investigated by considering fuzzy information within the material
properties [5]. According to that convergence study, the virtual model was constructed
by learning from 160 training samples. 1e3 Monte Carlo simulation (MCS) results
were considered as the benchmark. The established virtual model had relatively
high accuracy, with R-square nearly 1, and root mean squared error (RMSE) about
8e − 4. The Poisson’s ratio νC and density ρC of the ceramic are considered as fuzzy
parameters. νC was assumed to follow the triangular membership function, with
the support of [0.33, 0.37], and top of 0.35; ρC followed the trapezoid membership
function, with the support of [3.168, 3.232] g/cm3 , and top of [3.198, 3.202] g/cm3 .
The corresponding fuzzy-valued LB and UB of the concerned structural response
(i.e., critical load P (k N ) of the damage analysis) was estimated through the proposed
strategy, as shown in Fig. 1c.

(a) (b) (c)

Fig. 1 a Numerical model, b adopted FEM mesh of notched blade, and c estimated fuzzy-valued
lower bound and upper bound of the concerned structural response
Non-probabilistic Informed Structural Health Assessment with Virtual … 363

3 Discussion

This extended abstract introduced a virtual model-aided non-probabilistic structural

health assessment framework for engineering structures. The non-probabilistic infor-
mation was considered as a fuzzy parameter. Through a level-cut strategy, the problem
was formulated as a series of non-linear programming problems in the virtual model.
For the virtual model construction, a supervised machine learning technique, namely
T-X-SVR, has been adopted. The implicit, computational expansive constitutive rela-
tionship can be alternatively depicted as a mathematical equation, with continuous
and differentiable features. Thus, the optimal solution for fuzzy-valued bounds can
be effectively obtained in the virtual model. Moreover, information update can be
fulfilled by importing updated information into the virtual model, instead of recon-
structing the virtual model. The introduced virtual model-aided non-probabilistic
informed structural health assessment has great potential applicability in real-life
engineering.

Acknowledgements The work presented in this paper was supported by the Australian Research
Council projects IH150100006 and IH200100010.

References

1. Wang Q, Wu D, Tin-Loin F, Gao W (2019) Machine learning aided stochastic structural free
vibration analysis for functionally graded bar-type structures. Thin-Walled Struct 144:106315.1–
106315.19
2. Feng Y, Wang Q, Wu D, Gao W, Tin-Loi F (2020) Stochastic nonlocal damage analysis by a
machine learning approach. Comput Methods Appl Mech Eng 372:113371
3. Beer M, Ferson S, Kreinovich V (2013) Imprecise probabilities in engineering analyses. Mech
Syst Signal Process 37(s 1–2):4–29
4. Wang Q, Wu D, Li G, Gao W (2021) A virtual model architecture for engineering structures
with twin extended support vector regression (T-X-SVR) method. Comput Methods Appl Mech
Eng 386(1):114121
5. Wang Q, Feng Y, Wu D, Yang C, Gao W (2022) Polyphase uncertainty analysis through virtual
modelling technique. Mech Syst Signal Process 162(1–2):108013
364 Q. Wang et al.

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Modeling the Alkali–Silica Reaction
and Its Impact on the Load-Carrying
Capacity of Reinforced Concrete Beams

T. N. Nguyen, J. Li, V. Sirivivatnanon, and L. Sanchez

Abstract The alkali–silica reaction (ASR) is one of the most harmful distress mech-
anisms affecting concrete infrastructure worldwide. The reaction leads to cracking,
loss of material integrity, and consequently compromises the serviceability and
capacity of the affected structures. In this study, a modeling approach was proposed
to simulate ASR-induced expansion considering three-dimensional stress/restraint
conditions, and its impact on the structural capacity of reinforced concrete members.
Both the losses in concrete mechanical properties and prestressing effects induced
by the expansion under restraints are taken into account in the model. Validation
of the developed model is conducted using reliable experimental datasets derived
from different laboratory testings and field exposed sites. With the capability of
modelling both ASR-induced expansion and its impact on structural capacity, the
model provides valuable results to specify effective repair and/or mitigation strategies
for concrete structures affected by ASR.

Keywords Alkali-silica reaction · Concrete deterioration · Expansion · Finite

element

1 Introduction

Many concrete bridges and dam structures in Australia have been reported to be
affected by various degrees of deleterious alkali–silica reaction (ASR) [1]. These
affected structures require comprehensive diagnosis and prognosis protocols for
assessing the current degree of damage, forecasting the potential of further dete-
rioration, and evaluating the impact of ASR on structural capacity. Such information

T. N. Nguyen (B) · J. Li · V. Sirivivatnanon

School of Civil and Environmental Engineering, University of Technology Sydney, Sydney,
Australia
e-mail: thuc.nguyen@uts.edu.au
L. Sanchez
Department of Civil Engineering, University of Ottawa, Ottawa, Canada

© The Author(s) 2023 365

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_39
366 T. N. Nguyen et al.

is essential to specify efficient method(s) for remedial/rehabilitation and management

procedures for the structures.
In terms of the structural implication of ASR, it is interestingly observed from
several experimental studies that the load-carrying capacity of ASR-affected struc-
tures is not compromised, especially for shear capacity. An exception is evident in
Swamy and AlL-Asali [2], where the affected reinforced concrete beams lost up
to nearly 25% of their flexural strength at an expansion level of 0.518% measured
on the beams. The studies did, however, agree that the impact on load-carrying
capacity would become significant if the specimens were subjected to long periods
of exposure and underwent high expansion levels. As such, besides the adverse
impacts on the material performance of concrete, such as cracking and degradation
of mechanical properties, there is a certain favorable effect of ASR on the capacity
of structures at low expansion levels, such as a prestressing effect of restrained ASR
expansion (e.g., from the reinforcement [3]). The two most important questions
herein are: (1) to what extent does the prestressing effect contribute to the capacity
of affected structures; and (2) is this favorable outcome maintained or decayed as the
expansion increases? Therefore, investigations of the prestressing effect in relation
to the expansion advancement and degradation in mechanical properties are deemed
necessary.
In this study, an engineering-based finite element (FE) approach was developed to
model the ASR in reinforced concrete structures, considering the impact of reinforce-
ment restraints on anisotropic ASR expansion and the loss of concrete mechanical
properties. A case study of modeling the expansion and load-carrying capacity of
ASR-affected reinforced concrete beams is provided.

2 Methods

2.1 Modelling Approach

The modeling approach to assessing the impact of ASR on structural capacity is

presented in Fig. 1. The approach has three main steps: (1) estimating ASR-induced
expansion under non-restrained/ non-confined conditions, (2) modeling the effect of
restraints/confinements in 3-dimensional expansion, and (3) assessing the impact of
ASR on the load-carrying capacity.
In the first step, free ASE-induced expansion was forecasted using the ther-
modynamically based semi-empirical model proposed by Larive [4], then further
developed by Nguyen [5], which is capable of considering the effect of aggregate
reactivity, time-dependent temperature and relative humidity, concrete alkali content
and alkali leaching for forecasting expansion of field concrete. More details on the
semi-empirical model can be found in Nguyen [5].
Modeling the Alkali–Silica Reaction and Its Impact … 367

Fig. 1 Modelling of the expansion and load-carrying capacity of ASR-affected concrete members

In the second step, ASR-induced expansion in concrete under restraints/

confinements, which is significantly different and far more complicated in compar-
ison with free expansion of unrestrained concrete, was simulated in a FE model,
which required a constitutive model to consider the restraints/confinements effect
based on the multiaxial stresses developed in concrete elements. A general form of
the incremental ASR strain tensor to be implemented in the FE analysis could be
expressed as:
AS R, f r ee
ε̇ AS R = EW E T f (σ )ε̇V (1)

AS R, f r ee
where ε̇V is the free volumetric expansion of concrete which was calculated
per the semi-empirical model presented previously, f (σ ) is expansion-stress depen-
dent function accounting for the impact of stress state on ASR expansion, E is the
eigenvectors derived from the stress tensor, and W is the weight tensor that distributes
the volumetric expansion to each of three principal directions, given by:
⎡ ⎤
W1 0 0
W = ⎣ 0 W2 0 ⎦ (2)
0 0 W3

Determining the weight tensor was based on the empirical model from Gautam
et al. [6], which was derived from multiaxial testing schemes of ASR-affected
concrete. The weights calculated above were equivalent to the weights in three prin-
cipal directions, as such the incremental ASR strain tensor is as same as in Eq. (1)
to capture the ASR anisotropic behavior.
In addition, ASR causes loss of mechanical properties over time (i.e., modulus of
elasticity, compressive strength and tensile strength) to various degrees. Expansion-
dependent mechanical properties were implemented in the model to consider the
impact of the material degradation. With all these considerations, the developed FE
model was capable of assessing the impact of ASR on the structural capacity in the
third step of the approach.
368 T. N. Nguyen et al.

3 Case Study: Results and Discussion

In this section, a case study is presented for modelling of ASR-induced expansion and
consequently the ASR impact on the load-carrying capacity of reinforced concrete
beams tested by Fan and Hanson [7]. An overview of the test is presented followed
by a modelling briefing and some selected outcomes of the model.

3.1 Test Overview

Fan and Hanson [7] conducted a series of tests on reinforced concrete beams (150 ×
250 × 1500 mm) for ASR expansion and capacity. Two reinforced concrete beams
were prepared, namely, 5R1 and 5N1 (or reactive beam and non-reactive beam,
respectively), which used concrete mixtures containing reactive and non-reactive
aggregates, respectively, with the same mixture proportions. They were immersed
in an alkali solution at 38 °C with periodic expansion measurements for 1 year. The
expansion was measured from Demec studs mounted in the beams’ surfaces using a
Demec dial gauge at different locations.
After 1-year immersion in an alkali solution, the beams were tested for their
load-carrying capacity as shown in Fig. 2a. The load–deflection behaviors of the two
beams were almost identical despite a certain reduction in mechanical properties
of the concrete of 5R1 due to ASR. The behavior of the non-reactive beam can be
referred to as the undamaged concrete beam in comparison with the damaged reactive
beam.

Fig. 2 Geometric and boundary conditions of the reinforced concrete tested by Fan and Hanson
[7]
Modeling the Alkali–Silica Reaction and Its Impact … 369

3.2 Modelling for ASR-Induced Expansion

and Load-Carrying Capacity

Due to the symmetry of prism geometry and boundary conditions, only one-quarter
of the beam was simulated utilizing symmetric boundary conditions as shown in
Fig. 2b. The stress–strain behavior of the concrete defined at every 0.025% expansion
level (i.e., 0%, 0.025%, 0.05%, 0.075%, 0.1%, etc.) to represent the change in the
concrete’s mechanical properties as expansion increased.
Distribution of average expansion (FV1) throughout the beam is shown in Fig. 3a.
It shows a lower expansion in the area with both transverse and main longitudinal
reinforcement at the bottom, and higher expansion on the top and at the beam-end
with less reinforcement. Expansion in different locations and directions is plotted
in Fig. 3b alongside the measurements. With a higher ratio of reinforcement in the
longitudinal direction at the bottom, the expansion obtained at the bar level was
significantly lower than at other locations. Similar to experimental observations,
at the bar level, the expansion leveled off after 240 days of immersion, but kept
increasing in the longitudinal direction on the top.
Load–deflection results for the 5N1 beam are shown in Fig. 4a, indicating a
good agreement between the numerical and experimental results of load–deflection
behavior. Figure 4b shows the predicted load–deflection curve of the reactive beam
using the mean values of residual mechanical properties. First, the numerical results
are comparable to the experimental in terms of capacity. The predicted ultimate
loading value of the beam is ≈175.0 kN, and the value from test results was ≈177.3
kN. Similar to the test data, the numerical results showed an insignificant reduction
in the capacity of the affected beam despite the reduction in mechanical properties
as presented above. Second, the bending stiffness of the beam was slightly higher
than the measured result despite the reduction in concrete stiffness. The observation
aligned with observations from ISE [3], in which a favorable prestressing effect of
restrained ASR expansion helped to increase the stiffness and capacity of several
affected structures at low expansion levels.

4 Conclusions

This paper presents a modeling approach for the ASR expansion and capacity of
reinforced concrete members. The approach consists of both the semi-empirical
model and numerical model (i.e., FEM). The FE model could transfer the expan-
sion modeling results such as strains, stress state, residual mechanical properties of
concrete to modeling for load-carrying capacity of the affected concrete structural
members. Outcomes from utilizing the proposed approach for simulation of rein-
forced concrete beams tested in Fan and Hanson [7] show good agreement between
modeling and testing results, which indicates the capability of the model for fore-
casting long-term ASR-induced expansion and its impacts on structural capacity
 370

Fig. 3 Numerical and experimental ASR expansion at different locations for the reactive beam
T. N. Nguyen et al.
Modeling the Alkali–Silica Reaction and Its Impact … 371

Fig. 4 Load–deflection behavior of a the non-reactive beam 5N1 and b reactive beam 5R1

of reinforced concrete structures in the field. In addition, the case study shows an
insignificant impact of ASR on the load-carrying capacity at the expansion level of
lower than 0.2%.

References

1. HB79 S (2015) Alkali aggregate reaction: guidelines on minimising the risk of damage to
concrete structures in Australia, Cement and Concrete Association of Australia, and Standards
Australia, North Sydney, NSW
2. Swamy RN, AlL-Asali M (1989) Effect of alkali-silica reaction on the structural behavior of
reinforced concrete beams. Struct J 86(4):451–459
3. ISE (1992) Structural effects of alkali-aggregate reaction: technical guidance on the appraisal
of existing structures. UK
4. Larive C (1997) Apports combinés de l’expérimentation et de la modélisation à la compréhension
de l’alcali-réaction et de ses effets mécaniques. École Nationale des Ponts et Chaussèes, Paris
5. Nguyen TN (2021) Modelling alkali-silica reaction effects for condition assessment and capacity
evaluation of reinforced concrete structures (Doctoral dissertation)
6. Gautam BP, Panesar DK, Sheikh SA, Vecchio FJ (2017) Multiaxial expansion-stress relationship
for alkali silica reaction-affected concrete. ACI Mater J 114(1)
7. Fan S, Hanson JM (1998) Effect of alkali silica reaction expansion and cracking on structural
behavior of reinforced concrete beams. ACI Struct J 95:498–505
372 T. N. Nguyen et al.

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
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indicate if changes were made.
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Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Guidelines for Enzymatic Soil
Stabilization

B. O’Donnell, A. Swarup, A. Sidiq, D. Robert, and S. Setunge

Abstract There are numerous manuals to guide practitioners in utilizing traditional

additives in the construction of road, rail and dam construction but they fall short of
specific guidance for non-standard additive-based ecofriendly and cost-effective soil
stabilization. Increased attention has recently been on the use of non-standard addi-
tives for stabilizing weak soils due to environmental and cost concerns associated with
traditional additives. We summarize the specific guidelines of using environmental-
friendly enzymes to treat weak soils. We elaborate on the requirements and specifi-
cations for the Eko-Soil multi-enzyme product that is manufactured from water and
proteins extracted from fermented exudes of plants. Specific tests (laboratory and
field) and conditions required for soil stabilization using Eko-Soil enzyme are elab-
orated using the experience of past construction projects. The guide also elaborates
enhancing the efficiency of enzymatic soil stabilization by correctly incorporating
the required mixing proportions and pre-requisite condition tests. Professionals and
practitioners will benefit from using novel eco-friendly sustainable stabilization tech-
niques in the treatment of weak soils covering many applications including roads,
foundations, water containment areas, landfills, working platforms and slope erosion
control.

Keywords Environmental-friendly solutions · Enzymes · Guidelines ·

Practitioners · Soil stabilization · Weak soils

B. O’Donnell (B) · A. Swarup

Centre for Pavement Excellence, Asia Pacific Ltd (CPEAP Ltd), Langwarrin, VIC, Australia
e-mail: baod1942@bigpond.com
B. O’Donnell · A. Sidiq · D. Robert · S. Setunge
School of Engineering, Royal Melbourne Institute of Technology (RMIT University), Melbourne,
VIC, Australia

© The Author(s) 2023 373

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_40
374 B. O’Donnell et al.

1 Introduction

Worldwide, population growth at an unprecedented rate is leading to increased

demand for new civil infrastructure. The American Society of Civil Engineering
has estimated that private–public partnerships on roads alone have now increased to
US$67 billion per year in overall investments [1]. There is also insufficient infrastruc-
ture in major countries such as India and China due to limits on expansion imposed
by geographic boundaries and soil conditions [2]. In Australia, civil infrastructure
is increasing due to fewer land limitations than in other major countries, but there
are deficiencies in the soil conditions, leading to cracking in ≈50,000 residential
dwellings per year [3]. Due to these serviceability concerns, there is a high demand
for improving soil conditions through different techniques such as traditional method
of mechanical energy (i.e., hammering, vibration or rolling) or by the application of
additive-based soil stabilization.
Traditionally, lime has been used in soil stabilization and typically includes
hydrated high-calcium lime, dolomite lime, monohydrated dolomite lime and calcite
quicklime [4–8]. The physical mechanism of soil stabilization with the addition of
lime-based products involves fluctuation of clayey particles from plate-like to rod-
like, leading to a greater interlocking between particles. Hence a modified soil matrix,
and effectively, as the soil turns drier, it becomes less able to absorb water [9, 10].
Also, the physiochemical progress of the additional lime content in the soil matrix
potentially involves the hydration process; thus, when quicklime is added to the soil, a
chemical reaction takes place between the quicklime and the available water content
in the soil matrix. Subsequently, the hydrated lime reacts with alumina and silica,
leading to the formation of C–S–H (calcium–silica–hydrate) and C–A–H (calcium–
aluminate–hydrate), which are like Portland cement products. Hence, lime-stabilized
soil results in relatively impermeable layers of soil and significant load bearing [11,
12]. Although the usage of lime-based products has positive efficacy in soil stabiliza-
tion, there can also be drawbacks due to the carbonation process and sulfate attacks,
which can have negative economic impacts. The addition of lime product to the soil
leads to a change in the chemistry of pore water, thus increasing the pH level and
affecting the charge of the clayey particles. Also, the rate of carbonation is potentially
higher for soil exposed to the atmospheric environment, which places limitations on
the chemical reaction and pozzolanic activities and therefore becomes less effective
for the purpose of stabilization [13–15]. The addition of lime-based products leads
the soil to become more susceptible to sulfate attack which also limits the pozzolanic
activities, leading to less effective soil stabilization [16, 17].
Cement products are also used in soil stabilization applications and this process is
well known as cement stabilization in the geotechnical engineering field. By mixing
cement with the soil material, the cement particles chemically react with the available
water content in the soil [18], leading to the formation of crystal particles that bond
to each other and with the soil particles, achieving higher compressive strength
and stability [19, 20]. Thus, using cement for soil stabilization has been considered
Guidelines for Enzymatic Soil Stabilization 375

effective, but the drawback is the excessive cost and environmental impact during
the cement production process.
Fly ash (FA) is also considered as one of the most useful and sustainable addi-
tives for soil stabilization applications due to its unique characteristic of acting like
a cementitious material in the soil matrix. Although FA is solely incapable of densi-
fying the soil material, it can react chemically to achieve the cementitious compound
in combination with only a small amount of activator to improve the strength of the
soil. However, the implementation of FA for soil stabilization applications is limited
by the water content in the soil material. Thus, in order to achieve the optimum benefit
from the addition of FA, the water content in the soil matrix must be at the minimum,
and then dewatering is required to maintain the optimum moisture content of the soil
material [21]. On the other hand, research has shown that the sulfur content in FA
potentially forms expansive soil and, in the long term, leads to reduced strength and
durability of the soil [11].
In addition to the common stabilizing approaches outlined above, biotechnological
products, such as enzyme-based products, are currently being used as innovative
products for improving weak soils. Researchers recently reported on the Eko-Soil
enzyme (Eko Enviro Services) for soil stabilization and showed sustainable benefits
for the stabilization of expansive subgrades [22]. Pooni et al. [23] used the identical
enzyme product from the optimum enzyme content obtained by Rintu et al. [24] and
evaluated the hydraulic influence and sustainable benefits of the application to an
expansive clay material. The findings revealed that the addition of enzyme resulted
in an increase in California Bearing Ratio (CBR) by 58% under soaked conditions
due to densification effects. They further showed that the volumetric size of the
micropores in enzyme-stabilized samples was drastically decreased, in comparison
with raw soil, due to the improved density obtained from enzymatic stabilization
[23]. As a result of the high stabilization performance and low costs, enzymatic
stabilization using Eko-Soil has been adopted in field applications for unpaved road
construction [25].
Here we provide guidelines that demonstrate the methodologies of conducting
subgrade soil stabilization for road pavements, typically using the enzyme-based
soil stabilization technique. This guideline can be used as a tool for enhancing soil
conditions by providing stronger, impervious, and cost-effective subgrade soils under
rigid or flexible pavements. In addition, an impervious capping layer over reactive
soils can be productive by providing a safe all-weather pavement that is dust resistant
and requires minimal maintenance over the long-term life cycle. To ensure these
requirements are obtained, we performed the current study to identify the steps
that need to be taken by geotechnical practitioners in terms of safety instructions,
laboratory investigations and evaluation of site performance in compliance with
specifications.
376 B. O’Donnell et al.

2 Pavements

The safety and lifespan of pavement structures significantly depend on the condition
of the subgrade soil, because of the significant economic impact of frequent repairing
[26] if mechanical properties are weakened. Some naturally occurring soils are suit-
able for compaction, forming a homogeneous material that is capable of supporting
commercial and residential infrastructure. For instance, weathered rock extracted
from the ground is generally a suitable fill material for infrastructure, typically when
used with cementitious products for stabilization applications [27]. However, there
are other materials that are not suitable as fill materials because of factors such
as the in-situ conditions, environmental conditions, and applied loadings during
the construction timeframe. The following aspects are important to consider when
selecting a fill material from the economic viewpoint.
• Clay material that has a high plasticity index (PI), and is therefore a reactive soil,
must be considered under strict moisture and density control when selected as a
fill material.
• After completion of compaction, there can be large particles within the soil that
potentially limits trenching or drilling of piers excavation, footings, services, and
driving of piles.
• Over-wet soil materials, which can be found mainly in low-lying areas, have the
potential to dry out sufficiently within a shorter time during the project lifespan.
• Large-size individual graded rock fill material has limitations on the breaking
mechanism during the compaction process, which leads to higher porosity within
the soil that subsequently creates pathways for fine particles to migrate within the
soil matrix.
• The salinity of the soil, in relation to aggressive chemical products or unwanted
(polluted) soils.
• Soil carbonation leads to the occurrence of acidic products.
• The soil material has a PI between 6 and 20 with even graduation of particle size
from 20 mm minus.
There are soil types that cannot be used as fill material, due to high contamination
or undesirable performance of the soil material, and must be removed or transported
and used elsewhere. Unsuitable fill materials may include:
• Organic soils that contain severely root-affected material and peat, where these
soils are more likely to be the top soils.
• Soil materials contaminated through past site usage, which may contain toxic
substances or soluble compounds harmful to water supply or agriculture.
• Soil materials containing substances that can be dissolved or leached out in the
presence of moisture (e.g., gypsum), or are susceptible to volume change or loss
of strength when disturbed and exposed to moisture (e.g., sandstone), unless these
matters are specifically addressed in the design.
Guidelines for Enzymatic Soil Stabilization 377

• Silts or related materials that have the deleterious engineering properties of silt.
• Soil materials with properties that are unsuitable for forming structural fill.
• Fill that contains wood, metal, plastic, boulders or other deleterious material.
Furthermore, for road pavements, it is also important that high attention be given
to the shoulders abutting the pavement. Special treatments must be considered to
ensure that microbial metabolism and activities of all the substrates are minimal
under the permeable pavement, because they can lead to an increase in the total
organic carbon content and incremental water content, which, potentially, lower the
performance of the structure [28, 29].
Attention must be given to any erosion, which is the process by which the soil
is worn away by water or wind and sediment is produced. Some soils are more
susceptible to erosion than others, depending on their mechanical, chemical, and
physical properties and the terrain [19, 30]. Effectively, erosion can be increased by
several factors, including in high rainfall areas at the points where the overland flow
is concentrated; where roadside activities such as vehicular traffic and maintenance
practices increase the potential for erosion and sediment production or where any road
construction interrupts the natural topography or drainage flows. When the runoff
discharges turbid water into waterways, it can cause serious environmental harm, by
reducing the sunlight penetrating the waterway [29], which can affect the growth of
plant life and reduce the capacity of visual predators (e.g., fish and birds). Moreover,
road dust can have a significant detrimental effect on the environment, affecting adja-
cent crops, waterways, buildings, vehicle amenity, aesthetics and human health by
aggravating respiratory illness and road safety, through poor visibility and affecting
driver behavior [31].
The designed pavement for a road mainly depends on the type of road. Rural road
types are typically categorized as sealed, unsealed and stabilized. The sealed rural
road has a flexible pavement that is designed, as per geotechnical recommendation, to
include the addition of a top layer of bituminous concrete, asphalt or bituminous spray
seal. For example, sealed roads are usually formed by excavating and preparing the
existing ground (i.e., base) before placing crushed rock layers and a wearing course.
For urban roads or rural roads of significance, underground drainage, footpaths,
kerbing, and traffic management devices may also be considered [6, 31]. On the
other hand, for unsealed rural roads, the top surface of the road has no bituminous
layers but consists only of granular material (usually local gravel) or imported quarry
product [32]. Within Australia, there are numerous safety, economic, social and
environmental shortcomings regarding access for communities, and the extent of
such shortcomings is largely dependent upon the characteristics of each individual
road’s construction and traffic volumes. Significant regular maintenance is required
to ensure surface conditions do not change until the geometry and surface can be
improved to a safe acceptable level by the construction of an all-weather sealed road
[9, 33].
Soil stabilization is considered as an efficient method to ensure soil behavior
is within the required shear strength, permeability and compressibility parameters.
Various methods of soil stabilization have been implemented throughout the history
378 B. O’Donnell et al.

of subgrade soil stabilization, including mechanical and chemical methods [11, 34].
Soil stabilization through mechanical methods involves changing the soil mixture
by degrading and densifying the soil using compaction with heavy rollers, rammers,
and vibrational equipment and may sometimes involve blasting techniques for supe-
rior stability. Mechanical stabilization methods can be costly due to the requirement
for labor and specialized equipment, so soil stabilization using chemical additives
is becoming more common, using academic and practical engineering applications
to ensure soil stabilization and densification are obtained by mixing minerals or
biological additives [35]. In addition, chemical additives have the potential to reduce
the timeframe of the construction by using available construction equipment, which
is beneficial from the economic aspect. Stabilized road pavements are constructed
with one or more layers/courses mixed with an additive to bind the pavement mate-
rial [36, 37]. The preferred option for a conventional pavement is using the in-situ
subgrade soils and gravels as a subgrade. This is especially important for saving
natural resources, particularly where deficiencies in the existing in-situ gravels or
clays can be rectified by importing more suitable gravels or clays and mixing them
with a stabilizer additive to construct an unsealed pavement that is an environmental
friendly, cost-effective, impervious and strong road [38].

3 Construction Commencement

This section covers the specific working steps that need to be taken before
commencing the earthworks and soil stabilization process.

3.1 Construction Site Fencing

In the initial stage of construction, it is highly important that fencing around the
perimeter of the construction site is installed before any earthwork commences.
Fencing installation is one of the safest methods for identifying the boundaries of the
working zone because permission to enter the construction site is then only given to
authorized users, thereby preventing the public from entering the site and disturbing
the work performance while ensuring their protection.

3.2 Drainage, Erosion and Sedimentation Control

Before conducting the stabilization works, precautions must be taken to ensure the
earthworks will not cause siltation or erosion of adjoining lands, streams or water-
courses. Drainage, erosion and sedimentation controls should be installed before
the natural surface is disturbed. Sedimentation basins, stream diversion or other
Guidelines for Enzymatic Soil Stabilization 379

works may be appropriate in some environments or topographies. Careful plan-

ning is required to ensure both erosion and sedimentation controls are effective by
minimizing the area of disturbance and through progressive revegetation or redevel-
opment of the site. Wherever water may tend to accumulate, provision for temporary
drainage should be made by the contractor and care should be taken to guard against
scour during any part of the construction. All temporary provisions for drainage
should be installed to the satisfaction of the superintendent before stabilization and/
or pavement materials are placed. The cost of temporary drainage, unless otherwise
directed to be retained for use as catch or shoulder drains, is the responsibility of the
contractor. The location of each drainage line is determined on geotechnical advice
and with the approval of the superintendent.

3.3 Site Clearing

The site must be cleared (to the minimum extent required for the work) of all trees,
stumps and other materials unsuitable for incorporation in the works. The roots
of all trees and debris, such as old foundations, and buried pipelines are removed
to sufficient depth to prevent any inconvenience during subsequent excavation or
foundation work. The resulting excavations should be backfilled and compacted to
the same standard as required for subsequent filling operations. Disposal of cleared
combustible material may have to be off-site if clean air or bushfire regulations
prevent on-site burning.

3.4 Stripping

The area in which fill is to be placed and the area from which the cut is to be removed
are stripped of all vegetation and of such soils that may be unsuitable for incorporation
into fills, subject to density, moisture or other specified controls. Topsoil may need to
be stripped either as unsuitable material or as required for subsequent revegetation.
Extreme care needs to be considered to ensure that materials that will inhibit or
prevent the satisfactory placement of subsequent fill layers are not allowed to remain
in the foundations of the fills. Geotechnical assessment of the depth and quality of
topsoil or vegetable cover of the underlying soils and of the quality and depth of
the proposed fill may obviate the need for such stripping in some circumstances.
All stripped materials should be deposited in temporary stockpiles or permanent
dumps in locations available for subsequent re-use if required and where there is no
possibility of the material being unintentionally covered by or incorporated in the
earthworks.
380 B. O’Donnell et al.

3.5 Slope Preparation

Where a fill abuts sloping ground, benches should be cut progressively with each lift
as appropriate. It is unlikely that slopes flatter than 8:1 (horizontal to vertical) gradient
will require benching. The benches should be shaped to provide free drainage. The
boundary of cut-and-fill areas requires special consideration. All topsoil and other
compressible materials should be stripped prior to benching into the natural material
of the cut zone. The depth of the cut can vary depending upon the natural slope of
the ground, the nature and proposed end use of the fill and the equipment being used.

3.6 Foundation Preparation

The ground surface exposed after stripping should be shaped to assist drainage and
be compacted to the same requirements as for the overlying layers of fill. The surface
exposed upon completion of excavation works may also require preparation prior to
the fill placement proceeding. This will typically be the case when the subsequent
fill to be placed is for pavement construction or the base material of a project. In such
circumstances, it is necessary to loosen the exposed excavation surface to a certain
depth (depending on the soil conditions), then moisture-condition and compact this
loosened material. The depth to which this loosening is carried out should not exceed
that of the compacted soil layer above it. The degree of compaction achieved should
be consistent with the required subsequent filling operations unless design advice has
been obtained. In such cases a working platform generally of granular material, end-
dumped and spread in sufficient depth to allow the passage of earthmoving equipment
with minimal surface deflection, can provide a suitable foundation for subsequent
filling. Localized springs or seepages in the foundation area, detected during site
investigation for the work, should be noted and considered in the design. If such
problems are not detected before the work progresses, it is critical they be assessed so
that measures such as subsoil or rock rubble drains can be designed for incorporation
in the works.

4 Laboratory Test Investigations

This section elaborates on the required laboratory tests for evaluating the soil
condition and required stabilization measures prior to construction.
Guidelines for Enzymatic Soil Stabilization 381

Fig. 1 Representation of the clay particle entrapped by the enzyme strains

4.1 Enzyme Stabilization

An enzyme is a biological macromolecule that catalyzes biological reactions, found in

ribozymes (catalytically active RNA molecules) and some proteins (protein enzymes
that have the ability to catalyze specific biochemical reactions), which are capable
of initiating a biochemical reaction. The process involves changes (both formation
and breakage) in chemical bonds. This method creates an interaction between the
enzyme and bacterial strains within the soil product and is a replication of the natural
construction of termite mounds [23]. Enzyme soil stabilization is based on reducing
surface tension in soil particles through an ionic reaction, hence incrementing the soil
compaction conditions. After the absorbed water is reduced through the compaction
efforts, the soil particles agglomerate and as a result of the relative movement between
particles, the surface area is reduced and less absorbed water can be held, which in
turn reduces the swelling capacity. Figure 1 shows clay particles when entrapped
with the enzyme strains.

4.2 Laboratory Results Illustrating the Enzymatic Soil

Stabilization Mechanism

The behavior of the enzyme has been established through various studies. Laboratory
results have indicated that the addition of enzyme causes water content reduction,
as shown by different critical tests including FTIR (Fourier-transform infrared spec-
troscopy), SEM (scanning electron microscope) and microtomography (µ-CT). The
addition of the enzyme to the soil has shown a marginal reduction in the intensity
of the interlamellar water region due to the reduced affinity by the enzyme product
382 B. O’Donnell et al.

(a) [40]

Raw Soil Stabilized Soil

(b) [41]

Raw Soil
Stabilized Soil

(c) [24]

Fig. 2 Laboratory results for the mechanism of enzyme-treated soil. a FTIR evaluating the enzyme
base material in the soil and densification determination; b µ-CT analysis of the compactness of
treated soil compared with control soil material; c SEM images demonstrating the lower void
content and clay aggregation with the addition of enzyme product (Right) compared with control
soil material (Left)
Guidelines for Enzymatic Soil Stabilization 383

(Fig. 2) [23, 39]. Porosity analysis and clay microstructure via SEM images indi-
cate that enzyme-stabilized soil samples potentially show reduced permeability and
increased mechanical strength as ingress of water is restricted and density is enhanced
through clay aggregation [3]. Pooni et al. [40] showed by µ-CT analysis that the addi-
tion of enzyme can reduce the porosity from 2.67 to 1.44% (i.e., 46.07% reduction
in pore volume). Effectively, the enzyme mechanism is increasing the density with
decreased affinity for water.

4.3 Testing Techniques

Subject to the scale of the project, difficult conditions may be expected, and it is not
envisaged to relax the test frequencies specified herein; in some cases, more frequent
testing may be required. These testing frequencies relate to acceptance on a ‘not one
to fail’ basis.
In order to obtain optimized performance of the stabilized road, it is recommended
to estimate the performance of the stabilized soil through a comprehensive test plan
in the laboratory prior to the field application. This is mainly due to the performance
of the stabilized soil (i.e., treated road pavement), which is governed by the in-situ
soil type and its condition. Figure 3 shows the recommended laboratory tests that
can be conducted in the application of enzymes to stabilize pavements. The proposed
tests will facilitate determination of the suitability of in-situ soil in stabilizing the
pavement, as well as obtaining the optimal amounts of enzymes that will result in

Fig. 3 Typical soil stabilization using different testing techniques at different stages
384 B. O’Donnell et al.

the expected road performance. It is to be noted that oven-drying is not allowed in

any of the tests prescribed in this plan. The soil needs to be air-dried, or lime could
be used as a drying agent prior to testing.
The proposed test methods can be conducted in three stages. In stage 1, initial
tests are proposed to obtain the description and physical (and chemical) properties
of the raw soils (possibly in-situ/natural soil). These tests include gradation, miner-
alogy, chemical, Atterberg, compaction and permeability tests. Having obtained the
description and properties of the raw soils, samples are prepared at maximum dry
density and optimum moisture content and left to cure for 72 h before mechanical
testing (Stage 2 in Fig. 3). Should the natural soils be over-saturated, then add 3%
lime mixture to the soil, and allowed to dry for 3 days, repeat the procedure until
the natural soils are dried. This procedure can be used in the field to obtain the
desired result. Furthermore, every mechanical test can be repeated at 4 and 7 days
after curing to characterize the time-dependent strength gain of stabilized soils. Such
mechanical tests will be helpful to assess the performance of the pavement material
in accordance with relevant road standards. In addition to the tests noted in Stage
2, a permeability test of the stabilized samples is also recommended to determine
whether the stabilization has improved the permeability of the mix. If the material
is shown to be unsuitable, the strength of the stabilized mix could be improved with
the addition of cement. The proposed series of mechanical tests on the stable mix
with cement and enzymes will be able to ascertain the suitability of in-situ soil as a
pavement material.
In order to proceed with testing using the additive/s for stabilization, initially
samples should be taken from the field for each different soil type that is observed.
In the case of a road, samples should be taken every 300 m, or less if there appears to
be inconsistency within the sampled area. Moisture content within the sample bore
is also a good measure for indicating a change in soil characteristics. In a remote
area, geotechnical maps will assist in the location of samples if required. The initial
tests are conducted to evaluate the mechanical strength of the soils. Figure 3 shows
the types of tests that need to be conducted when enzyme is used as an additive
for stabilization. Soil stabilization enhances an in-situ soil to support loads well in
excess of those normally possible under natural conditions. Geotechnical analysis of
soil samples collected from prospective stabilized sites must be carried out.
In all practical applications where moisture enters a pavement section, the perme-
ability of the designed pavement mix must be tested in the laboratory to ensure
specification requirements are either met or exceeded. It is recommended that the
soil or soil with additives being considered for use at the construction site is tested
to ensure that it conforms with the ideal specifications when using Eko-Soil. These
specifications include the following.
1. Minimum 18% non-granular cohesive fines passing the 0.075 mm screen or a PI
≥ 6%. These fines will react best with Eko-Soil.
Test method: ASTM D-11, D-422 [41], AS1289.3.6.3 or similar particle size
analysis.
2. PI of 6–15 is the ideal soil condition when using Eko-Soil.
Guidelines for Enzymatic Soil Stabilization 385

Test method: ASTM D-4318 [42] Atterberg Limits AS1289.3.9.1 [43].

3. pH 4.5–8.5 is the most satisfactory. Soils with low pH may be amended with
crushed limestone. Alkaline soils may be treated with a cheap acid such as sodium
acid sulfate. Eko-Soil is classified as a naturally produced weak acid with pH
4–5.
Rintu et al. [44] used Eko-Soil to evaluate the effectiveness of different contents
of Eko-Soil enzyme on fine-grained field soil. The authors found that the optimum
enzyme stabilization is achieved with a diluted additive of 1% to the weight of soil
in a diluted mass ratio in water of 1:500. The performance at an optimum dosage
showed an increase in CBR value for the tested soil by 500% in comparison with the
control and a reduction in water demand and improved dry density.

5 Field Procedure to be Followed

To assess the quality of materials and workmanship provided on a project, regular

inspections and testing are required in the field at suitable time intervals. However,
site construction should not rely on test results alone; where good supervision is
essential for inspection measures such as test rolling. Such inspections should be
carried out by experienced and knowledgeable users in earthworks.
Having conducted laboratory tests as outlined in Sect. 4, careful measures should
be put in place in the field to apply the stabilization. Enzyme stabilization will prove
less than effective if the following minimum procedures are not followed.
• Ensure a soil analysis is undertaken on the subject soils and proposed additives
before treatment begins.
• During initial preparation, it may be necessary to add further plain water (Fig. 4a)
to bring the soil close to optimum moisture level for best compaction, or to aerate
and allow it to dry. Proper engineering supervision is essential (Fig. 4b).
• Add the appropriate quantity of enzyme additive to the water truck after the truck
is filled with water; as the quantity of enzyme depends on the in-situ soil type
and its condition. The recommended tests proposed in Fig. 3 will facilitate the
designer determining such optimized amounts of enzyme.
• After the soil/gravel and any other cementitious additives is mixed, the enzyme-
treated water is dispensed by the tanker and the material thoroughly mixed using a
purpose-built roto-mill (Fig. 4c). In most cases, adequate mixing can be achieved
using a single pass.
• The mixed material should then be spread and shaped before compaction by
rolling in maximum layers of 250 mm. For fill depths, refer to engineering plans.
• The rollers must make enough passes to ensure adequate compaction is taking
place. A vibrating pad foot roller is most effective for this application. On the
final passes, when the tire tracks no longer show, the vibrator should be turned
off to prevent excessive surface cracking caused by rapid drying. The top surface
should be rolled until it shows a uniform sealed appearance. The final stages of
386 B. O’Donnell et al.

Fig. 4 a For an optimum moisture content mix water lightly through the broken soil. b Ripping
and shaping an existing road base. c Adding stabilizer using a mixing machine and water tanker in
tandem. d Compaction by 16-tonne vibrating roller

rolling should be performed by a 16-tonne smooth drum roller (Fig. 4d), followed
by a pneumatic (rubber) roller to assist in drainage and preventing ponding of
water on the surface.
The ideal curing time for a 250-mm pavement depth would normally be 72 h.
Light traffic may be permitted, as soon as tyre tracks are not visible from the surface.
Light rain or high humidity will increase curing time. Application of enzyme is not
to be undertaken during rain unless otherwise approved by the superintendent. Once
field stabilization is completed, tests (Fig. 5) can be performed to ascertain field
efficiency.

Fig. 5 Different testing techniques that can be conducted on the basis of field samples. CBR,
California bearing ratio; UCS, unconfined compressive strength
Guidelines for Enzymatic Soil Stabilization 387

5.1 Unconfined Compressive Strength

Unconfined compressive strength (UCS) is an effective test in ensuring pavement

strength is within the range of 0.5 and 1.5 MPa or if the UCS value exceeds such a
range the pavement is classified as a rigid type and therefore subjected to cracking. To
assess the quality of the material through mechanical testing methods for areas of a
minimum of 1500 m2 , such as subdivisions, large industrial lots, road embankments,
etc., the following requirements must be considered in corresponding different testing
locations within the field.
Not less than:
• 1 test per layer or 250 mm thickness per material type per 2500 m2 or
• 1 test per 500 m3 distributed reasonably evenly throughout full depth and area or
• 3 tests per visit, whichever requires the most tests.
• Confined operations: 1 test per 2 layers per 50 m2 .
To ensure that enzyme and cementitious modified materials (clays mixed with
local gravels, cement, and enzyme) are achieving desired strengths in the order of
0.5–1.5 MPa (at 7 days), it is recommended that a minimum of two molded samples
of these materials be re-compacted in the laboratory and allowed to cure for 7 days
for the purpose of obtaining UCS values.

5.2 Deflection Test

Field deflection testing can be used to evaluate the strength of the soil. This test is
highly recommended when working with large platforms or pavements (Table 1).

5.3 Plasticity Index

The PI needs to be 6–15 for a minimum of 18% non-granular cohesive fines passing
the 75-micron sieve.

Table 1 Results of deflection testing of both treated and untreated pavement sections: Harvey
Norman/Ikea site, Springvale, November 2008
Pavement section Characteristic Tolerable deflection
deflection (mm) (mm)
In-situ stabilized material as found (chainage 1.486 1.5
0–0.210 km)
Unstabilized existing subgrade material as 2.157 1.5
found (chainage 0–0.220 km)
388 B. O’Donnell et al.

5.4 Maximum Density/Optimum Moisture

This test will determine the amount of enzyme required to obtain maximum results.
The test method follows ASTM D-1557 [45] modified proctor. Typical enzyme rate
(based on Eko-Soil) from field experience can be 1–1.5 L of the enzyme to 30 m3 of
compacted pavement in general. Moisture content to achieve maximum compaction
should be 1–2% below optimum. In the field, the moisture content is determined by
hand squeezing of the mixed material. If it crumbles, then add more water and retest.
If it has an excess of water, allow drying.

5.5 Bearing Strength or CBR

The bearing strength, or CBR, is an effective test for determining the bearing strength
of the soil. However, the laboratory CBR may not conform to or replicate the field
bearing strength because the compacted CBR samples must be allowed air dry for
72 h before submerging in water. CBR tests should be undertaken on cementitious
modified materials for each 2nd day’s production or every 2500 m3 whichever is the
lesser. Desired CBR values of cementitious modified materials should be >15%.

5.6 Permeability

With the addition of enzymes, the reduction in moisture directly affects the design
of the structural section of the pavement. Reductions of up to 100-fold are achieved.
The test method can be performed in accordance with ASTM D–5084 [46]. The
representative values of relative permeability of the different types of soils are shown
in Table 2. To ensure that the cementitious materials maintain the desired permeability
of <5 × 10–8 m/s, preferably <5 × 10–9 m/s, permeability testing should be undertaken
on cementitious modified materials for every 4th day’s production or every 5000 m3
whichever is lesser. The permeability for the enzyme treated soil is approximately
10–8 –10–11 m/s.

Table 2 Representative soil

Soil Permeability coefficient
permeability for various soil
types (obtained from AS1547 Sand 10– 5 m/s
[47]) Silt 10– 6 m/s
Clay <10– 7 m/s
Guidelines for Enzymatic Soil Stabilization 389

5.7 Field Density

Methods for the determination of field dry density are as described in Sects. 5.8–5.11
below.

5.8 Direct Density Test

In order to perform the direct density test, various standards describe specific methods
of conducting the test and evaluating the analysis including: AS1289.5.3.1 [48],
AS1289.5.3.2 [49], AS1289.5.8.1 [50] and AS1289.5.3.5 [51].

5.9 Indirect Density Test

This test provides an empirical measure of achieved density by measuring another

engineering property, principally shear strength, and may be used to further validate
density; however, the method of direct testing will govern acceptance.

5.9.1 Establishment of a Reference Density for Calculation of Relative

Compaction

To permit relative compaction to be calculated, it is necessary to establish a laboratory

reference density. Procedures for establishing such reference densities have been
developed empirically over many years and standardized with test procedures of
AS1289.5.1.1 [52], AS1289.5.2.1 [53], AS1289.5.7.1 [54], and AS1289.5.5.1 [55].

5.9.2 Sample Selection for Reference Density

For routine “compaction” testing, the sample for determination of laboratory refer-
ence density should comprise either the material recovered from the field density
determination, (see AS1289.5.3.1 [48]) or from the volume of material considered
in the field density.
For cement-modified stabilized materials, including enzyme, the reference density
may vary with time but the laboratory compaction should still be carried out on
material that has been mixed and compacted by on-site purpose-built machinery.
The density tested and re-compacted in the laboratory must be conducted as soon as
practicable to ensure minimum curing has occurred. For granular materials, including
pavement base and sub-base materials that have been manufactured from a hard rock
390 B. O’Donnell et al.

source under controlled conditions, consideration may be given to providing an

assigned value as further discussed within the procedures of AS1289.5.4.2 [56].

6 Construction Applications

This section describes various applications of enzymatic soil stabilization.

6.1 Flood Mitigation

Enzyme-stabilized pavements in the field are constructed in a similar manner to

conventional stabilization methods, using water tankers, motor graders with rippers,
and a 16-tonne vibrating steel roller. Additives should be verified for weight/square
meter to ensure that the quantity of additive meets specifications. For speed of
construction and to ensure a homogeneous mix, self-propelling and towing mixing
machines are recommended. Both types of machines can be fitted with a computer
water feed system and in a recent project that we were involved in, 7000 m2 of
300-mm deep stabilization were completed in an 8-h day. Should rock or cement be
required as part of the geotechnical design as an addition to the pavement mix, this
should be placed and mixed before the stabilization process.
Cross fall should be at least 3% and side drains should be cut into the pavement
(Fig. 6). Just as the pavement is losing its plasticity, a skim of one-size crushed
rock (10 mm) should be rolled into the top layer to provide a skid-resistant surface.
Light traffic can be allowed to traverse the pavement as soon as the tyre marks of
the stabilization equipment disappear. Full curing is achieved in 72 h after which the
road may be opened to all traffic.
Stabilized pavements subject to inundation must be tested in the laboratory to
ensure permeability either achieves or exceeds specification. Enzymes, or an enzyme
blended with up to 3% cement, in both laboratory and field testing, can achieve the
required specification but will have an advantage over clay pavements due to their

Wearing Course

Water Flow

Stabilised
Surface
Non-Stabilised Natural Soil

Fig. 6 Typical cross-section of road pavement subject to flooding

Guidelines for Enzymatic Soil Stabilization 391

residual tensile strength, which prevents cracking of the pavement once the water
recedes; hence, stabilized pavements must envelop the road or dam surfaces.

6.2 Embankments and Slopes

Enzymatic soil stabilization can be used successfully in the stabilization of embank-

ments and slopes. Preliminary work is presently underway on problems with some
highly plastic soils (e.g., colloid clays that are highly expansive). These soils contain
fatty clays and fine gravel or sands and silts with a PI ranging from 45 to 80.
In making calculations for enzymatic stabilization application for these types of
soils, the ratio of cohesive fines is different from the standard gradation specifica-
tion. Normal gradation is calculated on 18–30% cohesive fines. Therefore, correlating
these expansive clay results in a greater number of cohesive materials being present.
The enzymatic composition works on the molecules within these expansive clays
but not the other materials present. The standard application of 1 L per 30 m3 of
compacted pavement must be increased to compensate for the increased amounts of
cohesive fines. The standard rate for high amounts of expansive soil is calculated
using the same ratio as for a standard application. As an example, the rate of appli-
cation for soils containing 48% cohesive fines would be 1 L per 15 m3 of compacted
stabilized slope pavement. However, this has to be determined from adequate tests
as explained previously.
Recommended thicknesses for slopes are:
• Up to 30 degrees, 150 mm (Fig. 7)
• 30–50 degrees, 150–225 mm

Fig. 7 Erosion simulation equipment used in modified version of the tests [57]
392 B. O’Donnell et al.

• 50–65 degrees, 250–350 mm

• ≥65 degrees must be used in conjunction with soil nails.
Compaction has always been difficult in slope construction. In recent years, new
mechanical attachments have been introduced for compaction of trenches, slopes, and
embankments. Compaction load is applied by the backhoe. A filtration membrane,
mostly likely sand, must be placed between the slope and the pavement. Weep holes
are placed in the pavement at locations as directed by a civil engineer experienced
in such methods. Recent compaction equipment introduced into slope stabilization
enhances the ability to compact slopes at much steeper angles than was previously
possible.

6.3 Soil Nailing

Soil nails have been used for many years in slope stabilization. The normal soil nail
is constructed using concrete and steel reinforcement, which can be an expensive
additional cost in embankment stabilization projects.
It has been observed that concrete soil nails will catch water on their upper surface,
allowing water to penetrate around the soil nail, loosening adjacent soil and aggregate
materials and resulting in loss of support of surrounding materials.
Enzymatic-stabilized soil nails are economic alternatives. As with concrete soil
nails, they can be pre-manufactured or constructed on-site, and similarly, it is recom-
mended that enzymatic soil nails be reinforced with vinyl-coated reinforcement rods,
which should be ≈30% below the bottom of the soil nail for anchoring into the in-
situ material. Enzymatic soil nails are recommended to be pre-formed and pressed
into pre-bored holes that are slightly undersized. It is also recommended that enzy-
matic soil nails be placed on the slope after being dampened with a 1:10,000 mist of
enzymatic composition and water.
An alternating pattern of rows should be used. The spacing of the nails and the
rows should be approximately 1.5 times the diameter of the soil nails. A civil engineer
experienced in slope stabilization should design the use of soil nails. Construction
and replacement of soil nails should be performed in favorable construction climatic
conditions by avoiding freezing and wet weather conditions.

6.4 Enzymatic Composition Blocks (Hollow Blocks/Cement

Blocks) and Bricks

The construction of cemented crushed shell blocks can be achieved with the aid of
enzymatic composition. A reduction of ≈5% in the use of water will be realized and
a reduction of mold breakage rate of between 35 and 50% can be achieved. This is
an economical saving for any manufacturing process.
Guidelines for Enzymatic Soil Stabilization 393

Enzymatic-stabilized soils have been used to construct solid construction blocks

designed for the construction of interior walls and low retaining walls. They were
manufactured using small, fractured gravels and cohesive fines within standard grada-
tion specifications, and have proved to be exceptionally strong, durable, and easily
constructed with a hand-operated lever-style press and a steel mold.
More efficient methods are used to manufacture enzymatic-stabilized blocks/
bricks, where facilities are available. Hydraulic power units have the capacity to
produce 6000–8000 blocks per shift per machine.
Mortar for joining the enzymatic-stabilized blocks and bricks should be a combi-
nation of the makeup of the blocks/bricks (without the larger aggregate) and applied
as normal mortar to all connecting surfaces. Curing times of the mortar joints will
be in the order of 24 h. The utilization of enzyme is currently under investigation for
unfired brick fabrication.

6.5 Working Platform

A typical example of soil stabilization by utilizing enzyme product in the field was
undertaken on an area of 65,000 m2 for a whitegoods site in Australia (Fig. 8). The
pavement was 150 mm of in-situ fill material and 150 mm of recycled material recov-
ered from demolished site buildings. The testing was conducted in the laboratory and
the field as recommended here. The compaction density of the subgrade achieved
results >100% while the strength of the subgrade, in field testing on this site achieved
CBRs of ≈80% and permeability of 10–14 m/s.
The process for working platforms also applies to roads, large industrial/
commercial building sites, and most civil infrastructure sites. They are useful for
subgrade improvement of over-reactive clays to stabilize the subgrade moisture and
to limit differential movement in the subgrade. Figure 9 shows the comparisons of
road pavement based on unstabilized and stabilized base in the same locality.

6.6 Reclaimed Materials

Testing performed for the Hong Kong University recommended that the sludge
retrieved from the Hong Kong Harbour be used as a working platform for the Hong
Kong Housing Department. The material was delivered to the site and dried by
using 3% lime mixed into the sludge and left to dry for 3 days. A pavement mixture
consisting of 150-mm of dried reclaimed material, 150 mm of 19-mm of recycled
concrete, 3% cement (due to the variation of soil consistency) and 1% of enzyme
stabilizer produced a CBR of 80%.
394 B. O’Donnell et al.

Fig. 8 Using enzyme as an additive for soil stabilization of a construction site consisting of 65,000
m2

Fig. 9 a Cracked seal: base not stabilized. b Good seal: enzyme-stabilized base

6.7 Water Containment

Ponds were constructed at Warrnambool Airfield. The method of construction was

to build a base with an enzyme additive, build up the pavement mix for the walls
then stabilize the existing base. The stabilized material was then moved from the top
of the pond walls to the base on the 1 in-3 slope using vibrating rollers.
Guidelines for Enzymatic Soil Stabilization 395

7 Conclusions

Based on the procedures and criteria from different standards and guidelines, the
utilization of additives for soil stabilization is more effective when compared with
the mechanical methods traditionally used for soil stabilization. Soil stabilization
through enzymatic bonding, although highly dependent on the soil material, requires
soil materials to biochemically react with the additives in order to obtain an effec-
tive stabilization. An excellent understating of the topography and geology of the
construction site is imperative.
Consequently, different mechanical and chemical composition tests must be
undertaken on the soil types available from the site for each and every combination
of chemical additives that will potentially be used for stabilization of those materials.
Specifically, when using enzyme products, the soil properties need to be carefully
evaluated before any stabilization commences. Once the laboratory results are evalu-
ated, fieldwork can be performed based on the optimum additive content determined.
Subsequently, to maintain the safety of infrastructure, comprehensive construction
management of the site must be considered before conducting any construction work.
Thus, monitoring the field test in accordance with available guidelines, standards,
and contractor management protocols is equally important and the field tests are
undertaken regularly throughout the lifespan of the construction work to ensure that
construction is performed using the highest quality standards and with minimum
geotechnical issues.
Stabilization is a science, so a suitably qualified engineer must design and sign off
the pavement stabilization construction to ensure the pavement has met or exceeded
standards of permeability, strength and density. It is to be noted that the information
provided in this document is for guidance only and should not be used without
required tests and suitable evaluations as detailed herein.

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Deterioration Modeling of Concrete
Bridges and Potential Nanotechnology
Application

H. Tran and S. Setunge

Abstract Management of aging concrete bridges with limited resources can be a

challenge for state authorities. Deterioration modeling of concrete bridges at the
component level is essential to optimize maintenance actions and ensure the safety
and serviceability of the bridge network. In this study we examined the Level 2
visual inspection data of a concrete bridge’s components collected over 4–5 inspec-
tion cycles with the objective of predicting deterioration of components and the
bridge’s life cycle. With the increasing application of nanotechnology to increase the
mechanical properties and durability of concrete material for bridge structures, the
deterioration of nano-based concrete could be significantly different from conven-
tional concrete. A range of deterioration prediction methods, including determin-
istic models and stochastic models, were examined to understand the validity of
the different methods in predicting the deterioration of bridge components made
of conventional and nano-based materials. A case study with a demonstration on a
concrete open girder was investigated with regard to linear regression models and
the stochastic Markov deterioration model. The outcomes can be used to support
future study on the performance of conventional and nano-based concrete materials
and their lifecycles in the asset management of bridges.

Keywords Bridges · Deterioration · Nanotechnology · Markov model

1 Introduction

Bridges are important assets of the transport infrastructure network that play a crucial
role in ensuring the well-being of populations and the economy. Concrete is the
popular material used in bridges, but after many years of service, concrete bridge
structures are aging and deteriorating, which becomes a hazard to traffic users in the

H. Tran (B) · S. Setunge

Civil and Infrastructure Engineering, School of Engineering, RMIT University, Melbourne, VIC,
Australia
e-mail: huu.tran@rmit.edu.au

© The Author(s) 2023 399

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_41
400 H. Tran and S. Setunge

event of structural failure or falling apart. The causes of concrete deterioration have
been intensively studied, such as the carbonation of concrete and chloride attack [1,
2]. Contributing factors to deterioration include traffic volume, exposure to corro-
sive soil and airborne chloride near the coastline and acidic gases (including airborne
carbon dioxide, nitrous and sulfurous oxides) in urban area. The deterioration process
can be single, such as concrete creep, or combined processes, such as stress corrosion
cracking in steel, and coupled with random damage events such as flooding and earth-
quake. The search for better concrete materials against natural and man-made hazards
is ongoing. Recent advances in nanotechnology and its application in concrete bridge
construction can help not only ensure safety and serviceability throughout service
life but also cost-effective maintenance, rehabilitation and replacement (MRR).
Nano-concrete has been developed in the past decade to improve key characteris-
tics of normal concrete such as tensile and compressive strength, anticorrosion and
durability [3, 4]. The efficiency of the nano-concrete depends on its density, which
can be maximized by minimizing the particle gaps. The optimal particle gap within
the concrete mass can be achieved by incorporating a homogeneous gradient of fine
and coarse particles in the mixture. In this regard, nanomaterials such as nano-silica,
nano-graphite platelets, carbon nanotubes, graphene, nano-titanium dioxide and nano
clay have been used to reinforce cementitious composites (cement paste, mortar, and
concrete) [5]. They can be highly effective because, with their extremely small size,
nanomaterials can fill the voids between cement and silica fume particles, leading to
higher level of compaction and generating a denser binding matrix. This high level
of compaction of concrete particles can significantly improve both the durability and
mechanical properties of the nano-concrete. Several studies have reported signif-
icantly improved properties of the nano-concrete. For example, the use of 0.02%
graphene oxide in ultra-high performance concrete can increase its strength charac-
teristics, such as compressive, tensile, and flexural strength, up to 197%, 160%, and
184%, respectively [5]. Kancharla et al. reported that replacing 0.5 and 1.0% cement
with nanosilica showed good improvement in bending strength of 7.8 and 15.7%,
respectively, in the crushing stage and slight improvement in bending strength of
0.42 and 1.26%, respectively, in the failure stage [6]. Mostafa et al. found that nano
glass waste can increase the bending strength of ultra-high performance concrete up
to 1.5-fold if it is added at 1% [7].
With the increasing use of nano-concrete, it is useful to compare deterioration rates
between traditional concrete and nano-concrete over service life. Such a comparison
is useful for decision making on the use of nano-concrete in bridge engineering
and to provide better knowledge for managing the maintenance and life cycle costs
for current and future use of the nano-concrete. One method is to use deterioration
models to compare the deterioration rates. Deterioration models can be divided into
knowledge-based, data-driven and mechanism-based models. Few bridge manage-
ment agencies use knowledge and experience to determine future deterioration from
inspected defects [8]. The data-driven models use inspection data to predict future
deterioration [9–11]. The mechanism-based model is based on deterioration mecha-
nisms such as corrosion of reinforcing steels [1, 12]. The physical models are consid-
ered more advanced and accurate but they require intensive and detailed data, which
can be costly and difficult to obtain. Therefore, they were not selected for this study.
Deterioration Modeling of Concrete Bridges and Potential … 401

The deterministic linear model is well known for its ease of use and implementation,
but is criticized for failure to capture the uncertainty of the deterioration process,
which was addressed by the stochastic Markov model in this study.
We aimed to investigate the deterioration rate of traditional concrete as the bench-
mark for future study of deterioration of nano-concrete used for bridge compo-
nents such as bridge girders and the bridge deck. For traditional concrete, visual
inspection data of bridge components using traditional concrete are available and
collected from the bridge agency for deterioration modeling. However, such data
are not available for the nano-concrete. Therefore, scenario analysis was conducted
for the deterioration rate of the nano-concrete based on its reported performance in
public literature to assess its potential economic benefit. The outcomes of this study
will demonstrate the benefit of condition monitoring and data collection for deterio-
ration modeling traditional concrete for expanding the application of nano-concrete
in bridge engineering.

2 Case Study

VicRoads is the registered business name of the Roads Corporation in the State
of Victoria, Australia (www.vicroads.com.au). It is a Victorian statutory authority
established under the Transport Act 1983 and continued in the Transport Integra-
tion Act 2010. One of VicRoads’ core services is to plan, develop and manage the
arterial road network, including roads, bridges, culverts and traffic signs. The bridge
management of VicRoads is supported through its computerized database, which
contains basic data of 6207 bridge structures including bridges, culverts and tunnel
with their attributes such as number of spans, length and width. The oldest structure
was built in 1899. The database also stores 26,1324 records of Level 2 inspections as
of 2017. The first recorded inspections in the computerized database began in 1995.
The VicRoads’s inspection practice is published in an open-access inspection
manual, which basically has three levels of inspection [13]. Level 1 is considered
a screening inspection with a maximum interval of 6 months. Level 2 is a routine
inspection with a typical interval of 2 years and Level 3 is an in-depth inspection for
special cases. The VicRoads’ inspection manual describes the breakdown of bridge
structures into superstructures and substructures and into structural elements such
as piles, decks and bearings. Each bridge element is coded by its function and one
of five types of material (i.e., timber, steel, in-situ concrete, precast concrete and
others); for example, the 2C mean open girder/stringer with the concrete material.
The Level 2 inspection provides q condition rating for individual elements based
on their inspected defects as per inspection guidelines. The condition rating is a
1–4 scale, with 1 being good and 4 being worst. The condition state percentage
distribution is used to provide inspection reports. For example, an inspection report
of [90% 10% 0% 0%] of 2C means 90% of the open girders in condition 1, 10% in
condition 2 and 0% in conditions 3 and 4. The case study dataset of Level 2 inspection
402 H. Tran and S. Setunge

data from 1997 to 2017 was used for deterioration modeling of bridge components
in this study.
A collaborative research project between VicRoads and RMIT University with
support from the ARC Nanocom Hub was established to develop deterioration models
using Level 2 inspection data for bridge structures. The outcomes of the collaborative
project could provide justification for future state funding of bridge management and
also provide support for more effective and efficient asset management programs.

2.1 Deterioration Models

2.1.1 Linear Regression Model

Equation (1) shows the linear relation between bridge condition (output) with age
(input) [9]. More input factors can be added into Eq. (1) in a similar manner.

Y = a0 + a1 ∗ age + ε (1)

where a0 is a constant coefficient, a1 is slope coefficient of age (year) and E is the

measurement error assumed to have zero mean and independent constant variance.
The interpretation of Eq. (1) is that (a) for the case of 1 input and 1 output, the
relation is a straight line; (b) for the case of ≥2 inputs, the relation is a plane and
hyperplane; (c) a linear relation means a constant deterioration rate over time; (d)
measurement error means that for the same values of the inputs, every time the output
is measured or assessed, the output value might not be the same due to equipment
sensitivity or human subjectivity; and (e) deterioration increases with age, meaning
the unchanged condition is not valid.
For calibrating or estimating the model coefficients, the least square method is
often used to minimize the error term between the observed output and model output.
The predictive performance of the linear model is often assessed using the coefficient
of determination R2 between observed values and model outputs.

2.1.2 Markov Model

The Markov model is based on the stochastic theory of Markov chain [14], which
describes a system that can be in one of several defined condition states at any time,
and it can stay still or move to another state at each time step with some transition
probabilities over time. This theory is well suited for observation of Level 2 condition
data of bridge elements because the snapshot inspection reveals the condition state
at the time of inspection and condition state movement over time can be captured
with the transition probabilities.
The most important property of the Markov chain is that the probability of move-
ment depends only on the current condition regardless of the history of movement,
Deterioration Modeling of Concrete Bridges and Potential … 403

called memoryless property. This means that (a) the currently known condition can
represent the accumulated deterioration up to the current time and (b) the future
condition does not depend on how long it stays in previous conditions. This property
is well suited for long-life assets such as concrete. The opposite is usage-based assets
such as light bulbs and machinery in which their future condition depends on how
long they have been used in the past.
The Markov model is mathematically expressed as a matrix M of transition
probabilities:
⎡ ⎤
P11 P12 P13 P14
⎢ P21 P22 P23 P24 ⎥
M =⎢
⎣ P31
⎥ (2)
P32 P33 P34 ⎦
P41 P42 P43 P44

where Pij is the probability of moving from condition state i to condition state j over
a unit time step. In this study, Pij was assumed to be 0 if i > j, meaning improved
condition was not modeled due to the lack of maintenance data.
The memoryless property and the theory of total probability can be used to predict
the probability vector [P1 P2 P3 P4] at any future time T, given the known current
condition with certainty or probabilities [C1 C2 C3 C4] as shown in the Chapman–
Kolmogorov equation [15]:

[P1 P2 P3 P4] = [C1 C2 C3 C4] ∗ M T (3)

Equation (3) is a matrix multiplication of the current condition with the transition
matrix powered to future time T. To utilize Eq. (3), the transition matrix M of Eq. (2)
needs to be estimated. Among several calibration/estimation techniques, we used the
proven Bayesian Markov chain Monte Carlo simulation technique [16] to estimate the
transition matrix M of the Markov model. In brief, the Bayesian technique transforms
the unknown elements Pij of the transition matrix M into a multivariate distribution by
using observed data, Eq. (3) and the Bayesian theory. The Markov chain Monte Carlo
simulation is then used to generate sampling data of the elements, which become the
estimator of the unknown elements (see [16]).
The predictive performance of the Markov model is commonly validated using
the Chi-square test on a separate dataset that is not used in the calibration of the
Markov model [15, 16]. The validation dataset is often randomly selected from 15
to 20% of the entire dataset.

3 Results

The deterioration models were applied to the case study and the results are
demonstrated for a concrete open girder.
404 H. Tran and S. Setunge

3.1 Model Fitness and Effect of Long-Term Data Collection

The effect of long-term and short-term data collection was investigated using the
linear and Markov models. Table 1 shows that the Markov model passed the fitness
test for both datasets. The negative values for R2 of the linear model on both datasets
indicate the poor fitness of the linear model and it should not be used. The poor fitness
of the linear model can be seen in Fig. 1 where the observed data have high uncertainty
as shown by the unpatterned scattering. Instead of trying to fit an impossible line
through the observed percentage in condition 1 with age, the Markov model takes
a different approach by capturing the percentage changes between condition states.
This shows the good generality of the Markov model in such cases.
Despite the unacceptable fitness of the linear models, Fig. 1 can still be used to
illustrate the significant effect of long-term data collection. This effect showed the
predicted deterioration rate is sensitive to the range of calibration data. The predicted
deterioration rate for long-term data collection (Fig. 1a) has a mild slope as compared
with the steep slope of the predicted deterioration rate for short-term data collection
(Fig. 1b).
The calibrated transition matrix for an old bridge is shown as a demonstration in
Eqs. (4) and (5) for long-term and short-term data collection. Figure 2a, b shows the
effect of long-term data collection on the Markov model. It appears that removal of
the first inspection accelerated the deterioration. For example, at the age of 100 years,
the percentages in condition 1 were 45 and 18% between the long-term and short-
term data. Similarly, the percentage of 20 and 30% in condition 4 can be observed
between the long-term and short-term data, implying a 15% difference.

Table 1 Results of fitness

Data Markov model Linear model
test for Markov and linear
models for percentage Chi-square test R2 value
prediction on condition 1 Long-term 55.8 < 85.9 −0.544
Short-term 7.8 < 67.5 −0.247

Fig. 1 Linear model fitted with long-term (a) and short-term (b) data for percentage in condition 1
Deterioration Modeling of Concrete Bridges and Potential … 405

Fig. 2 Markov model using long-term data (a) and short-term data (b)

⎡ ⎤
0.9934 0.0052 0.0008 0.0006
⎢ 0 0.9984 0.0009 0.0007 ⎥
M(all) = ⎢
⎣ 0
⎥ (4)
0 0.9972 0.0028 ⎦
0 0 0 1
⎡ ⎤
0.9894 0.0082 0.0018 0.0006
⎢ 0 0.9889 0.0104 0.0007 ⎥
M(second) = ⎢
⎣ 0
⎥ (5)
0 0.9991 0.0009 ⎦
0 0 0 1

3.2 Potential Benefit of Nano-concrete

The deterioration rate of traditional concrete was estimated using visual inspection
data. There were 3000 bridge structures in the case study. If it is assumed that at
current year 2022, all bridge structures in failure condition 5 have already been
repaired or replaced, then the budget for proactive asset management over the next
10 years (selected as an example) can be estimated as: (a) unit cost of major repair
or replacement AUD 100,000 per bridge girder for its average length of 30 m. The
penalty cost is assumed being equal to unit cost of inspection; (b) Markov deteri-
oration model predicts the increase by 4.0% of 3000 bridge structures (i.e. ≈120
structures) that will have its girder in failure condition 5 and require major repair or
replacement. The replacement budget is therefore AUD 12 million (which is 120 ×
$100,000); and (c) if nano-concrete is used instead of traditional concrete, the rate
of deterioration could be 20% lower because of its better strength and durability,
resulting in 4%*0.8 = 3.2*3000 = 96 girders that require major repair or replace-
ment. The budget cost is AUD 9.6 million dollars and the saving is AUD 2.4 million
over 10 years.
406 H. Tran and S. Setunge

It should be noted that the cost figures are hypothetical values that were used in
this study only for demonstration of methodology.

4 Discussion

This study was limited to a case of concrete material with a demonstration of an

open girder 2C. However, the methodology can be extended to all bridge elements
with different materials. The effect of contributing factors such as traffic volume and
exposure condition will be investigated in further studies.
This study revealed the weakness of the linear model and thereby supports the
use of the stochastic Markov model in modelling deterioration of bridge elements.
Despite this advantage, the Markov model still shows deterioration in the first few
years, which is considered unrealistic as found in a previous study [17]. Sobanjo and
Thompson claimed that new structures are expected to stay in condition 1 for the first
10–30 years before transitioning into a deteriorated condition. If all structures follow
this pattern, then the Markov model should be applied to the data after transition start.
If some new structures start to move to condition 2 after a few years due to unknown
causes such as uncertainty of material batch, inconsistent construction practice or
damage events, then the Markov model can reflect that. Therefore, it is not the
shortfall of the Markov model but how valid data are compiled for the deterioration
models. The comparison between the Markov model and other stochastic models
such as the Gamma process model [11] will be conducted in further works.

5 Conclusions

Nanomaterials including nano-silica, nano-alumina and nano-titanium oxide are used

in current research into developing nano-concrete. The addition of nanomaterials to
concrete can be viewed as having a similar effect to that with micro-based mate-
rials such as metakolin and silica fume. The effect comes from pore refinement and
thereby increases the strength and durability of concrete. The long-term performance
of nano-concrete used in bridge structures is still unknown. We used deterioration
models to estimate the deterioration rate of traditional concrete and conducted a
scenario analysis to understand the potential benefit of the nano-concrete. This can
significantly improve the efficiency of MRR management of bridge networks. We
also compared the deterministic linear model against the stochastic Markov model.
The demonstration on a concrete open girder showed that the Markov model is more
suitable than the linear model. Further investigations will be conducted to study the
effect of contributing factors, and to compare with other stochastic models.
Deterioration Modeling of Concrete Bridges and Potential … 407

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Transfer and Substrate Effects on 2D
Materials for Their Sensing and Energy
Applications in Civil Engineering

Q. Zhang, C. Zheng, K. Sagoe-Crentsil, and W. Duan

Abstract The recent emergence of two-dimensional (2D) materials such as

graphene and transition metal dichalcogenides (TMDs) of the family (Mo, W)(S,
Se)2 has attracted interest from a broad range of engineering applications, including
advanced sensing and energy harvesting and conservation, because of their distinctive
properties. However, it is critical important to achieve intact delamination and transfer
of these atomically thin materials, as well as to understand the effects of the target
substrates on their optical and electronic properties. Therefore, we developed and
compared techniques for transferring as-grown WS2 crystals to arbitrary substrates.
Polystyrene-assisted wet transfer can realize improved preservation of monolayer
WS2 crystals than the commonly used poly(methyl methacrylate) (PMMA)-assisted
wet transfer method, due to minimal chemical etching involved in the 2D material
delamination process. The intercalation of alkali ions in the PMMA-based transfer
method induces chemical doping over the transferred 2D crystals, leading to the
formation of trions. Moreover, the edges of the crystals on hydrophilic substrates,
such as sapphire or SiO2 /Si, are subject to ambient water intercalation, which
locally affects the photoluminescence behavior of the monolayer WS2 by doping and
changing of the dielectric environment. This non-uniform optical behavior is absent
when the crystal is transferred onto a hydrophobic substrate through which ambient
water cannot penetrate. These results have important implications for the choice of
target substrate and transfer method adopted for 2D TMD-based applications such as
next-generation strain sensing, photodetectors, gas sensing, bio sensing, solar energy
harvesting and radiative cooling in which uniform behavior of the channel material
is required.

Keywords Two-dimensional materials · Transfer · Photoluminescence · Sensing ·

Dielectric screening

Q. Zhang (B) · K. Sagoe-Crentsil · W. Duan

Department of Civil Engineering, Monash University, Clayton, VIC, Australia
e-mail: zqhmonashcsu@gmail.com
C. Zheng
Department of Physics, School of Science, Key Laboratory for Quantum Materials of Zhejiang
Province, Westlake University, Hangzhou, Zhejiang Province, China

© The Author(s) 2023 409

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_42
410 Q. Zhang et al.

1 Introduction

Because of their distinctive optical, electrical and mechanical properties, two-

dimensional (2D) materials such as graphene and transition metal dichalcogenides
(TMDs) are promising candidates for the high performance electronics and opto-
electronics needed for next-generation sensing and energy applications [1–3]. For
example, monolayer WS2 in the TMDs family has been reported as a promising
strain sensing that are integrated into infrastructure such as building windows [5]. To
realize their ultimate utilization in diverse applications, a feasible non-defective and
quality-preserving transfer technique is essential to create functional devices or van
der Waals heterostructures [6, 7]. The polymer-assisted wet transfer technique using
poly(methyl methacrylate) (PMMA) is widely used because of its better applicability
and material conservation [4, 8, 9]. However, this method usually involves soaking
the sample in heated alkaline solution, resulting in chemical doping as well as a
certain level of damage to the transferred materials [4]. To overcome this drawback,
hydrophobic polystyrene (PS) has been adopted as the supporting polymer layer for
the Chemical Vapor Deposition (CVD)-grown WS2 single crystal transfer process,
which enables less exposure of the materials to the alkaline solution [10]. Neverthe-
less, there is still a lack of detailed characterization of the crystals before and after
transfer to fully understand the doping effect of the transfer process.
Furthermore, because of their atomic thickness and reduced dielectric screening,
the properties of 2D materials are greatly influenced by the underlying substrate [11].
For 2D TMDs, the reduced Coulomb screening gives rise to high exciton binding
energies and quasiparticles of trions, which are charged excitons. Trions have been
observed in doped monolayer TMDs [12]. The characteristics of excitons and trions,
which significantly affect electronic transport and optical transitions in the mono-
layer TMDs, are found to be strongly coupled with the dielectric environment [13],
as well as doping effects [14] induced by the underlying substrate. In this regard,
non-uniform photoluminescence (PL) emissions of CVD-grown TMDs crystals on
dielectric substrates are widely observed [4, 15–18]. However, the underlying physics
have not been clearly demonstrated and further investigation is required to understand
the mechanism of this non-uniform PL behavior.
In this study, the distinctive effects of the substrate and the influence of doping
from the transfer process on the optical properties of CVD-grown monolayer WS2
triangular crystals were demonstrated. Optical and morphological characterizations
indicated that the non-uniformity of the monolayer WS2 PL behavior is highly likely
related to the intercalation of a layer of ambient water molecules between the WS2
crystal and the underlying hydrophilic substrate, where the intercalated water changes
the dielectric environment and the doping state. Comparison was made between the
commonly used PMMA transfer and the modified PS transfer methods in transferring
single WS2 crystals from as-grown sapphire substrates to target SiO2 /Si substrates.
Detailed morphology mappings showed that PS transfer maintained better preser-
vation of the crystals in transfer using PMMA, due to the avoidance of soaking in
alkaline solution. Optical characterizations of crystals before and after PS transfer
Transfer and Substrate Effects on 2D Materials for Their Sensing … 411

indicated that the residual alkali metal ions facilitated the formation of trions in
monolayer WS2 under ambient conditions, leading to variations in PL behavior. In
contrast to the transferred WS2 crystal on hydrophilic SiO2 /Si substrates, where non-
uniformity of PL along the edges remained, the transferred WS2 on hydrophobic PS
presented uniform PL emission across the crystal, supporting the theory that water
intercalation is the source of the inhomogeneous PL behavior on hydrophilic dielec-
tric substrates. After removal of alkali dopants by annealing at 100 °C in an argon
gas environment, the trion peak diminished, with only the exciton peak contributing
to the PL emission, indicating that the trion peak formation closely correlated with
the electron doping caused by alkali metal ions.

2 Methods

2.1 CVD Growth of Monolayer WS2 Crystals

The WS2 crystals were grown by CVD on substrates of sapphire [Al2 O3 (0001)]
using WO3 and S powders as precursors. The detailed CVD setup for monolayer
single WS2 crystal growth can be found in Zhang et al. [4].

2.2 Polymer-Assisted Wet Transfer Using PMMA and PS

For PMMA wet transfer, the CVD-grown monolayer WS2 on the sapphire substrate
was spin-coated at 3000 rpm for 60 s by PMMA (A4) and soft baked at 80 °C for
3 min. Next, the spin-coated sample was soaked in 2 mol/L KOH solution and heated
to 100 °C on a hot plate for 1 h. Before soaking, the edges of the PMMA thin film on
the sample were scratched with a scalpel to facilitate penetration of alkali between
the polymer film and the substrate. After soaking, the PMMA film was separated
from the as-grown sapphire substrate at a deionized (DI) water surface with the help
of surface tension. The separated PMMA film floated on the surface of the DI water
with the WS2 crystals attached. The film was then fished out by the target substrate
from underneath (WS2 side). The PMMA film was removed by soaking in acetone
and isopropanol (IPA) in turn.
For the PS transfer, PS (Mw ~ 192,000) in toluene solution (50 mg/mL) was used
to spin-coat as-grown WS2 on a sapphire sample. The steps of the PS transfer process
were the same as those introduced in Xu et al. [10]. Similarly, the PS film was fished
out by SiO2 /Si from underneath (WS2 side) and the polymer was washed off by
acetone and IPA. A further cleaning step of soaking the sample in PG remover was
used for thorough removal of the PS. To transfer WS2 onto a hydrophobic PS surface,
a rigid substrate of SiO2 /Si was used to attach the detached PS film from the top,
leaving the WS2 sitting on the PS surface.
412 Q. Zhang et al.

2.3 Optical and AFM Characterizations

Raman and PL characterizations were performed with a confocal microscope system

(WITec alpha 300R) with ×50 objective lens and 532 nm laser excitation. A weak
laser power of 50 µm and a short integration time of 1 s was adopted. The PL intensity
mappings were obtained by totaling the PL intensity from 1.9 to 2.1 eV. The Raman
and PL spectra were collected using a 600 line mm−1 grating. The atomic force
microscopy (AFM) measurements were carried out on the Bruker Dimension Icon
in tapping mode.

3 Results and Discussion

Characterizations of a representative as-grown monolayer WS2 crystal by CVD on

an atomically flat sapphire substrate are presented in Fig. 1. The optical image in
Fig. 1a shows a typical triangular single WS2 crystal ~5 µm in size with uniform
contrast. The AFM morphological mappings of the crystal (Fig. 1b, c) revealed that
the grown WS2 followed the atomic steps of the surface of the bare sapphire with step
heights of 0.2 nm, indicating uniform height and atomic flatness. However, it was
observed that the heights of the edges were slightly raised by ~0.37 nm, as shown in
Fig. 1c, g. This was inferred to be raised by a single layer of water molecules from the
ambient environment that intercalated between the sapphire and the monolayer WS2
crystal. This intercalated water layer can induce localized strong dielectric screening
to the monolayer WS2 because water has a much higher dielectric constant value
than sapphire [19].
This finding was further supported by the optical features reflected by Raman
intensity, PL intensity and PL peak position mappings for this WS2 crystal on

Fig. 1 Characterizations of a CVD as-grown monolayer WS2 crystal. a Optical image; b AFM
height scan; c zoomed-in AFM mapping of the region shown in the white square in (b); d–f Raman
intensity, PL intensity and PL peak position mappings; g height profile along the yellow line in (c);
h Raman spectrum collected in (d); i PL spectra taken from points 1–3 indicated in (e, f). Scale
bars: 5 µm in (a, b, d–f); 1 µm in (c)
Transfer and Substrate Effects on 2D Materials for Their Sensing … 413

sapphire. A Raman spectrum collected from the crystal (Fig. 1h) shows a distinc-
tive in-plane E , Raman peak of monolayer WS2 at 350 cm−1 . Although uniform
Raman signals were detected across the crystal (Fig. 1d), distinctive non-uniform PL
behavior was observed (Fig. 1e, f) where the edges of the crystal exhibited stronger
PL emissions and redshifted peak positions. The regions of edges with varied PL
behavior were consistent with the regions where the height was raised by water
intercalation, indicating that the trapped water had a notable influence on the optical
properties of the monolayer crystal.
To further understand the inconsistent PL behavior, three PL spectra were collected
at points 1–3 from the edge to the center of the crystal as indicated in Fig. 1e, f and
shown in Fig. 1i. Information regarding the three spectra is summarized in Table 1.
Lorentzian fitting indicated that the PL spectra at points 2 and 3 featured a single peak
each at 1.998 eV, coinciding well with the exciton peak of the CVD WS2 monolayers
on sapphire [20]. For the PL spectrum collected at point 1 located at the edge of the
crystal, however, the peak position was redshifted. Lorentzian fitting revealed two
distinctive peaks of an exciton (X 0 , magenta curve) at 1.999 eV and a trion (X − ,
cyan curve) at 1.975 eV, with an energy difference of 24 meV. This value was in
good agreement with the negative trion binding energy of the monolayer WS2 [21].
The exciton to trion intensity ratio was 1.4, agreeing well with the value for the
monolayer WS2 with 10 V positive gate voltage [21], indicating n-type doping in the
water-intercalated regions.
Besides the change in peak composition and the shift in peak position, the inte-
grated PL emission at the edge with water trapped underneath (point 1) also had
fivefold higher intensity than the inner region of the crystal (points 2 and 3). This
was likely due to the high dielectric constant (κ) of the water trapped underneath,
which led to prolonged lifetimes of the quasiparticles and increased recombination
efficiency of the excitons and trions [22, 23].
Next, the influence of the wet transfer method on the properties of the transfer
monolayer WS2 was investigated. First, the morphological changes of the crystal
transferred by processes using different polymers were studied. The transfer steps
with the two different polymers (i.e., PMMA and PS) are schematically illustrated in
Fig. 2a. As can be seen, compared with PMMA transfer, the PS film could be directly

Table 1 Spectral information for WS2 crystal on sapphire (corresponding to Fig. 1i)
Trion Exciton Trion Exciton to
Peak Peak Peak Peak Peak Peak binding trion intensity
position intensity width position intensity width energy ratio
(eV) (a.u.) (meV) (eV) (a.u.) (meV) (meV)

Point 1.975 208 58 1.999 322 44 24 1.4

1
Point 1.998 97 65
2
Point 1.998 100 64
3
414 Q. Zhang et al.

separated from the sapphire at the KOH solution surface with minimal exposure
to the alkaline environment. The effect of the alkali exposure is reflected in the
morphological characterization of the transferred crystals presented in Fig. 2b–e.
Obvious damage can be observed along the edges of the PMMA-transferred crystal,
with jagged shapes evident by the optical contrast (Fig. 2b) and AFM scan (Fig. 2d).
This damage to the edges was presumed to be caused by KOH etching from the long
soaking at elevated temperature [10]. Specifically, the raised edges of the crystals
by ambient water intercalation prior to the transfer process caused them to be more
easily subjected to etching by the alkali. In contrast, the morphology of the WS2
crystal transferred by PS indicated good preservation, including that at the edges
(Fig. 2c, e).
However, PS cannot easily be removed by acetone and IPA washing, so polymer
residue on the crystals can be observed. Therefore, PG remover, an NMP-based
solvent stripper, was further used to thoroughly remove the PS residue. As shown
in Fig. 3, the polymer residue was cleanly removed with edges well preserved.
These results suggested that PS transfer was superior to PMMA transfer, with better
preservation of the morphology of the transferred 2D crystals.
The change in the optical properties of the WS2 crystal after the PS transfer process
was investigated for the same crystal characterized in Fig. 1. To avoid confusion, the
crystal before PG remover washing is here referred to as the “as-transferred crystal”
and the crystal after PG remover washing is denoted as the “PG washed crystal”.
Figure 3a–d shows the optical images and PL mappings of the as-transferred crystal.
As can be seen clearly, non-uniform behavior of PL remains after the crystal was

Fig. 2 a Schematic of PMMA- and PS-based wet transfer methods. b, c Optical images of PMMA-
and PS-transferred WS2 crystals on SiO2 /Si; d AFM mapping of the PMMA-transferred crystal
shown in the square in (b); e AFM mapping of the PS-transferred crystal in (c). Scale bars: 10 min
(b, c); 2 min (d); 5 min (e)
Transfer and Substrate Effects on 2D Materials for Their Sensing … 415

Fig. 3 Optical characterization of the PS-transferred WS2 crystal onto SiO2 /Si. a Optical image,
b PL intensity mapping, c PL peak position mapping and d PL width mapping of the crystal as
transferred onto SiO2 /Si. e Optical image, f Raman intensity mapping, g PL intensity mapping and
h PL peak position mapping of the crystal transferred onto SiO2 /Si and soaked in PG remover. i PL
spectra taken at points 1–3 in (b–d). j PL spectra taken at points 1–3 in (g, h). Scale bars: 5 µm in
(a–h)

Table 2 Trion and exciton peak information in spectra for WS2 crystal transferred onto SiO2 /Si
by PS transfer (corresponding to Fig. 3i)
Trion Exciton Trion Exciton to
Peak Peak Peak Peak Peak Peak binding trion intensity
position intensity width position intensity width energy ratio
(eV) (a.u.) (meV) (eV) (a.u.) (meV) (meV)

Point 1.995 791 58 2.024 1143 38 29 1.4

1
Point 1.995 421 72 2.031 532 42 36 1.2
2
Point 1.995 352 75 2.033 382 42 38 1.1
3

transferred onto SiO2 /Si, with the edges exhibiting stronger PL intensity, redshifted
peak position and narrowed peak width. Because SiO2 has a hydrophilic surface and
a very small dielectric constant of 3.9 compared with water, it was deduced that this
PL non-uniformity was induced by water intercalation. It has been reported that the
trapped water can be removed by annealing at high source-drain bias or prolonged
baking in air [24, 25]. Figure 3i and Table 2 shows three PL spectra collected at
points 1–3 from the edge to the center. Unlike the spectra taken before the transfer,
all three PL spectra for the as-transferred crystal feature trion emissions comparable
to excitons. This strongly suggested that the WS2 crystal was subjected to doping
after the transfer process. The doping was presumed to have been induced by the
alkali metal ions of K+ from the transfer process acting as electron donors. The
notable PL peak position blueshift compared with before transfer (~25–35 meV) can
be explained by the release of strain induced by high temperature CVD growth after
the PS transfer process [20].
416 Q. Zhang et al.

After PG washing, the crystal was free of polymer residue, as indicated by the
optical image in Fig. 3e. However, the optical characterizations suggested severe
degradation of the optical properties of the crystal, reflected in the weak signals in
Raman and PL intensity mappings in Fig. 3f–h. The PL spectra fittings in Fig. 3j
show a weak single peak at the low energy of 1.950–1.975 eV for each spectrum
of points 1–3, consistent with a defect-bound exciton emission derived from defects
across the crystal [26]. Further study is required to address the negative impact of
PG remover on the transferred WS2 .
To further study the doping induced by alkaline solution, a CVD as-grown mono-
layer WS2 crystal was transferred onto a hydrophobic PS surface, using the PS wet
transfer to eliminate the influence of intercalated water. As can be seen from the PL
intensity, peak position and peak width mappings in Fig. 4a–c, the as-transferred WS2
crystal on PS exhibited uniform contrast over the entire crystal, with no variation
along the edges. This observation strongly supported the theory that the non-uniform
PL behavior along the edges stems from ambient water intercalation between the WS2
crystal and the underlying hydrophilic substrate. The PL spectra collected at points
1–3 on the as-transferred crystal on PS are shown in Fig. 4g, with details presented
in Table 3. The features of the three spectra are nearly identical, implying consistent
optical behavior over the crystal without trapped water along the edges. The nega-
tive trion emission across the crystal with a binding energy of 28 meV was strongly
indicative of n-type doping of the alkali metal ions after the PS transfer process.
Ar annealing of the crystal at 100 °C for 1 h was then performed to confirm the
effect of alkali doping. The optical mappings of the crystal after Ar annealing are
presented in Fig. 4d–f, with the PL spectra at points 1–3 shown in Fig. 4h. At elevated
temperatures, impurities, including the alkali dopants, could be detached from the
crystal and exhausted by Ar gas flow. Thus, after Ar annealing, it can be seen in
Fig. 4h that only a single peak was required to fit the PL spectrum for each point.
The peak positions are at ~2.013 eV, corresponding to exciton emission energy. The
suppression of trion peaks and the decrease of the PL intensity indicated that electron
doping of the crystal was significantly diminished by removal of residue alkali during
the annealing process. The exciton peak position was blueshifted compared with

Table 3 Trion and exciton peak information in spectra of WS2 on PS by PS transfer (corresponding
to Fig. 4g)
Trion Exciton Trion Exciton to
peak Peak Peak Peak Peak Peak binding trion intensity
position intensity width position intensity width energy ratio
(eV) (a.u.) (meV) (eV) (a.u.) (meV) (meV)

Point 1.976 127 55 2.003 223 37 27 1.8

1
Point 1.975 119 58 2.003 249 37 28 2.1
2
Point 1.974 137 55 2.002 214 39 28 1.6
3
Transfer and Substrate Effects on 2D Materials for Their Sensing … 417

Fig. 4 Optical characterization of a WS2 crystal transferred onto hydrophobic PS using the PS
wet transfer method. a PL intensity, b PL peak position and c PL width mappings of the crystal as
transferred onto PS. d PL intensity mapping, e PL peak position and f PL width mappings of the
crystal transferred onto PS and annealed by Ar at 100 °C. g PL spectra taken at points 1–3 in (a–c).
j PL spectra taken at points 1–3 in (d–f). Scale bars: 2 min (a–f)

that of the crystal as transferred onto PS (2.003 eV) as a result of reduced electron
doping, which also led to decreased intensity and narrowed width of the PL peaks
after annealing.

4 Conclusions

We investigated the effects of polymer-assisted wet transfer process and type of

substrate on the optical properties of CVD-grown monolayer WS2 crystals. We found
that transfer using PS instead of the traditionally adopted PMMA led to better preser-
vation of the morphology of the transferred WS2 crystals due to minimized exposure
to alkali solution. The electron doping induced by the alkali metal ions from the
transfer processes caused trion formation as well as intensified PL emission across
the transferred crystal. Annealing by Ar at 100 °C can efficiently remove alkali
residue on the transferred crystals. We also demonstrated that ambient water inter-
calation between the edges of the monolayer crystal and the underlying hydrophilic
substrates induces variations in PL due to doping and dielectric screening effects.
The non-uniform PL behavior of the WS2 crystals disappeared when the crystals
were transferred onto a hydrophobic PS substrate. These results contribute to the
understanding of trion formation in atomically thin TMDs via chemical doping and
non-uniform PL behavior stemming from the dielectric screening effect of water
intercalation when the material is deposited on hydrophilic substrates. This work has
implications for the provision of consistent crystal behavior of 2D TMDs for their
applications in sensing and energy harvesting and conservation devices.
418 Q. Zhang et al.

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Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
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the copyright holder.
Experimental and Numerical Studies
on the In-Plane Shear Behavior
of PVC-Encased Concrete Walls

Kamyar Kildashti and Bijan Samali

Abstract The effective application of lightweight stay-in-place concrete forms for

casting shear walls subjected to wind and seismic loading is of particular concern to
practitioners. Insufficient technical data available for new kinds of wall systems, such
as Polyvinyl Chloride (PVC) form walls, hinder their implementation in construc-
tion practice. To that end, an effective experimental and numerical campaign was
launched at Western Sydney University to investigate the structural performance
of PVC form walls when subjected to in-plane shear loading. A set of push-out
specimens was designated to conduct monotonic in-plane shear tests until failure.
All failure phenomena, capping strengths, and ductility capacities were monitored.
Test results indicated that the embedded PVC latticed webs could efficiently protect
the concrete web from sudden crushing and improve ductility capacity and failure
pattern of the specimens. Nonlinear finite element analysis on test specimens was
also conducted and good correlation with experiment results was achieved.

Keywords Stay-in-place forms · Squat concrete walls · Finite element analysis

1 Introduction

The common construction technique of stay-in-place (SIP) concrete forms delivers

low installation costs and high speed of construction. A novel SIP form, cast with
high-strength fabrics, was developed by Li and Yin [1] and its composite action
with inner concrete was studied. Both finite element (FE) analyses and experimental
lab works were conducted by various researchers to explore the performance of SIP
encasements under mechanical and thermal loading conditions [2–4]. The application
of cold-formed sections in the off-site manufacturing of SIP forms was studied by Wu
et al. [5]. The so-called Insulated Concrete Forms (ICF) were extensively employed

K. Kildashti (B) · B. Samali

Centre for Infrastructure Engineering, School of Engineering, Design, and Built Environment,
Western Sydney University, Penrith, NSW, Australia
e-mail: k.kildashti@westernsydney.edu.au

© The Author(s) 2023 421

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_43
422 K. Kildashti and B. Samali

Fig. 1 Geometry of PVC

form

to cast either squat or slender walls and their structural composite action with the
concrete core were mostly studied under monotonic and cyclic loading regimes [6–
10]. A novel stackable Polyvinyl Chloride (PVC) SIP form was introduced to cast
concrete walls even with complex geometries [11–13], as shown in Fig. 1. As the PVC
form engages with the inner concrete, the new design represents a departure from
traditional solid concrete core in responding to in-plane shear loads. Therefore, the
in-plane shear behavior of PVC form walls needs to be explored through experimental
and numerical studies. In this paper, we present in-plane shear behavior of PVC form
walls using experimental tests on push-out specimens as well as FE analysis.

2 Methods and Results

2.1 Test Outline

A total of three PVC form wall push-out specimens were tested to failure to acquire a
better understanding of the in-plane shear capacity and ductility with the influence of
PVC encasement. Further, one push-out wall specimen was cast using the traditional
form to be enabled for direct comparison with PVC form walls. The geometry of
specimens along with the test parameters are, respectively, shown in Fig. 2 and
in Table 1. These tests were performed to investigate the effects of vertical and
horizontal reinforcement ratio, wall thickness, and the PVC encasement. Constants
were concrete compressive strength and steel yield stress. Concrete compressive
strength and steel tensile yield stress were, respectively, 40 and 550 MPa. A 10,000 kN
servo hydraulic machine was employed to apply vertical force parallel to the shear
surface. Figure 3 Shows photographs of PVC and traditional form walls secured
in the testing rig. Linear Potentiometers (LPs) were employed to measure vertical
Experimental and Numerical Studies on the In-Plane Shear Behavior … 423

Fig. 2 Geometry of specimens

Table 1 Push-out test specimens

Specimen name Length Height Wall PVC Horizontal Vertical
L w (mm) Hw thickness encasement reinforcement reinforcement
(mm) tw
(mm)
F16-RW156S 1165 1000 150 Yes N12-350 mm N12-250 mm
F17-RW200D 195 Yes N12-233 mm 2N12-250 mm
F18-RW200D 195 Yes N16-233 mm 2N12-250 mm
F19-STW200D 195 No N16-233 mm 2N16-250 mm

displacement that occurred along the shear plane. The location of LPs at the top and
bottom face of specimens is shown in Fig. 4.

2.2 Test Results

Qualitative observations of the response of PVC form specimens were recorded

during the tests. Most of the observations were visual including crack initiation,
growth, and propagation, as well as excessive deformation through PVC encase-
ments. Figure 5 shows the failure shapes of PVC form and traditional form concrete
specimens. In Fig. 5a, it is identified while the slip occurred due to concrete failure
along shear plane, the excessive deflection in PVC encasement took place due to
outward dilatation of concrete. As seen in Fig. 5b, the concrete crushing near the
bottom corner close to the shear plane is mostly attributed to substantial slip along
the shear plane. In Fig. 5c, inclined shear cracking near the shear plane as well
424 K. Kildashti and B. Samali

(a) PVC form specimen (b) Traditional form

specimen

Fig. 3 Specimens in testing rig

(a) Bottom face (b) Top Face

Fig. 4 Instrumentation of specimens

as crushing and spalling of the concrete is associated with failure at shear plane.
More specifically, this failure occurred abruptly accompanied with explosive sound.
Figure 6 depicts force-slip relationships pertinent to different specimens. As seen,
specimens cast with PVC forms demonstrate substantially high ductility compared
to one cast with the traditional form. The major reason is the confining action of
PVC encasement to protect concrete core against dominating premature spalling
off. Further, the higher in-plane shear strength of the traditional form specimen is
attributed to the thickness and amount of vertical and horizontal reinforcement.
Experimental and Numerical Studies on the In-Plane Shear Behavior … 425

(a)

(b) (c)

Fig. 5 Phenomena of failure a excessive deformation of PVC encasement in a PVC form specimen
b concrete crushing at the bottom face in a PVC form specimen c diagonal concrete crushing in
traditional form specimen

Fig. 6 Force-slip behavior 3000

of different specimens
2500
Lateral load (kN)

2000

1500

1000 F16-RW156S
F17-RW200D
F18-RW200D
F19-STW200D
500

0
0 2 4 6 8 10 12 14 16 18 20 22 24 26 28 30 32 34
Lateral displacement (mm)
426 K. Kildashti and B. Samali

2.3 FE Simulation

In this study, the Abaqus program [14] was employed for FE analyses. The proper
selection of constitutive material model played an important role in verifying FE
models against experiments. Three built-in material models including Concrete
Damaged Plasticity (CDP), Crushable Foam (CF) [20], and J 2 Plasticity were used,
respectively, to model concrete, PVC encasement, and horizontal/vertical reinforce-
ments. The ascending and descending branches of uniaxial stress–strain curves in
compression pertinent to CDP model were assumed as follows [15]:

κ.η − η2
σc =(1 − H [|εc | − |εc1 |]) f cm
1 + (κ − 2).η
 −1
2 + γc f cm εc1 γc εc2
+ H [|εc | − |εc1 |] + γc εc + (1)
2 f cm 2εc1

where
σc = compression stress in (MPa)
εc = compression strain

η = εc /εc1

κ = E ci /E c1

γc = constant crushing energy [16]

π 2 f cm εc1
γc =   
2 gcl∗ − 1
2
f cm εc1 (1 − βc ) + βc f cm
Ec

gcl∗ = volume specific localised crushing energy

εc1 = strain at maximum compressive stress
1/3
E ci = tangent modulus = 104 f cm in (MPa)
f cm = characteristic compressive strength
E c1 = secant modulus from the origin to the peak compressive stress
H = Heaviside function
The concrete stress–strain relation in uniaxial tension consisted of a linear phase
up to tensile strength, followed by a nonlinear strain softening phase that depended
on the specimen geometry. The stress-crack opening relation can be expressed as
[17, 18]
Experimental and Numerical Studies on the In-Plane Shear Behavior … 427
 
c1 w 3
c2 w w 
σct (w) = 1+ exp − − 1 + c13 exp(−c2 ) f ctm
wc wc wc
c1 = 3, c2 = 6.93 (2)

where
σct = tensile stress in (MPa)
w = lt εctck = lt (εct − σct /E ci ) = crack opening in (mm)
wc = 5G F / f ctm = crack opening when σct = 0 in (mm)
G F = 0.073 f cm
0.18
= fracture energy in (N/mm)
2/3
f ctm = 0.3 f cm = tensile strength in (MPa)
εct = tensile strain
lt = characteristic length in FE modelling
εctck = cracking strain
List of parameters used for CDP model calibration was presented in [19]. To
capture nonlinear behavior of the PVC encasement, parameters of the CF model were
calibrated against test results as reported in [11]. To establish yield surface for J 2 Plas-
ticity constitutive law, steel Young’s modulus, Poisson’s ratio, yield stress, ultimate
stress, and ultimate strain was, respectively, assumed as 200 GPa, 0.3, 550 MPa,
650 MPa, and 0.05. Figure 7 shows the comparison between failure phenomena
obtained from the experiment and FE simulations. The equivalent plastic strain was
chosen from FE analysis to represent localization of damage. As seen, the FE simu-
lation reasonably replicates the failure modes triggered and evolved across the PVC
encasement and concrete core. As shown in Fig. 8, there is a good correlation between
experimental results and those obtained from FE analysis.

3 Conclusion

Push-out experimental lab tests were conducted to explore in-plane shear capacity
of PVC form concrete walls under monotonic loading. Further, a numerical FE
model was established to predict the shear behavior of the walls. Experimental
observations revealed gentler failure phenomena for PVC form walls compared to
the one cast using the traditional form. It was also concluded that PVC encase-
ment provided confinement pressure against concrete outward dilation and therefore
enhanced ductility capacity was observed. Further, the established FE model showed
correlative results with those obtained from experiments.
428 K. Kildashti and B. Samali

(a) (b)

(c) (d)

Fig. 7 a A photo of excessive deformation on PVC encasement b localized plastic deformation on

PVC encasement obtained from the FE c a phot of concrete crushing at the bottom end d damage
evolution at the bottom end obtained from FE

Fig. 8 Comparison between 2000

FE and test results 1800
Shear load (kN)

1600
1400
1200
1000
800
F18-RW200D-Test
600
F18-RW200D-FE
400
200
0
0 10 20 30
Slip (mm)
Experimental and Numerical Studies on the In-Plane Shear Behavior … 429

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Recycled Glass-Based Capping Layer
for Foundations in Expansive Soils

H. Karami, D. Robert, S. Costa, J. Li, S. Setunge, and S. Venkatesan

Abstract Construction on weak expansive soils is challenging due to their low

bearing capacity and high-volume susceptibility under moisture fluctuation. The
uplift pressure from expansive clay can induce significant swelling pressure on foun-
dations, but on the other hand, shrinkage of clay can result in substantial foundation
settlement during dry seasons. This differential movement of the foundation can
distress the superstructure of a building, resulting in serviceable and ultimate limit
state failures. The current approach to dealing with foundations in expansive clay
soils is to construct a rigid slab that can withstand the anticipated movement or a pier-
type approach using engineering design principles or a normal-type slab based on the
fill being placed under controlled conditions. We introduce a capping layer under the
foundation to control moisture fluctuation and increase the bearing capacity of the
foundation using a recycled glass-based stabilization approach. A prototype foun-
dation was constructed in the laboratory using an optimum stabilization mix design
that was derived using glass and other sustainable additives. Slab movements and
soil conditions were monitored over a 6 month period under simulated dry/wetting
moisture fluctuations and operational loads. Results revealed that the performance
of the foundation under controlled conditions was outperformed by the capping
layered foundation during service loadings including seasonal moisture fluctuation.
The outcomes from this research will have a significant impact on improving foun-
dation performance in expansive soils, as well as proposing a sustainable foundation
construction process using recycled glass waste.

Keywords Capping layer · Expansive soils · Foundations · Glass waste ·

Prototypes · Recycling

H. Karami · D. Robert (B) · J. Li · S. Setunge · S. Venkatesan

Civil Engineering Department, School of Engineering, RMIT University, Melbourne,
VIC, Australia
e-mail: dilan.robert@rmit.edu.au
S. Costa
School of Engineering, Faculty of Science, Engineering and Built Environment, Deakin
University, Waurn Ponds, VIC, Australia

© The Author(s) 2023 431

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_44
432 H. Karami et al.

1 Introduction

Expansive soils are a significant hazard for lightweight building foundations and
highway pavements, because they experience changes in volume due to seasonal
moisture fluctuation, swelling during wet seasons and shrinking during dry periods. If
a pavement rests on expansive soil, longitudinal surface cracks may occur because of
the seasonal volume change of the subgrade expansive soil [1]. Infrastructure damage
due to expansive soils is commonly reported in many countries, such as Australia,
Canada, England, China, India, and the USAs. Consequently, the need for change in
conventional foundation construction systems is becoming a very important require-
ment to construct more sustainable houses and buildings with low maintenance costs
over their lifetime.
Soil stabilization is an effective way to enhance the durability, mechanical charac-
teristics and to reduce or eliminate the amount of volume change in expansive soils.
Using lime and cement for soil stabilization are less cost effective and not environ-
mentally friendly strategies due to the use of energy, resources and carbon footprint
produced during the manufacturing process. Moreover, investigations revealed that
cyclic wetting and drying cause arresting volume change behavior to be lost after
the first wet–dry cycle, and consequently swelling potential increases after each
cycle due to the formation of expansive material such as ettringite in calcium-based
stabilized soils [2].
Much attention has been given to reducing the utilization of natural resources in
cement and other traditional construction materials. For example, the construction
sector generated roughly 92.5 million tonnes of asphalt concrete and 4.6 million tons
of cement in 2013 and 2015, respectively [3, 4], necessitating alternative construction
materials to minimize environmental impact and to preserve natural resources. Hence,
the main goal of research by both scientists and engineers has been reducing the
demand for natural resources as well as to minimizing the disposal of wastes such
as glass [5–7]. Significant research has been performed into the use of waste glass
in construction. For example, cullet has been tested as aggregate in the construction
sector ranging from concrete and cementitious materials to roadway and asphalt
construction. In addition, waste glass can be utilized in the manufacture of ceramic-
based products [8].
The mechanical behavior of clay soils can be significantly improved by adding
glass powder to the raw soil [9]. Adding fly ash and cement together with recycled
glass powder increases the shear strength and CBR of soil [10–12]. Individually, glass
powder can increase the strength of cement-stabilized expansive soil and decrease
the plasticity index of the soil mixture. Furthermore, the addition of glass powder
increases dry density, CBR and UCS with a reduction in optimum moisture content
[13]. However, the effect of crushed glass on soaked CBR can be less pronounced
compared with unsoaked conditions. In terms of other additives, Phanikumar showed
that the addition of fly ash to expansive soils can modify the pore orientation of the
soil while significantly improving its compaction behavior [14].
Recycled Glass-Based Capping Layer for Foundations in Expansive Soils 433

Previous studies identified the optimum mix design using waste glass and
secondary additives, proposing specifications for a novel capping layer that can be
applied to minimize the impact of soft subgrades on foundations [15]. A series of
mechanical tests (UCS, standard compaction test and direct shear test), hydraulic
conductivity test (permeability) and microscopy test (X-ray diffraction, scanning
electron and porosity test) were carried out to explore the optimum combination of
glass wastes and secondary additives for field trials in expansive soils.
The aim of current study is to evaluate the performance of a capping layer by
applying it in large-scale tests as verification of the proposed novel capping layered
foundation for buildings and roads. A laboratory experiment was carried out to
investigate the performance of a foundation slab placed on an expansive soil. The
proposed capping layer was placed between the foundation and weak subgrade clays
to evaluate the foundation’s performance under environmental and operational loads.
Performance of the foundation and soil conditions were carefully monitored across
a period of time under operational loads for verification of the proposed capping
layered foundation system.

2 Methods

2.1 Materials

2.1.1 Soil Characteristics

Soil was obtained from a land excavation site in Melbourne, Australia. The soil
was categorized as fine-grained (CL) [16] with a high degree of expansion. All the
physical properties of the soil are shown in Figs. 1, 2 and Table 1.

Fig. 1 Particle size

distribution of soil 90
Percentage Passing (%)

80
70
60
50
40
30
20
10
0
0.1 1 10 100
Particle Size (µm)
434 H. Karami et al.

Fig. 2 Soil standard 1.84

1.83
compaction test result

DRY DENSITY (gr/cm³)

1.82
1.81
1.80
1.79
1.78
1.77
1.76
1.75
1.74
1.73
1.72
1.71
1.70
1.69
14 15 16 17 18 19 20 21 22
MOISTURE CONTENT (%)

Table 1 Summary of soil characteristics

Property MDD g/cm3 OMC % LL % PL % PI
Value 1.832 17.10 36.9 20.3 16.6

2.1.2 Crushed Recycled Glass Powder

Crushed glass behaves like natural rock and is totally inert and non-biodegradable.
Glass powder (GP) was used in this study. Particle size distribution and samples are
shown in Figs. 3 and 4 respectively.

2.1.3 Fly Ash

Class F fly ash from a local supplier was the primary additive used in the capping
layer. For all types of fly ash, the particle size distribution is similar to that of silt

Fig. 3 Particle size

distribution of glass powder 90
Glass Aggregate
and glass aggregate [15] 80 Glass Powder
Percentage Passing (%)

70

60

50

40

30

20

10

0
1 10 100 1000 10000 100000
Particle Size (µm)
Recycled Glass-Based Capping Layer for Foundations in Expansive Soils 435

Fig. 4 Glass powder used in

the study [15]

(Fig. 5). Typical physical and mechanical properties for fly ash are shown in Table 2.
Other characteristics of the fly ash including chemical composition can be found in
Karami et al. [17].

Fig. 5 Fly ash particle size

distribution [17] 90
Percentage Passing (%)

80
70
60
50
40
30
20
10
0
1.00 10.00 100.00 1000.00
Particle Size (µm)

Table 2 Physical properties

Property Value
of fly ash [14]
Class F or low lime fly ash
Specific gravity 2.14
Liquid limit 43
Plastic limit Non-plastic
Optimum moisture (%) 34
Max. dry density (gr/cm3 ) 1.1
436 H. Karami et al.

Table 3 Physical and chemical properties of hydrated lime [14]

Property Value Property Value
Specific gravity (g/cm3 ) 2.05 Calcium hydroxide Ca(OH)2 (%) 82
Density (kg/m3 ) 510 Chloride (%) 0.01
Normal consistency (%) 43.50 Silicon dioxide (SiO2 ) (%) 2.5
Initial setting time (min) 165 Aluminum (Al2 O3 )/iron (FeO) (%) 3.5
Final setting time (h) 46.25 Water (H2 O) (%) 0.6
Fineness (% by weight on 300 µm 2.65 Arsenic (%) 0.0004
sieve)
Soundness (Le Chatelier’s expansion 1.8 Lead (%) 0.0001
mm) (mm)
Compressive strength (14 days) (N/ 1.45 Sulfate (SO3 ) (%) 0.9
mm2 )
Compressive strength (28 days) (N/ 2.18 Magnesium oxide (MgO) (%) 3.5
mm2 )

2.1.4 Lime

Hydrated lime, which is the most concentrated form of lime, was used as a secondary
additive. Lime decreases the liquid limit of soil mixtures, resulting a reduction in
plasticity index. Typical physical and engineering properties of hydrated lime are
similar to what was used in this study are shown in Table 3.

2.2 Experimental Set-Up

2.2.1 Stage 1: Control Test

Preparation of the soil required an extremely thorough and careful process to obtain
as homogeneous material as possible. In preparation for the test, the natural soil was
first sequentially sieved with a 25 mm sieve. A large-scale soil box (L: 1900 mm, W:
750 mm, H: 1000 mm) was built to enable the control of hydraulic and mechanical
processes of the soil by changing the water table in the soil. A drain pipe and a valve
system were installed at the bottom of the box (i.e., at a depth of 1000 mm) to allow
drainage and to control the water table inside the box. Sieved soil was mixed with
water in a concrete mixer until a homogenous mixture with optimum moisture content
was obtained. The soil was then placed in the test box in 4 layers and compacted in
100 mm rises (Table 4). Figure 6 shows the details of the test box setup. A geotextile
sheet was placed over the bottom surface of the box, a 100 mm thick drainage layer
made of gravel was placed on top and finally a filter paper before the soil. The top soil
surface at the top was left exposed to atmospheric conditions to facilitate evaporation
Recycled Glass-Based Capping Layer for Foundations in Expansive Soils 437

and infiltration. The experiment began late December 2022 during the summer, and
continued for 6 months.

Table 4 Soil layer characteristics and thickness for stages 1 and 2

Stage 1
Layer Depth Material Color Stage 2
(mm) Layer Depth (mm) Material Colour
Reinforced Reinforced
A 650–820 A 650–820
concrete slab concrete slab
B 500–650 Capping layer
B 100–500 Raw Soil
C 100–500 Raw soil
C 0–100 Gravel
0–100 Gravel

Fig. 6 Proposed sensor locations for prototype test: a top view, b stage 1, c stage 2
438 H. Karami et al.

Table 5 Instruments
No. Description
1 UBIBOT soil moisture probe
2 UBIBOT soil moisture probe
3 Humidity and temperature probe
4 Surface moisture meter
5 Middle moisture meter
6 Bottom moisture meter
7 UBIBOT soil moisture monitoring station
8 Edge movement gauge
9 Edge movement gauge
10 Central movement gauge

Table 6 Load calculations

Item Standard residential load Actual applied load to the
(kPa) soil (kN)
Dead 3.35 1.38
load
Live 2 0.82
load
Total 5.35 2.2

As shown in Fig. 6 and Table 5, six moisture probes and three digital movement
sensors were installed in different locations. Two temperature/humidity sensors were
also installed in the test box to record both the soil and room temperatures. Data
were recorded in the controller at a frequency of 1 h and transferred to a computer
daily. In addition, weekly manual readings of the movement were taken. A sprinkler
system was constructed above the sandbox to simulate rain events similar to field
conditions. Water spray was based on the water table depth at specific times and
amounts. Water percentage was obtained from the swelling test results for raw soil
according to Australian Standard (AS 1289.7.1.1-2003). The amount of water was
calculated based on the raw soil amount in the box (Table 7).
For Stage 1, a concrete slab scaled down to L: 750 mm, W: 550 mm, and H:
100 mm was placed on top of the compacted soil to apply the dead loads during
the test period. The size of the concrete block was scaled down from an actual
designed slab for a residential building following AS 2780. In this test, the dead load

Table 7 Water added to the

Water percentage (%) Time (weeks) Amount (L)
sample
2.20 W2 22.36
1.91 W4 19.41
3.15 W6 32.01
Recycled Glass-Based Capping Layer for Foundations in Expansive Soils 439

was added to the slab for increasing the thickness of the slab from 100 to 170 mm
for a residential-purpose building. The live load was applied to the soil by specific
weights at the end of wetting and drying periods. The standard load amounts and the
calculated applied load based on the slab area are shown in the Table 6.

2.2.2 Stage 2: Test with Capping Layer

In Stage 2, all steps in Stage 1 were repeated, but a capping layer was constructed
on top of the raw soil (Fig. 6c). The design mix of the capping layer was based on
previous studies [15]. Materials were mixed with soil in percentages of 25% WG
powder, 7.5% fly ash and 3% lime by weight compared with raw soil. To prepare the
stabilized soil, all ingredients were mixed properly in the dry state in a mechanical
mixer. Water was added to the mixture up to the optimum moisture content according
to the Standard Compaction test and well mixed. Similar to Stage 1, the soil (subgrade
soil) was uniformly compacted in the box. The size of the installed capping layer
was same as the raw soil in length and width to cover the whole soil surface area.
Thus, the size of the capping layer was L: 1900 mm, W: 750 mm, and H: 100 mm.

3 Results and Discussion

3.1 Slab Settlement

At the beginning of both stages of the prototype test, the slab started settling down
for the first few days after it was placed on the soil surface, due to the weight of both
the slab and the capping layer on the raw soil, which caused a secondary compaction
process during those first few days. The results show that settlement in Stage 2 was
greater than in Stage 1 at early stages. The maximum settlement in Stage 1 was
0.06 mm and for Stage 2, it was 0.10 mm after week 1 for the control and stabilized
capping layer tests respectively (Fig. 7). The larger settlement in Stage 2 resulted
from elastic compression of the capping layer, which consisted of coarser materials
than raw soil. However, this increment in initial settlement was negligible and did
not imply a reduction in bearing capacity.
It is important to note that the subsequent slab movement due to swelling was
reduced when the capping layer was introduced. The difference in maximum slab
movement between the two stages was 0.2 mm. More importantly, the rate of increase
in slab movement was less with the capping layer in place. In the control test, slab
movement reached 0.4 mm in less than 5 weeks, whereas it took nearly 9 weeks to
move by the same amount under stabilized conditions. Therefore, it can be concluded
that the capping layer was effective in mitigating the magnitude and the rate of the
vertical slab movement.
440 H. Karami et al.

a)
0.70 35.0%

0.60
30.0%
0.50
Slab Movement (mm)

Moisture content (%)

25.0%
0.40

0.30 20.0%

0.20 15.0%
0.10
10.0%
0.00 Slab Movement Average (Stabilized)
Slab Movement Average (Control)
Surface Moisture (Stabilized) 5.0%
-0.10
Surface Moisture (Control)
-0.20 0.0%
W1 W2 W3 W4 W5 W6 W7 W8 W9
Time (weeks)
b)
0.70 40.0%

0.60 35.0%

Moisture content (%)

0.50
Slab Movement (mm)

30.0%
0.40
25.0%
0.30
20.0%
0.20
15.0%
0.10
Slab Movement Average (Stabilized) 10.0%
0.00
Slab Movement Average (Control)
-0.10 Middle Moisture (Stabilized) 5.0%
Middle Moisture (Control)
-0.20 0.0%
W1 W2 W3 W4 W5 W6 W7 W8 W9
Time (weeks)
c)
0.70 30.0%

0.60
25.0%
Moisture content (%)

0.50
Slab Movement (mm)

0.40 20.0%

0.30
15.0%
0.20

0.10 10.0%

0.00 Slab Movement Average (Stabilized)

Slab Movement Average (Control) 5.0%
-0.10 Bottom Moisture (Stabilized)
Bottom Moisture (Control)
-0.20 0.0%
W1 W2 W3 W4 W5 W6 W7 W8 W9
Time (weeks)

Fig. 7 Slab movement and moisture profiles for stage 1 and stage 2 during the wetting period from
week 1–9. Horizontal profiles: a at soil surface level, b 200 mm below the soil surface level, and
c at the base of the soil layer for Stage 1 and 100 mm above the base for stage 2
Recycled Glass-Based Capping Layer for Foundations in Expansive Soils 441

3.2 Expansive Soil Movement (Swelling)

The soil started swelling and increasing in volume as expected for any expansive soil.
During the wetting period, the soil in the control test swelled by 0.6 mm on average
(maximum 0.72 mm for individual gauges). In Stage 2, with the capping layer in
place, the slab moved by only 0.397 mm (maximum 0.61 mm for individual gauges)
(Fig. 7a–c). As shown in Fig. 7, the swelling in Stage 2 of the test with the stabilized
capping layer was lower than with the control soil during the wetting period.
Because the capping layer was more permeable than the control soil, it assisted
water to drain away faster and remove it from the slab and the expansive soil under-
neath. Moreover, the load applied to the soil from the weight of the capping layer
helped to arrest or reduce the swelling of the expansive soil.

3.3 Moisture Distribution

Water was sprayed on top of the soil with the same pattern (time, amount, and size
of spray nozzle) for both stages of the test. As shown in Fig. 7a, the capping layer
tended to be mostly dryer than the raw soil in the control test, because of the higher
permeability of the capping layer (Fig. 8). In the control test, water pooled on the
surface or remained at a shallow depth (based on visual observations) after a wetting
event whereas more water infiltrated the soil through the capping layer in Stage 2.
This was evident from the moisture variation at 200 mm from the surface (Fig. 7b).
Moisture content at the raw soil-capping layer interface was higher than the moisture
content recorded at 200 mm within the raw soil in the control test (Fig. 9). This was
due to the rapid flow of water through the capping layer. However, the capping layer

0.00001

0.000001
Permeability (cm/s)

0.0000001

1E-08

1E-09
Glass Powder (GP)
Soil+7.5% FA+3% Lime+GP
1E-10
0 5 10 15 20 25 30
Glass %

Fig. 8 Permeability results for soil–glass powder and soil–fly ash–lime–glass powder mix [17]
442 H. Karami et al.

Concrete Slab
0.000

-0.100
Increasing Depth (m)

-0.200

-0.300
Control Test - W2
Control Test - W6
Control Test - W9
-0.400
Stabilised Capping Layer - W2
Stabilised Capping Layer - W6
Stabilised Capping Layer - W9
-0.500
10.0% 15.0% 20.0% 25.0% 30.0% 35.0%

Moisture content

Fig. 9 Variation of water content in control soil and stabilized capping layer in the wet season of
the test: vertical profile

was effective in delaying water reaching the raw soil layer. Without the capping layer,
the moisture content at the surface of raw soil was >30% in less than 3 weeks. When
the capping layer was in place, it took almost 5 weeks for the water content at the
top of the raw soil layer (interface between the raw soil and capping layer) to reach
>30%. Consequently, the raw soil layer remained drier for a longer period, which
will reduce the vertical movement of the slab.
The moisture variation shown in Fig. 7c corresponded to the condition at the base
of the control test and at 400 mm depth below the surface for Stage 2. Interestingly, the
difference in moisture content at these locations was narrow, indicating the raw soil
layer remained drier in Stage 2 throughout the wetting period. Thus, it was evident
that the capping layer kept the water away from the expansive soil underneath.

4 Conclusions

This research investigated a novel sustainable foundation system that could mitigate
the adverse influence of expansive soil conditions for construction. The proposed
approach is based on a capping layer constructed from recycled glass waste and other
sustainable additives and placed directly above the existing expansive soil. A proto-
type foundation was constructed in the laboratory using an optimum stabilization mix
design derived from a detailed investigation of mechanical and hydraulic character-
istics. The slab movement and soil conditions were monitored over 6 months under
Recycled Glass-Based Capping Layer for Foundations in Expansive Soils 443

simulated dry/wet moisture fluctuations and operational loads. Results showed that
the foundation’s performance on the novel capping layer was significantly produc-
tive compared with control conditions. The higher efficiency of the new capping
layer system was evidenced by 35% reduction in slab displacement during wet
season simulation, which is a significant improvement in foundation serviceability
on expansive soils. The outcome from this research will have a significant impact
on minimizing infrastructure maintenance costs, which are a heavy burden for asset
managers in any country. The proposed capping layer, which incorporates waste
materials (32.5% of total), is a waste management strategy that minimizes the service-
ability concerns of civil infrastructure while introducing a value-added benefit for
waste materials. It is to be noted that the results provided in this document are based
on the referred tests/materials conditions and should not be used in field applications
without appropriate verification.

Acknowledgements This research was funded by the RMIT postgraduate scholarship

and the Australian Research Council Industrial Transformation Research Hub (grant no.
IH150100006, IH200100010).

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Submicroscopic Evaluation Studies
to Minimize Delayed Ettringite
Formation in Concrete for a Sustainable
Industry and Circular Economy

Yogesh Kumar Ramu, Paul Stephen Thomas, Kirk Vessalas,

and Vute Sirivivatnanon

Abstract The high cost of maintenance, repair and retrofitting of concrete infrastruc-
ture to keep these structures durable and serviceable is not sustainable, so the design
process needs to consider all aspects of deterioration mechanism/s that can potentially
occur in a concrete structure. The ideal solution should contribute to sustainability
by enhancing the durability of concrete elements and supporting a circular economy.
We studied delayed ettringite formation (DEF), a potential deterioration mecha-
nism, including mitigation measures, in various heat-cured cementitious systems.
The results showed that continuously connected pore/crack paths at the submicro-
scopic level favor the transportation of DEF-causing ions in heat-cured systems. DEF
increases the chance of developing cracks, which is a durability concern. To mitigate
DEF, fly ash produced from an Australian bituminous coal-burning power station
was incorporated in the binder to support the circular economy concept. Changes
in heat-cured cementitious systems were evaluated using expansion, electrical resis-
tivity, dynamic modulus, and microstructural studies. The pozzolanicity of fly ash
was found to greatly enhance the formation of denser calcium-silica-hydrate, which
in turn restricted the transportation of DEF-causing ions at the submicron level,
leading to less DEF occurrence and enhancement of the durability and sustainability
of concrete in field structures.

Keywords Circular economy · Delayed ettringite formation · Fly ash · Precast

concrete · Sulfate attack

Y. K. Ramu (B) · P. S. Thomas · K. Vessalas · V. Sirivivatnanon

School of Civil and Environmental Engineering, University of Technology Sydney, Ultimo, NSW,
Australia
e-mail: yogeshaccet@gmail.com; yogesh.ramu@uts.edu.au

© The Author(s) 2023 445

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_45
446 Y. K. Ramu et al.

1 Introduction

Modern-day concrete infrastructure requires routine maintenance, repair and/or

retrofit to ensure structures achieve their intended service life without compromising
durability and sustainability. Unfortunately, the modernization of the construction
industry solved a problem but created another. The precast concrete industry produces
offsite concrete members to enable faster construction throughput rates for infras-
tructure megaprojects, but deleterious/expansive delayed ettringite formation (DEF),
which has been linked to crack formation, is a significant concern for this industry [1–
5] Elevated temperature curing (>70 °C) in precast concrete production is believed
to be primarily responsible for inhibiting or decomposing early ettringite and, at a
later stage, recrystallizing ettringite results in expansion and cracking [6]. Cracking
in concrete causes permeability, which increases the risk of infiltration of external
harmful agents (e.g., sulfates and chlorides) that negatively impact concrete dura-
bility. However, despite the potential risk, Australia lacks research in DEF, creating
uncertainty within the industry. In this study we addressed the issue by evaluating
Australian cements and their susceptibility to DEF under heat-cured conditions. The
influence of sulfate and alkali spiking of these cements on DEF was also systemat-
ically studied and as mitigation of expansive/deleterious DEF, the incorporation of
fly ash (FA) to enhance the durability, and thereby the sustainability of concrete, was
also studied.

2 Materials and Methods

Two Australian general-purpose cements conforming to AS 3972:2010 [7] and

locally sourced Gladstone FA, complying with AS 3972:2010 [8], were used. The
composition of the binders was determined by X-ray fluorescence analysis (Table
1). The specific surface area of Cement 1 (C1) and Cement 2 (C2) was 395 m2 /kg
and 420 m2 /kg, respectively. The Na2 Oeq content of both cements (C1: 0.5%; C2:
0.35%) complied with the Australian Technical Infrastructure Committee’s specifi-
cation ATIC SP-43 [9]. C1 contained 7.5% limestone as mineral addition and C2
was free from mineral addition. Modified Bogue’s phase composition analysis indi-
cated that C2 had a higher tri-calcium aluminate, C3 A (11.15%) content than C1
(7.48%). Sand sourced from Rockhampton, Queensland, was used to prepare the
mortar prisms. Distilled water was used to prepare all mortar and paste specimens
used in the study. Calcium sulfate and sodium hydroxide were used to chemically
modify the cements.
Four different cement systems were used for mortar prisms prepared with
cement:sand:water mass ratio of 1:3:0.45. Mixing of mortars was carried out
according to AS2350.12-2006 [10]. Mortars were cast in steel molds (complying with
AS 2350.11-2006 [11]) of size 40 × 40 × 160 mm, incorporating steel expansion
studs. Three prisms per mix were prepared for the expansion studies. Cement paste
Submicroscopic Evaluation Studies to Minimize Delayed Ettringite … 447

Table 1 Chemical composition of binders by oxide analysis

Oxide % CaO SiO2 Al2 O3 Fe2 O3 SO3 MgO K2 O Na2 O TiO2 LOI Total
Cement 1 64.2 19.7 4.8 3.1 2.4 0.9 0.4 0.2 0.2 4.1 99.9
(C1)
Cement 2 65.7 22.8 4.4 0.2 3.0 0.7 0.3 0.1 0.2 2.2 99.6
(C2)
Fly ash 3.9 55.5 24.4 6.9 0.2 1.2 0.9 0.7 1.0 2.5 97.2

specimens were also prepared for characterization studies. To investigate the impact
of different cement systems and heat curing on DEF, three cementitious systems
and heat-curing regimes were designed: (1) cements without any additional sulfates
(C1, C2); (2) cement systems with 4% sulfate and 1% alkali added (C1SN, C2-N);
and (3) cement systems with 25% FA, 4% sulfate and 1% alkali added (C1SNF,
C2SNF). These cement systems were used to make mortar specimens that were
subjected to two different regimens (Fig. 1): ambient temperature (23 °C) cured in a
humidity chamber (relative humidity (RH) 90%) for 24 h (C1-A, C2-A); heat-curing
in a programmable laboratory oven. For the heat-curing process, the specimens were
pre-cured at 30 °C for 4 h, followed by a temperature ramp to 90 °C, which was
maintained for 12 h before natural cooling to 30 °C (≈6 h) (C1-H, C2-H, C1SN-H,
C2SN-H, C1SNF-H, C2SNF-H). In both curing regimes, specimens were demolded
24 h after casting and transferred to a lime-saturated water bath for further curing.
Changes in the length of the specimens were monitored in accordance with AS
2350.13-2006 [12], using a digital length comparator with an accuracy of 0.001 mm.
The frequency of length change monitoring was every 7 days for the first 90 days,
then every 30 days up to 600 days and after that, every 90 days.
For the characterization studies, the hydration of the cement pastes was arrested
with solvent replacement methods at the respective ages. Hydration-arrested samples
were kept in a desiccator with silica gel (as the drying agent) under laboratory-
controlled conditions (23 ± 2 °C). The stored samples were tested for phase changes
using Thermo gravimetric analysis. Thermal analysis was conducted with a Netzsch
Jupiter F5 STA instrument using helium flowing at 40 mL/min on 20 ± 0.3 mg cement
powder over a temperature range of 40–1000 °C. Dynamic modulus was calculated
by the ultrasonic pulse velocity test method. Bulk electrical resistivity was evaluated
by conducting electrochemical impedance spectroscopy on mortars and the results
are provided in a previous study [13].

3 Results and Discussion

Figures 2 and 3 show the linear expansion trends of the cement mortar systems
made with the two cements, which were exposed to ambient (23 °C) and heat-
cured (90 °C) cycles as described. It can be seen that the ambient-cured mortars
448 Y. K. Ramu et al.

Fig. 1 Representation of ambient and heat curing regimens for cement mortars

(C1-A and C2-A) showed no expansion. Cement mortars C1-H and C2-H (in as
received condition) satisfied Australian cement standards [7] and specifications [9]
by showing no expansion, despite being cured at 90 °C, likely due to the lesser
volume precipitation of DEF typically observed in low sulfate cements [14].
In contrast, the chemically modified cement mortars (C1SN-H, C2SN-H)
containing higher levels of sulfate (4%) and alkali (1%) showed significant DEF
expansion, aligning with other research studies [15, 16]. Of the chemically modified
cements, C2SN-H expanded more than C1SN-H, likely because of the higher C3 A
content of C2 (11.15%) compared with C1 (7.5%), which translates to higher DEF
and corresponding expansion. Furthermore, from the 6 month (0.5 year) expansion
data, the rate of expansion of the C2SN-H mortar was higher than the C1SN-H, also
likely due to the higher quantity of C3 A, which is known to accelerate hydration
reactions.
Although C1-H and C2-H did not show any significant DEF expansion, they may
have formed porous hydrates because of the accelerated heat-curing regimen. The
data in Table 2 support this hypothesis as the dynamic modulus and bulk resistivity
values were less than those for the ambient-cured systems. This finding infers poor
quality microstructure despite no observable expansion. For the highly expansive
mortars, C1SN-H and C2SN-H, dynamic modulus and bulk resistivity values were
found to be lower than for C1-H and C2-H, which indicated that other than the
formation of porous hydrates, expansive DEF may be causing microcracks and further
deteriorating the microstructure. Hence whether heat-cured cementitious systems
undergo expansive DEF or not, there will be durability issues. DEF only further
Submicroscopic Evaluation Studies to Minimize Delayed Ettringite … 449

Fig. 2 Linear expansion of cement 1 mortars over 3 years

Fig. 3 Linear expansion of cement 2 mortars over 3 years

450 Y. K. Ramu et al.

Table 2 Later-age
Mix Physical characteristics at 1 year
mechanical characteristics of
cement mortars Expansion (%) Dynamic Bulk resistivity
modulus (GPa) (Ω)
C1-A 0.006 46 30
C2-A 0.015 37 22
C1-H 0.067 25 9
C2-H 0.049 26 12
C1SN-H 0.729 21 7
C2SN-H 1.058 18 7
C1SNF-H 0.001 350 51
C2SNF-H 0.003 309 40

complicates the durability issue. Therefore, the solution for DEF mitigation must
also address mitigation of porous hydrates forming in these cementitious systems.
FA was included in this study to investigate its ability to mitigate expansive DEF
and restrict the formation of porous hydrates. Although there has been significant
research undertaken in understanding mitigation of DEF using FA [17–19], this study
is unique because we tested the ability of FA to mitigate DEF in high-sulfate and
high-alkali cements. As shown in Figs. 2 and 3, complete elimination of expan-
sive DEF occurred with additional FA in both C1SNF-H, and C2SNF-H mortars
despite the cement being spiked with 4% sulfate and 1% alkali. This result could be
partially due to a change in the dissolution behavior of the cement, thereby reducing
the overall content of aluminum and sulfate ions available in the pore solution.
Thus, less ettringite would precipitate in the cementitious system. From the data in
Table 2, the FA addition led to significant improvement in the dynamic modulus and
bulk resistivity values, contributing to a denser and high-quality microstructure. The
denser microstructure is likely due to the pozzolanicity of FA, which converts porous
(10 µm–10 nm) portlandite to denser (10–0.5 nm) calcium–silica–hydrate (C–S–H).
For comparison, the portlandite (calcium hydrate (CH)) content of the heat-cured
cementitious systems was also studied, as shown in Fig. 4. It is evident that the port-
landite content in C1SN-F was the lowest, which represents that the greater reaction
of CH with reactive silica in the FA formed additional C–S–H, thus making the
microstructure denser, in turn restricting the transportation of DEF-causing ions at
the submicron level and eliminating any increase in deleterious DEF occurrence.
Thus, it can be deduced that incorporating FA in heat-cured cementitious systems
eliminates DEF expansion and improves the microstructure by densification through
the production of more gel pores and less capillary pores due to the formation of
additional C–S–H. Based on our results, the following microstructural pattern (Fig. 5)
is hypothesized from the circuit model proposed by Guangling [20]. The model
considers concrete/mortar as a circuit when a potential difference is applied to the
sample. The free transportation of ions in the pore solution is controlled by the action
of continuous conductive paths. In contrast, blocked passages by the arrangement
Submicroscopic Evaluation Studies to Minimize Delayed Ettringite … 451

Fig. 4 Portlandite consumption of heat-cured mortars with and without fly ash

of a series of gel particles prevent the transfer of pore solution ions and are termed
“insulator paths”. If some gel particles exist in the conductive paths, then points of
discontinuation lead to discontinuous conductive paths. For further understanding of
the concrete/mortar circuit models, please refer to Guangling [20] and Ramu et al.
[13].
Figure 5a shows that the ambient-cured mortars (C1-A and C2-A) may have more
insulator paths than conductive paths due to the absence of heat-induced accelerated
curing, as reflected in their higher dynamic modulus and bulk resistivity values shown
in Table 2. However, for the heat-cured mortars (C1-H, C2-H, C1SN-H, and C2SN-
H), because of accelerated curing at 90 °C, microcracks may develop, which leads
to the gel particles not being tightly arranged and giving rise to more conductive
paths as represented in Fig. 5b. This scenario correlates well with the significant
reduction in dynamic modulus and bulk resistivity values noted in Table 2, which
creates a favorable situation whereby various ions such as Na, K, S, Ca, and Al are
readily transported. Although this scenario is the same for all heat-cured mortars
in our study, the C1SN-H and C2SN-H mortars contained more sulfates than C1-H
and C2-H. Therefore, greater transportation and deposition of sulfur ions at different
locations would increase the likelihood of DEF. Higher expansion of DEF causes
expansion and microcracks, which justifies the observation from Table 2 for further
reduction in dynamic modulus and bulk resistivity values.
With the incorporation of FA, greater C–S–H formation is inferred by the
consumption of (CH) portlandite (Fig. 4) and this creates a pore blocking effect
that leads to tightly packed gel particles creating more insulator paths, as repre-
sented in Fig. 5c. More tightly packed gel particles reduces the likelihood of ion
452 Y. K. Ramu et al.

Fig. 5 Schematic representation of different mortars’ microstructure (not to scale): a cement mortar
cured at ambient temperature (25 °C) [20]; b cement mortar cured at 90 °C; c cement mortar with
fly ash cured at 90 °C

transportation, thereby reducing ettringite precipitation and eliminating deleterious

DEF. Thus, the incorporation of FA in heat-cured mortars has two significant benefits:
(1) in highly DEF-prone cementitious systems, it can eliminate harmful expansive
DEF; and (2) in non-DEF-prone cementitious systems, it can be helpful in creating a
denser microstructure. Both these benefits contribute to more durable and sustainable
concrete.
Submicroscopic Evaluation Studies to Minimize Delayed Ettringite … 453

4 Conclusions

From this research work, the following conclusions can be made.

(1) Cements satisfying the requirements of AS 3972:2010 [7] and ATIC SP-43 [9]
for sulfate and alkali content are not prone to expansive DEF even if the concrete
is heat-cured. However, due to elevated temperature curing, porous hydrates
form, which leads to poor microstructural quality and eventual reduction in the
durability performance of the concrete.
(2) Chemically modified cement (sulfate 4%, alkali 1%) was prone to expansive
DEF when heat-cured. A reduction in physical characteristics such as dynamic
modulus and bulk resistivity was observed. Thus, if such cements are used in
concrete that is heat-cured, there is a high likelihood of durability issues arising
from the combined action of expansive DEF and permeability, leading to the
detriment of concrete service life.
(3) The incorporation of FA eliminated expansive DEF even in chemically modified
cement mortars (sulfate 4%, alkali 1%). Furthermore, the consumption of port-
landite to form additional C–S–H resulted in a much denser microstructure, as
inferred from the high values for dynamic modulus and bulk resistivity. Thus, the
addition of FA undoubtedly contributes to increasing the durability of heat-cured
(precast) concrete, thereby also contributing to enhanced sustainability.
(4) The conclusions presented in this study are limited to the two Australian general-
purpose cements used in the research and mortar specimens. This study needs
to be extended to a range of cements/clinkers and concrete specimens.

Acknowledgements We acknowledge the financial support provided by Holcim Australia and

the award of a UTS international research scholarship for carrying out this research project. The
Construction Materials Research Laboratory, IIT Madras, is gratefully acknowledged for laboratory
support to do part of the work presented in this report.

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A Novel Concrete Mix Design
Methodology

D. Kumar, M. Alam, and J. Sanjayan

Abstract Concrete mix design is the methodology for mixing binder, aggregate and
water to achieve required physical, mechanical, and thermal properties. In particular,
the physical properties depend on the volume fraction of each element in the concrete
recipe. In this study we considered cement mortar, complying with ASTM C105, as
the reference concrete with cement as the binder and silica sand as the aggregate. The
reference mortar was denser with high thermal conductivity and compressive strength
at given rheological properties. A denser concrete presents difficulty in material
handling and imposes a safety risk, and high thermal conductivity increases building
energy consumption. Therefore, lightweight concrete (LWC) has been developed by
replacing silica sand with porous materials. LWC includes cement as the binder, with
silica sand and other porous materials as the primary and binary fillers. The mass
of the filler materials is determined by their particle density and volume fraction.
LWC has low thermal mass, thereby exacerbating the summertime overheating and
peak cooling demand of buildings. Therefore, there is a need to design a LWC
with high thermal mass by incorporating phase change materials (PCM), which are
mainly incorporated as tertiary filler. Here, we propose a novel concrete mix design
methodology to incorporate PCM composite as a partial replacement of the porous
material without changing binding materials.

Keywords Capric acid · Cement mortar · Expanded perlite · Mix design

methodology · Silica aerogel

D. Kumar · M. Alam (B) · J. Sanjayan

Centre for Sustainable Infrastructure and Digital Construction, Department of Civil and
Construction Engineering, Swinburne University of Technology, Hawthorn, VIC, Australia
e-mail: malam@swin.edu.au
D. Kumar
Department of Mechanical Engineering, Mehran University of Engineering and Technology,
(SZAB Campus), Khairpur Mir’s, Sindh, Pakistan

© The Author(s) 2023 457

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_46
458 D. Kumar et al.

1 Introduction

Energy use in buildings is increasing due to population growth, changes in lifestyle,

and urbanization, and it accounts for 30–40% of the total energy use in many
countries [1]. In buildings, the heating, ventilation, and air conditioning (HVAC)
system is responsible for 42–68% of the total energy use in building to maintain
acceptable levels of thermal comfort [2, 3]. Energy consumption by the HVAC
system depends on internal heat gain due occupancy, lighting and appliances, and
heat transfer through the building envelope (50–60% of HVAC system energy
consumption). The latter is reduced by using various insulating materials to reduce
building energy consumption by 20–80% [4]. However, inappropriate use of insula-
tion materials accounts for 30% of heat loss in buildings [5]. Traditional insulation
materials such as extruded polystyrene boards, expanded polystyrene panels, glass
felt, and polyurethane foam have lower resistance, compressive strength, and poor
hydrophobicity [4].
Lightweight concrete (LWC) was developed using silica aerogel (SA) as an
alternative insulating material with comparable thermal conductivity (0.026–0.37
W/m–K) to foam and autoclave aerated concrete, high compressive strength, and
high fire resistance [6]. Ibrahim et al. [7] reported that the thermal performance of
SA-based LWC materials was better than that of insulation. Gao et al. [8] devel-
oped an aerogel-incorporated concrete by replacing 10–60 vol% of sand with SA,
having thermal conductivity and compressive strength of 0.26 W/m–K and 8.3 MPa,
respectively. Fickler et al. [9] used SA granules instead of SA particles to develop
LWC with minimum thermal conductivity and compressive strength of 0.16 W/m–K
and 3.0 MPa, respectively. Suman et al. [10] developed ultra-LWC using expanded
glass particles, SA particles and prefabricated plastic bubbles to replace the weight
percentage of cement. They found a 42% decrease in compressive strength at 15%
decrease in cement content with SA particles. The thermal conductivity of the cement
mortar was reduced by 60, 53 and 48% by replacing 20 wt% of fine aggregates
with SA, expanded perlite and vermiculite, respectively. They achieved the same
thermal performance to the reference coating material by reducing its thickness by
60% with the addition of SA [11]. Li et al. [12] used alkali-activated binder instead
of cement because of its low thermal conductivity. They developed an ultra-LWC
with thermal conductivity equivalent to insulating materials (0.0434 W/m–K) with
minimum compressive strength of 0.38 MPa at 30 °C. The volumetric heat capacity
and thermal conductivity of the SA-LWC was lower than those of the reference
concrete by 0.17–0.227 J/m3 -K and 1.22–1.27 W/m–K, respectively. The thermal
mass of LWC is lower than that of normal concrete because of its low density and
high porosity [13].
Some researchers have integrated form-stable phase change material (FSPCM)
into expanded perlite [14], expanded clay [15] and hollow ceramsite [16]-based
LWCs to develop low thermal conductivity and high thermal storage compos-
ites. They first developed LWC composites by volumetrically replacing sand with
lightweight fillers, the mass of which was calculated using a particle density approach.
A Novel Concrete Mix Design Methodology 459

Fig. 1 Research
methodology. Note: FSPCM,
form-stable phase change
Development
material; LWC, lightweight
concrete . Binder . Physical
. Fillers . LWC . Thermal
. FSPCM-
Material LWC
Properties
information

The FSPCM composites were added as a mass fraction of dry mixture [14] and total
aggregate [15, 16]. In these mixing and trial methods [17], the mass of binder and
of aggregate was changed to those of the integrated FSPCM composite, resulting in
a significant loss of materials to achieve desirable thermophysical properties. Han
et al. [18] collected data regarding concrete laboratory waste for the period of 2011–
17. They found that a cylindrical specimen was annually producing 50 m3 (61%)
concrete waste, followed by 18% of tensile specimen, 15% cubes and remaining
pavement block and mortar specimen. Annually, the concrete and cement waste was
80 m3 and 20 m3 , respectively, with CO2 emissions >15 m3 . Therefore, there is
need to develop a scientific methodology to design concrete mixtures for achieving
desirable thermophysical properties and performance without wasting materials.
In this study we aimed to develop a novel concrete mix methodology using the
particle density of lightweight filler and FSPCM (Fig. 1). This study considered
SA granules as the filler. The FSPCM is made up of capric acid and hydrophobic
expanded perlite [19]. We also investigated the thermophysical properties of the LWC
and FSPCM-LWC composites by considering bulk density, compressive strength,
thermal conductivity, and latent heat storage.

2 Methods

2.1 Materials

Ordinary Portland Cement (OPC; purchased from Bunnings, Australia) was used
as the binder in accordance with AS3972. Silica sand was selected as the aggre-
gate for the reference concrete in compliance with ASTM C105. SA granules
(SAG; purchased from Enersen, France) was the lightweight aggregate for the LWC
composite. Finally, CAHEP was used to develop the FSPCM-LWC composite, as
described previously [19, 20] (Table 1).
460 D. Kumar et al.

Table 1 Properties of different aggregates

Property Units Sand CAHEP SAG
Surface – Hydrophilic Hydrophobic Hydrophobic
Particle Density kg/m3 2657 600 120
Particle size Å 327,376 35 69
Thermal conductivity W/m.K NA 0.38 0.024
Latent heat kJ/kg – 95 –
CAHEP, capric acid/hydrophobic expanded perlite; SAG, silica aerogel granules

2.2 Development of LWC and FSPCM-LWC Composite

Our proposed novel concrete mix design methodology for developing LWC and
FSPCM-LWC composites is shown in Fig. 2. The reference concrete was according
to ASTM C105 regarding cement, water and sand as a binder, activator, and primary
filler, respectively. At a given quantity of binder, the sand-to-cement and water-to-
cement ratios were 2.75 and 0.485, respectively. LWC was developed by volumet-
rically replacing the sand particles (primary filler) with SAG (secondary filler). The
complete replacement of sand with SAG makes SAG the primary filler due to the
presence of only one filler. The primary filler is replaced by a secondary filler from
top to bottom at a given mass of cement and sand. The secondary filler (SAG) is
volumetrically replaced by a tertiary filler (CAHEP) to develop the FSPCM-LWC
composite without changing the primary filler (sand) from left to right. When the
secondary filler (SAG) is completely replaced by CAHEP, the developed concrete
has sand as the primary filler and CAHEP as the secondary filler.

Binder Fillers

Primary (silica sand) ASTM C105

Binary (Sand & Tertiary (Silica Sand, Silica Binary (Silica

Cement

Water

Silica Aerogel) Aerogel & FSPCM) Sand & FSPCM)

Primary (Silica Binary filler (Silica Aerogel Primary

Aerogel) & FSPCM) (FSPCM)

Fig. 2 Novel concrete mix design methodology. PCM, phase change material
A Novel Concrete Mix Design Methodology 461

The reference concrete composite was developed using a water to OPC ratio of
0.485 and silica sand to OPC ratio of 2.73, in accordance with ASTM C105. The LWC
composite was developed by volumetrically replacing silica sand with SAG at 20–
80%. The mass of sand and SAG was calculated using Eq. 1 and Eq. 2, respectively. To
develop the CAHEP-LWC composite, the mass of SAG was volumetrically replaced
by CAHEP at 20–80% without changing the mass of sand. The mass of CAHEP was
calculated using Eq. 3.

m Sand = ρsand Vsand (1)

ρ S AG m sand (Vsand − VS AG )
m S AG = (2)
ρsand Vsand
ρC AH E P m S AG (VS AG − VC AH E P )
m C AH E P = (3)
ρ S AG VS AG

where, m, ρ and V show the mass (kg), density (kg/m3 ) and volume (m3 ) of sand,
SAG and CAHEP composites, respectively, to develop the reference concrete, LWC
composite, and FSPCM-LWC composite. The workability of the LWC and FSPCM-
LWC composites was kept the same as that of the reference concrete by using a
superplasticiser. The mass of OPC, water, and the calculated mass of fillers are given
in Table 2.

Table 2 Mix design recipes

Notation Binder Activator Filler Superplasticizer
OPC Water Sand SAG CAHEP
kg kg kg kg kg %
RC 520 252 1430 0 0 –
LWC1 520 252 1144 13 0 –
CAHEP-LWC1 520 252 1144 10 13 –
CAHEP-LWC2 520 252 1144 8 26 –
CAHEP-LWC3 520 252 1144 5 39 –
CAHEP-LWC4 520 252 1144 3 52 –
LWC2 520 252 172 52 0 –
CAHEP-LWC5 520 252 172 41 52 1.2
CAHEP-LWC6 520 252 172 31 103 3.7
CAHEP-LWC7 520 252 172 21 155 5.0
CAHEP-LWC8 520 252 172 10 207 6.0
CAHEP, capric acid/hydrophobic expanded perlite; LWC, lightweight concrete; RC, reference
concrete; SAG, silica aerogel granules
462 D. Kumar et al.

Fig. 3 Techno-test machine a and test specimen b

2.3 Properties of LWC and FSPCM-LWC Composites

2.3.1 Physical Properties

Cubic specimens were cast using our mix design recipes in 50 × 50 × 50 mm metallic
molds, in accordance with ASTM C1009. Three specimens were cast for each mix
design for precision and accuracy of results. They were demolded after 24-h curing
in an environmental chamber at temperature and relative humidity of 23 °C and 90%,
respectively. The demolded specimens were water cured until the test date. The mass
of the cubes was measured by a simple balance with accuracy of 0.1 g. The techno-
test machine with accuracy of 0.1 kN was used to measure the compressive strength
of the test specimens, as shown in Fig. 3.

2.3.2 Thermal Properties

Thermal properties include thermal conductivity and latent heat storage. Thermal
conductivity of the developed cementitious composites was measured using a tran-
sient line source (TLS-100), complying with ASTM D5334, as shown in Fig. 4.
A 50-mm diameter cylinder of length 120 mm was cast and demolded after 24-h
curing in an environmental chamber. The samples were air dried in the environmental
chamber at temperature and relative humidity of 23 °C and 50%, respectively, until
the test date. The TLS-100 probe was inserted into the test specimen and kept for
15 min to achieve thermal equilibrium between the specimen and probe surface. Final
measurements by executing the test were obtained using a digital display meter.
A Novel Concrete Mix Design Methodology 463

Fig. 4 Thermal conductivity meter a and sensor b

Differential scanning calorimetry provided inaccurate data due to the smaller

sample size of 5–20 mg. The mass fraction of each material in the developed compos-
ites could not be accurately balanced in such a small sample. Consequently, the latent
heat storage of the FSPCM-LWC composite (h F S PC M−L W C ) was calculated using
Eq. 4 [20].

m F S PC M h F S PC M
h F S PC M−L W C = (4)
m w + m O PC + m S + m F S PC M

where, m w , m O PC , m S and m F S PC M are the mass of water, OPC, sand and FSPCM,
respectively, and h F S PC M is the latent heat storage of the FSPCM composite.

3 Results and Discussion

Figure 5 shows the density of the LWC and CAHEP-LWC composites. The density
of the reference concrete was 2226 kg/m3 and that of the composites was 9 and 40%,
respectively, lower than the reference concrete due to replacement of 20 and 80%
volume of silica sand with SAG, because the particle density of SAG is 20-fold lower
than that of silica sand. The effect of CAHEP on density depends on the presence
of SAG. For instance, the density of LWC with 20 vol% of SAG was only 3%, but
the density of LWC with 80 vol% of SAG was maximally increased by 30% with
the addition of CAHEP. Thus, the density of LWC increases dramatically at higher
volume fractions of the lightweight filler. LWC must contain the highest proportion
of lightweight fillers as given binding materials.
464 D. Kumar et al.

2400
2200

Density (kg/m3) 2000

1800
V20 V80
1600
1400
1200
1000

Replacing SAG with CAHEP composite

Fig. 5 Effect of capric acid/hydrophobic expanded perlite (CAHEP) on bulk density of lightweight
concrete (LWC)

30
Compressive strength (MPa)

25

20

15
7-days CS (V20) 7-days CS (V80)
10 28-days CS (V20) 28-days CS (V80)

Replacing SAG with CAHEP composite

Fig. 6 Effect of capric acid/hydrophobic expanded perlite (CAHEP) on compressive strength of

lightweight concrete (LWC)

Figure 6 exhibits the compressive strength of LWC and CAHEP-LWC. The

compressive strength of 31 and 46 MPa were measured for the reference concrete after
7 and 28 days, respectively, of water curing. The addition of 20 vol% and 80 vol% of
SAG as a partial replacement of sand reduced compressive strength by 56% and 81%,
respectively. The decrement percentage of density was smaller than the compressive
strength even at the same proportion of SAG because SAG have smaller particle
A Novel Concrete Mix Design Methodology 465

Thermal conductivity (W/m-K)

1.6

1.2
V20 V80
0.8

0.4

Replacing SAG with CAHEP composite

Fig. 7 Effect of capric acid/hydrophobic expanded perlite (CAHEP) on thermal conductivity of

lightweight concrete (LWC)

density and more open porous structure, making SAG more fragile and breaking at
very small loading. Moreover, SAG have an amorphous shape, promoting heteroge-
neous porosity, and resulting in lower compressive strength. The addition of CAHEP
increased the compressive strength of the LWC by filling the large pores of the
LWC. The compressive strength hardly changed with the addition of CAHEP to the
LWC with 20 vol% of SAG. At higher volume fractions of SAG, the compressive
strength of LWC doubled with the addition of CAHEP. The developed LWC and
LWC-CAHEP composites both had higher than the minimum compressive strength
(4.14 MPa) required for nonload-bearing structural material as required by ASTM
C129–17 [15].
Thermal conductivity and theoretical latent heat storage of LWC and CAHEP-
LWC are shown in Figs. 7 and 8, respectively. Thermal conductivity of the refer-
ence concrete was 2.27 W/m–K, which decreased by 60 and 93% by adding 20
vol% and 80 vol% of SAG, respectively. The effect of SAG on thermal conductivity
was higher than on density and compressive strength because of the 100-fold lower
thermal conductivity of SAG (0.01–0.02 W/m–K [5, 18]) compared with sand, and
the creation of macro-porosity due to the heterogeneous porous structure of SAG.
The addition of CAHEP almost increased thermal conductivity of the LWC from 0.92
to 1.87 W/m–K and from 0.16 to 0.532 W/m–K at 20 vol% and 80 vol% of SAG,
respectively. The addition of CAHEP dramatically increased the thermal conduc-
tivity of the LWC due to the higher thermal conductivity of CAHEP (0.38 W/m–K)
and its smaller particle size which filled the macro-porosity of the LWC. The latent
heat storage increased linearly with CAHEP proportion. The reference concrete and
466 D. Kumar et al.

30

Latent heat storage (J/g)

25

20

15

10

Replacing SAG with CAHEP composite

Fig. 8 Effect of capric acid/hydrophobic expanded perlite (CAHEP) on latent heat storage of
lightweight concrete (LWC)

LWC stored heat directly, whereas the CAHEP-LWC composite stored latent heat
by changing the material’s phase. Heat storage increased by 3.3 kJ/kg and 28 kJ/kg
at 20 vol% and 80 vol%, respectively, of SAG replaced by CAHEP (Fig. 8).

4 Conclusions

We proposed a new concrete mix design methodology for developing tertiary filler
cementitious composite to meet structural and thermal properties. We considered
three fillers—silica sand, silica aerogel granules and CAHEP composites—with
different particle densities. Silica sand was used for high compressive strength of
cementitious composites, but the addition of SAG decreased the thermal conduc-
tivity of the reference concrete to an acceptable level for non-load bearing appli-
cation in buildings. The developed LWC with the lowest thermal conductivity had
low thermal mass due to the high porosity of SAG. To increase the thermal mass of
the LWC, CAHEP partially replaced the volume fraction of SAG. The mass of the
LWC-CAHEP composite was determined by a particle density approach instead of
the hit and trail mixing methodology. The particle density-based mix design method-
ology revealed that the high fraction, denser filler increased compressive strength,
thermal conductivity, and density, whereas the porous material reduced thermophys-
ical properties. Thermophysical properties are a function of particle density, surface
morphology and porous structure. The use of a denser filler increases compressive
strength at given binding materials. A hydrophobic surface results in lower compres-
sive strength due to less affinity with cement paste. Finally, an amorphous porous
structure promotes macro-porosity, reducing the strength of the concrete. Future
A Novel Concrete Mix Design Methodology 467

studies should consider the structure and surface morphology of fillers to investigate
optimal mass fractions of silica sand, SAG and CAHEP in thermally enhanced LWC
composites.
The selection of the CAHEP-LWC composite depends on thermal conductivity
and latent heat storage. Both properties are essential for energy-efficient building
design. We found that increasing the thermal storage of LWC and increasing its
thermal conductivity was undesirable due to heat transfer. Therefore, there is a need
to conduct a sensitivity analysis to investigate the optimum thermal conductivity
and latent heat storage of heat resistive and storage panels to design energy efficient
buildings.

Acknowledgements This research was funded by the Higher Education Commission (HEC),
Pakistan and Swinburne University of Technology, Australia with grant no. 5-1/HRD/HESTPI/
(Batch-VI)/6021/2018/HEC.

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Characterization of the Nano-
and Microscale Deterioration Mechanism
of the Alkali–Silica Reaction in Concrete
Using Neutron and X-ray Scattering
Techniques: A Review

E. Nsiah-Baafi, M. J. Tapas, K. Vessalas, P. Thomas, and V. Sirivivatnanon

Abstract Alkali–silica reaction (ASR) is one of the most recognized chemical reac-
tions that lead to the deterioration and premature failure of concrete. The severity
of ASR is largely dependent on the expansive nature of the reaction product (ASR
gel). As such, it is important to expound the developed knowledge on the forma-
tion, structure, composition, and swelling mechanism of ASR gel, to provide a
greater understanding of ASR deterioration and to facilitate the development of
more reliable prediction and mitigation methods. We present a summary of existing
methods for assessing ASR and the state-of-the-art techniques that use neutron and
X-ray scattering methods to characterize the nano- and microstructural properties of
concrete and elucidate the potential transport dynamics of reactants that determine
the mechanism and extent of ASR.

Keywords Alkali–silica reaction · Characterization · Microstructure ·

Nanostructure

1 Introduction

The alkali–silica reaction (ASR) is a deleterious chemical reaction that occurs in

concrete when certain reactive silica phases in aggregates react with alkali ions (Na,
K) from the pore solution [1–3]. Through decades of research the mechanism of
ASR is relatively well documented, but controversy remains regarding the reported
reaction sequence leading to the formation of ASR gel [4, 5] and the expansion

E. Nsiah-Baafi (B) · M. J. Tapas · V. Sirivivatnanon

Innovation Factory, UTS-Boral Centre for Sustainable Building, Sydney, NSW, Australia
e-mail: elsie.nsiah-baafi@uts.edu.au
K. Vessalas · P. Thomas · V. Sirivivatnanon
School of Civil and Environmental Engineering, University of Technology Sydney (UTS),
Sydney, NSW, Australia

© The Author(s) 2023 469

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_47
470 E. Nsiah-Baafi et al.

mechanism of the gel to form cracks in concrete [6]. Undoubtedly, ASR begins with
the dissolution of reactive silica in the alkaline pore solution of concrete and the
dissolution rate depends on several factors including the degree of alkalinity of the
pore solution, the type and particle size of the silica mineral present, temperature, and
the presence of other cations such as Li+ , Al3+ and Ca2+ in the pore solution [1, 2].
Following silica dissolution, Ca2+ and alkali cations in the pore solution react with
the dissolved silica to form a C-(Na, K)-S–H reaction product known as ASR gel, as
well as other calcium-rich hydrates [2]. The composition and structure of the ASR
gel may vary in the same concrete system, and in one concrete system from the other,
depending on the location of the reaction site, the type and amount of silica and cations
at the reaction site, and the age and curing conditions of the concrete. The contention
in the sequence of ASR gel formation stems from an earlier proposal described by
Hou et al. [4]. By comparing ASR gels in laboratory specimens to gels in field
structures, those authors concluded that the continuous formation of a calcium-rich
C-S-H product occurs when calcium is locally available at the reaction site [4]. This
initial C-S-H is typically dense and acts as a physical barrier that isolates the reaction
sites in the concrete structure [7]. Upon depletion of calcium at the localized sites, the
concentration of silicon increases until a low calcium and high alkali C-(Na,K)-S-H
ASR gel is formed at silicon saturation. Although this sequence of ASR gel formation
has been supported by other studies [4, 8, 9], it is worth noting that the results were
obtained from batch experiments using model reactant methods; thus, they may not
be representative of actual concrete systems. Furthermore, a recently reported study
[10] demonstrated that ASR gel may first form in cracks on the aggregate surface
and around the aggregate, then penetrate towards the inside of the aggregate, which
suggests that a C-S-H physical barrier may not be evident in the sequence of ASR
gel formation. Moreover, once the ASR gel is formed, the mechanism by which it
expands upon moisture absorption is still a topic of discussion [6, 11]. It is, however,
well recognized that the addition of supplementary cementitious materials (SCMs)
such as fly ash and slag effectively mitigates ASR in concrete [1, 12–14]. With
the current depletion of these conventional SCMs, several studies are emerging to
discover and optimize potential alternative SCMs and techniques for mitigating ASR.
Understanding the dynamics of ASR gel formation, the transport of the gel through
the concrete structure, the expansion mechanism of the gel and the effects of its
expansion at the micro-and nanoscale is ultimately the key to developing effective
mitigation against deleterious ASR. Currently, a number of techniques exist for
assessing the structure and composition of ASR gel. In this paper we recap some
of the reported studies on ASR characterization methods and present the state-of-
the-art techniques that use neutron and X-ray scattering to identify the micro- and
nanostructure of concrete to characterize additional features and propagation of ASR
in concrete.
Characterization of the Nano- and Microscale Deterioration Mechanism … 471

2 Traditional Characterization Techniques for Assessing

ASR in Cement-Based Materials

Microscopic techniques such as petrography and scanning electron microscopy

(SEM) are by far the most commonly used to identify ASR gel in concrete and
other cement-based materials [15–18]. These methods can be used in conjunction
with other techniques such as electron dispersive spectroscopy (EDS) [17], electron
backscattered diffraction [19], Raman spectroscopy [20], nuclear microwave reso-
nance (NMR) [21] and nonlinear impact resonance acoustic spectroscopy [22] to
identify the morphology, composition and effect of ASR and its reaction products
on concrete. A typical example of petrographic characterization of an ASR-affected
mortar and concrete is presented in Fig. 1. Other less commonly used techniques that
have shown proficiency in providing information on the nanostructure of cement-
based materials and ASR are scanning transmission electron microscopy (STEM)
[23, 24], atomic force microscopy [25] and scanning confocal microscopy [26].
Studies carried out by Leeman et al. [10, 27] to characterize the structure and
composition of ASR gel formed in concrete using a combination of SEM–EDS,
Raman spectroscopy and 29 Si NMR demonstrated that both crystalline and amor-
phous ASR products form in concrete aggregates. To follow the sequence of ASR gel
formation and structure of the ASR products at the nanoscale, the studied concrete
mixes were doped with CsNO3 and KNO3 tracers [10]. The authors established that
ASR gel first forms in pre-existing pores or cracks in aggregates close to the cement
paste and is generally amorphous in structure. The initial cracks that result from the
expansion of the formed gel are mostly empty; therefore, secondary ASR products,
identified as crystalline ASR gel, begin to fill the cracks while progressing from the
aggregate–cement interface to the aggregate’s interior. Further characterization in
that study revealed that both the amorphous and crystalline ASR products have a
structure dominated by Q3 -sites; however, there is a difference in their composition

Fig. 1 Petrographic images of a concrete specimen showing ASR gel lining an air bubble that
is thinly outlined by low birefringent ettringite and b ASR-filled crack passing through an acid
volcanic fragment and a silica-depleted (porous) acid volcanic fragment
472 E. Nsiah-Baafi et al.

such that the amorphous product has a higher Na/K ratio. Additionally, ASR prod-
ucts in concrete samples cured at 38 °C exhibited a similar structure to ASR gel in
field concretes, whereas at temperatures above 50 °C, a K-shylkovite structure was
observed [27]. A similar observation has been reported in other studies [28, 29]. The
difference in the ASR gel structure with temperature potentially contributes to the
expansion capacity of the gel. This outcome supports that temperature is a signifi-
cant factor to consider when selecting a suitable accelerated test method for assessing
ASR, such as the accelerated concrete prism test and accelerated mortar bar test, and
rationalizes the differences observed in the laboratory and field reactivity predictions
of some aggregates.
Although these characterization methods are effective in assessing the microstruc-
ture and composition of ASR and other cement reaction products, they are generally
destructive methods that depend on proper representative sampling from bulk mate-
rial and somewhat rigorous sample preparation, which may influence the outcome of
the characterization studies. For example, there are reports in the literature that cite
instances where the sectioning and polishing of suspected ASR-affected samples for
SEM analysis may have resulted in dislodgement of the ASR gel from cracks [30, 31].
Similarly, to obtain TEM lamellae for nanostructural characterization, cement-based
samples are usually milled down to a low micron thickness (<3 µm) to allow the
penetration of electron beams [24]. This potentially affects the structure of the formed
product. The destructive nature of these methods also inhibits time-lapse character-
ization and continuous or in-situ monitoring of ASR development and crack propa-
gation in the same region of the bulk sample. This is a major drawback to improving
insight on the transport of reactants in concrete and identifying features of ASR that
can contribute to the development of novel mitigation strategies.

3 X-ray and Neutron Scattering Techniques

Cement-based materials are characteristically porous. Pores play a key role in the
durability and mechanical performance of concrete. For one, they act as a conduit,
thus determining the extent of permeation of chemical agents and triggers of deteri-
oration mechanisms. For example, during ASR, moisture containing alkali and other
solutes may ingress from the service environment of the concrete structure or be
transported from one region to the other within the concrete system through pores.
Furthermore, as ASR gel takes up water and expands to form cracks, the propagation
of the cracks provides a channel for the spread of less rigid, high-expansive ASR
gel [32] through the concrete, promoting ASR. These cracks also become potential
sites for the repolymerization and crystallization of new ASR products. However,
it is worth noting that pores may be closed (air bubble) or open with a network of
micro- and nanosized distribution. These features of the pore system are critical in
understanding the influence and extent of porosity on the durability and strength of
a concrete structure. For instance, during salt attack, the crystallization pressure of
salts in the concrete will vary with pore size. Similarly, the mechanism of drying
Characterization of the Nano- and Microscale Deterioration Mechanism … 473

shrinkage in the concrete is dependent on pore size and relative humidity such that
in larger pores with higher relative humidity, capillary pressure is the driving force
for drying shrinkage [33, 34]. Considering that the nature of the pore system in
concrete has a major influence on its durability properties, destructive characteriza-
tion techniques that generally sample thin sections from bulk concrete material may
not provide precise information on pore features, including size, volume fraction,
distribution, and network.
In the past decade, there has been a significant increase in the use of non-
destructive neutron and X-ray scattering techniques, such as ultra-/small-angle
neutron scattering (USANS, SANS), small angle X-ray scattering, X-ray computed
tomography (X-ray CT) and neutron tomography, to characterize the micro- and
nanostructure of the concrete and particularly to elucidate the chemo-poromechanics
of ASR reactants through time-lapse damage evolution monitoring [35–37]. USANS
and SANS use the elastic scattering of neutrons passing through a sample (neutron
diffraction) to study the atomic structure of the bulk material and determine structural
inhomogeneity at the mesoscopic scale length, typically ranging from 1 to 300 nm
[38]. This technique is similar to X-ray diffraction as both principles obey the Beer
Lambert law [39]. However, neutrons are unaffected by electrons, therefore when
encountering matter, they penetrate to interact with the atomic nuclei whereas X-
rays intermingle with the electron cloud around the atom [40]. As such, neutrons
are relatively more sensitive to atoms with low atomic number such as hydrogen.
This explicates the proficient use of neutron scattering techniques to characterize
ASR gel (C-S-(Na,K)-H) and other cement hydrated reaction products, as well as
the transport of reactant through the pores in the pore solution of concrete structures.
In tomography, neutron and X-ray beams passing through a sample are attenuated
according to the sample’s composition and geometry. A series of transmission images
(tomographs) representing slices of the sample at several rotation angles are gener-
ated. These tomographs can be superimposed to form a 3D representation of the bulk
sample showing the surface and internal features in respective volumetric locations
[35, 41]. Typically, incident neutrons will provide a high imaging contrast for hydrates
in concrete and a good transmittance of metals (e.g., steel reinforcements), whereas
X-rays display a high contrast for metals and an adequate transmittance for other
light element materials. Therefore, the information obtained from both scattering
techniques is complementary. In characterizing concrete structures, the combined
use of neutron and X-ray diffraction for imaging has proven to be very efficient for
investigating the pore system, propagation of cracks, the presence and dynamics of
reaction products, and monitoring of concrete reinforcement materials. For example,
the neutron imaging facility at the Paul Scherrer Institute in Switzerland (NEUTRA)
has in the past 7 years shown a development in imaging techniques by mounting an
X-ray tube before the initial collimator to enable characterization of samples with
X-rays and neutrons under the same geometric conditions and the use of an identical
detection system [42]. A typical application of this bimodal approach to investigate
the internal structure of stainless steel fiber-reinforced concrete is presented in Fig. 2.
474 E. Nsiah-Baafi et al.

Fig. 2 Neutron and X-ray tomography of a concrete sample [42]

In Fig. 2, it can be seen that the neutron tomographic data set provides information
on the segregation of aggregates and pore network in the concrete, whereas the X-
ray tomographic results show detailed information on the steel reinforcement. In
another study, the authors reconstructed a 3D microstructure of concrete using finite
modelling by stacking 2D slices of the segmented concrete’s microstructure obtained
from both neutron and X-ray CT [41]. The reconstructed 3D microstructure, shown
in Fig. 3, revealed the distribution and volume fraction of pores, aggregate and paste
binder, as well as the sizes of the respective constituents.
Recently, the inclusion of nanomaterials in construction such the addition of
carbon nanotubes as reinforcement in concrete to increase fire resistance, reduce
porosity and improve strength properties, and the addition of novel nano- and micro-
sized SCMs to improve strength and reduce ASR in concrete have become the focus
of emerging research [43, 44]. The application of nondestructive neutron and X-
ray techniques to understand how these new materials alter the microstructure of
concrete to deliver the desired properties and performance is undoubtedly crucial in
accomplishing such ground-breaking innovations.

Fig. 3 Segmented aggregate and void phases obtained from a X-ray computed tomography (CT),
b neutron CT, and c combined CT [41]
Characterization of the Nano- and Microscale Deterioration Mechanism … 475

4 Conclusions

The characterization of ASR in concrete using microscopic and spectroscopic tech-

niques have to date provided useful information that has contributed to the under-
standing of the mechanism of ASR and the development of mitigation measures.
However, due to the destructive nature of these conventional techniques, crucial
microstructural features that influence the dynamics, progression, and extent of ASR
with age under different conditions are unobserved. To obtain these characteris-
tics, non-destructive X-ray and neutron scattering techniques have shown excellent
results, providing 3D volumetric insights into the morphology, interfacial bonding
and interaction between constituents, and the pore system, which has a key effect
on the durability of concrete. Nonetheless, X-ray and neutron scattering techniques
usually do not provide information on the composition of the material. With advance-
ments in the engineered materials used in construction, the ultimate characterization
tool kit that combines microscopy and spectroscopy with neutron and X-ray scat-
tering techniques will provide the best approach to understanding and optimizing the
performance of concrete for modern structures.

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the copyright holder.
Life Cycle Assessment
of the Environmental Impacts of Virgin
Concrete Replacement by CO2 Concrete
in a Residential Building

M. Ma, Y. Zhou, V. W. Y. Tam, and K. N. Le

1 Introduction

Concrete is one of the most consumed materials in construction, with 25 billion tons
produced globally per year [1]. However, it is considered as the most non-sustainable
material. The acquisition of virgin aggregate consumes a considerable amount of
energy and emits a large amount of greenhouse gasses [2]. Recycled aggregate from
construction and demolition (C&D) waste could be a viable substitution in concrete
production, both avoiding landfills and conserving natural resources [3]. Recycled
aggregate could replace part or all virgin aggregate in concrete and the product
is referred to as “recycled concrete” [1]. Although recycled concrete containing
recycled aggregate is considered as comparable to virgin concrete, it is not widely
accepted by the industry, because of uncertainty about material performance [4].
Recycled concrete has lower mechanical properties and higher shrinkage and creep
than virgin aggregate with the same mix design [1]. In order to improve the properties
of recycled concrete, CO2 gas is injected into the recycled aggregate and the CO2 -
treated aggregate is mixed into concrete as normal. This new concrete is known
as “CO2 concrete” and rivals the virgin concrete in its mechanical and durability
qualities.
Environmental performance of the concrete has attracted increasing attention from
academics [5]. Marinkovic et al. [1] summarized two research focuses on sustainable
solutions for concrete production: (1) using recycled aggregate to partly or entirely
replace virgin aggregate, and (2) replacing cement with cementitious materials [1].
Life cycle assessment is a commonly used tool to evaluate the environmental impact
of a product [5]. Specifically, Xing et al. [6] compared the environmental benefits of

M. Ma · Y. Zhou · V. W. Y. Tam (B) · K. N. Le

School of Design, Engineering and Built Environment, Western Sydney University, Penrith,
NSW, Australia
e-mail: vivianwytam@gmail.com

© The Author(s) 2023 479

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_48
480 M. Ma et al.

virgin concrete, recycled concrete and CO2 concrete, and found that CO2 concrete
was the best-performing product for greenhouse gas reduction, because the CO2
was retained in the recycled aggregate during the carbonation process [6]. Residen-
tial buildings use a wide range of resources in their construction, including a great
amount of concrete [7]. The environmental performance of a residential building
using CO2 concrete as a partial replacement for virgin concrete remains unknown, so
our aim was to conduct a lifecycle assessment to evaluate the environmental impact of
CO2 concrete as a replacement in a residential building. Specifically, virgin concrete
replaced by 0, 30, 50, 75 and 100%.

2 Methods

To fulfil the aim of this study, a building information modeling (BIM) and life cycle
assessment integration program was used to conduct the life cycle assessment of a
building in five scenarios where 0, 25, 50, 75 and 100%, respectively, of the virgin
concrete was replaced by CO2 concrete. The life cycle assessment was conducted
according to the ISO 14040 framework, which provides a standard process of four
phases, namely, goal and scope definition, life cycle inventory analysis, life cycle
assessment analysis and life cycle interpretation phases. In this study, the process
started with creating a BIM of a residential building as the goal and scope defi-
nition phase. The lifespan of the building was assumed to be 50 years. The anal-
ysis accounted for the full cradle-to-grave life cycle of the building studied across
all stages, including material manufacturing, transportation, building construction,
maintenance and replacement, and eventual end of life. In the life cycle inventory
analysis phase, the bill of quantities for each building component was extracted
from the BIM, and the life cycle inventory data of each component was retrieved
from the GaBi 2018 databases, the Australian life cycle inventory database, and
a literature review. The quantities of building components and their corresponding
environmental impact coefficient were recorded in a spreadsheet. The life cycle envi-
ronmental impact of the building in five scenarios was assessed by multiplying the
quantities of building components by the corresponding environmental impact coef-
ficient. In the life cycle interpretation phase, the building’s environmental impact was
expressed as global warming potential (reported in kg CO2 eq) based on the Traci 2.1
method [8].

3 Results and Discussion

The life cycle assessment results of the building in the five scenarios are presented
in Table 1.
The results showed that replacing virgin concrete with CO2 concrete in a building
could greatly reduce its carbon emissions. By increasing the proportion of CO2
Life Cycle Assessment of the Environmental Impacts of Virgin Concrete … 481

Table 1 Life cycle environmental impact of a building in five scenarios

Scenario 1: Scenario 2: Scenario 3: Scenario 44: Scenario 5:
0% CO2 30% CO2 50% CO2 75% CO2 100% CO2
concrete concrete + concrete + concrete + concrete
70% virgin 50% virgin 25% virgin
concrete concrete concrete
Global warming 275.830 (0) 271.33 268.36 264.632 260.848
(t CO2 eq/m2 ) (1.6%) (2.7%) (4.1%) (5.4%)
(declined ratio
compared with
scenario 1)

concrete in a building, its carbon emission decreases over its life cycle. As much as
5.4% of the CO2 eq/m2 can be reduced when 100% of the virgin concrete is replaced
by CO2 concrete.
This study evaluated the life cycle environmental performance of a residen-
tial building using CO2 concrete as a replacement for virgin concrete. The results
suggested that the application of CO2 concrete in the building sector will bring great
benefits in terms of environmental performance. However, the mechanical and dura-
bility qualities of CO2 concrete have been considered in this study. In future work,
more emphasis should be put on the mechanical and durability qualities of CO2
concrete for application in the building sector.

References

1. Marinkovic S, Dragas J, Ignjatovic I, Tosic N (2017) Environmental assessment of green

concretes for structural use. J Clean Prod 154:633–649. https://doi.org/10.1016/j.jclepro.2017.
04.015
2. Kisku N, Joshi H, Ansari M, Panda SK, Nayak S, Dutta SC (2017) A critical review and
assessment for usage of recycled aggregate as sustainable construction material. Constr Build
Mater 131:721–740. https://doi.org/10.1016/j.conbuildmat.2016.11.029
3. Tam VWY, Soomro M, Evangelista ACJ (2021) Quality improvement of recycled concrete
aggregate by removal of residual mortar: A comprehensive review of approaches adopted. Constr
Build Mater 288:123066. https://doi.org/10.1016/j.conbuildmat.2021.123066
4. Makul, N. 2020, ‘Cost-benefit analysis of the production of ready-mixed high-performance
concrete made with recycled concrete aggregate: A case study in Thailand’, Heliyon, vol. 6, no.
6, pp. e04135-e, DOI https://doi.org/10.1016/j.heliyon.2020.e04135.
5. Serres N, Braymand S, Feugeas F (2016) Environmental evaluation of concrete made from recy-
cled concrete aggregate implementing life cycle assessment. Journal of Building Engineering
5:24–33. https://doi.org/10.1016/j.jobe.2015.11.004
6. Xing W, Tam VWY, Le KN, Butera A, Hao JL, Wang J (2022) Effects of mix design and
functional unit on life cycle assessment of recycled aggregate concrete: evidence from CO2
concrete. Constr Build Mater 348. https://doi.org/10.1016/j.conbuildmat.2022.128712
482 M. Ma et al.

7. Islam H, Jollands M, Setunge S (2015) Life cycle assessment and life cycle cost implication
of residential buildings—A review. Renew Sustain Energy Rev 42:129–140. https://doi.org/10.
1016/j.rser.2014.10.006
8. Ryberg M, Vieira MDM, Zgola M, Bare J, Rosenbaum RK (2014) Updated US and Canadian
normalization factors for TRACI 2.1. Clean Technol Environ Policy 16(2):329–339. https://doi.
org/10.1007/s10098-013-0629-z

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Economic Impacts of Environmentally
Friendly Blocks: The Case of Nu-Rock
Blocks
V. W. Y. Tam, K. N. Le, I. M. C. S. Illankoon, C. N. N. Tran, D. Rahme,
and L. Liu

Abstract There are numerous industry byproducts that have negative environmental
impacts. Pond ash accumulated from coal power plants is one such byproduct that
creates major environmental and social issues, especially with regard to decom-
missioning a coal power plant. Using pond ash in the block manufacturing process
is a promising solution proposed by Nu-Rock. This research study evaluated the
economic impact throughout the life cycle of Nu-Rock blocks. Nu-Rock block
production use a technology called “Nu-creeting” in ash dams to prevent dust gener-
ation followed by the block manufacturing process using pond ash as a raw material.
Nu-Rock technology can process approximately 250,000 tonnes of ash per annum
and manufacture the equivalent of up to 330,000 tonnes volume of traditional building
materials. This manufacturing process already generates jobs and pays tax and royalty
fees to local governments, which is an added advantage. The total operating cost for
1 tonne of Nu-Rock blocks amounts to approximately AUD48, and the cost of a
Nu-Rock block is AUD1.50–2.40, which is within the range of common bricks.
Although there is a considerable initial cost to this process, it derives significant
economic benefits in terms of manufacturing blocks using industrial byproducts, job
creation and even tax revenue. Apart from these economic benefits, the Nu-Rock
block manufacturing process generates environmental benefits through the reuse of
pond ash from decommissioned coal power plants.

V. W. Y. Tam (B) · K. N. Le · L. Liu

School of Engineering Design and Built Environment, Western Sydney University, Sydney, NSW,
Australia
e-mail: vivianwytam@gmail.com
I. M. C. S. Illankoon
School of Architecture and Built Environment, University of Newcastle, Sydney, NSW, Australia
C. N. N. Tran
Ho Chi Minh City University of Technology, Ho Chi Minh, Vietnam
D. Rahme
Nu-Rock Australia Pty Ltd, Sydney, NSW, Australia

© The Author(s) 2023 483

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_49
484 V. W. Y. Tam et al.

Keywords Coal power plants · Economic issues · Industrial byproducts ·

Nu-Rock

1 Introduction

The manufacturing process of most conventional building materials poses various

environmental issues, and the brick and block manufacturing process is no exception.
To name a few of these issues, when manufacturing burnt clay bricks there is deple-
tion of fertile topsoil and emissions of CO2 to the atmosphere [1]. Conventional
bricks are produced from clay with high-temperature kiln firing or from ordinary
Portland cement concrete, and thus contain high embodied energy and have a large
carbon footprint [2]. To protect the clay resource and with the vision of developing
ecofriendly building materials, certain cities in China have banned the use of clay
bricks [3]. Brick manufacturing with higher resource efficiency and less carbon foot-
print is the priority. Industrial byproducts such as fly ash/pond ash are alternatives in
brick and block manufacturing process. Fly ash is a fine industrial waste byproduct
often found after coal combustion for power generation.
There are research studies of using byproducts in the brick manufacturing process.
Cicek and Tanriverdi [4] conducted research on using fly ash from thermal coal power
plants for brick manufacturing in Turkey. Several research studies have used various
percentages of fly ash in the process of manufacturing bricks; for example, Wang
et al. [5] introduced a technology to use up to 50% of fly ash in brick manufacturing.
Using pond ash as an alternative to sand is another viable alternative in the block
manufacturing process [6]. Similarly, there are many research studies on using fly
ash as a replacement material in both brick and block manufacturing [7, 8].
Although research has been conducted on ecofriendly brick/blocks using indus-
trial waste, the commercial production of such bricks is still very limited [2]. Slow
acceptance of waste materials-based bricks by industry and the public is one of the
challenges faced when commercializing these innovative products [2]. One of the
main reasons for this limited use in the commercial space is economic consider-
ations, because incorporating these byproducts is deemed to be costlier compared
with conventional materials.
Most research studies rarely discuss the economic aspects from initial investment,
construction cost, operational cost and recycling or reuse opportunities. Although
these studies used industrial waste, the environmental impacts were rarely quantified
using valid parameters. Therefore, this research study conducted a whole of life-cycle
cost analysis of Nu-Rock blocks.
Economic Impacts of Environmentally Friendly Blocks: The Case … 485

1.1 Nu-Rock Blocks: Properties and Manufacturing Process

In Australia, 54% of electricity is generated from coal combustion [9]. Thermal

coal power plants generate fly ash and bottom ash, and these are mixed with water,
resulting in a slurry that is carried to ash ponds [10]. When the Port Augusta power
plant closed, it created a huge pile of pond ash covering hectares of land, giving rise
to many environmental and social issues [11]. Nu-Rock technology uses this pond
ash to manufacture a range of building materials.
Nu-Rock Technology Pty. Ltd. is a privately owned company providing solutions
for environmental and economic issues related to the growing stockpiles of industrial
byproducts occurring worldwide. Nu-Rock produces a range of products, including
common or rendered bricks and blocks, masonry bricks, blocks, pipes, pavers, tiles,
sheeted products and any other shaped building products [12]. Our study focused
on the manufacturing of three types of Nu-Rock blocks, namely Mighty block (90
× 245 × 470 mm), 200 series (190 × 190 × 390 mm), and 100 series (90 × 190
× 390 mm). The Nu-Rock Technology has a Proprietary IP. The recycling process
usually converts 100% of the pond ash in abandoned power plants into building
blocks and brick products. The product has undergone rigorous testing including
(but not limited) AS3700D Flexural strength, AS1191 Acoustic fire and AS1530.4
Fire resistance.
The first step in the Nu-Rock block manufacturing process is called Nu-creeting,
which prevents dust generation and builds “modules” to process the pond ash. Nu-
Rock can process 250,000 tonnes of ash per annum and manufacture the equivalent
of up to 330,000 tonnes of traditional building materials using the pond ash as an
input [13]. Each module in Nu-Rock can process approximately 0.25 million tonnes
of ash per annum and each module can produce a specific type of product. Nu-Rock
Blocks require only 3% of the embodied energy required to produce traditional
concrete blocks and the Nu-Rock manufacturing process produces zero waste [14].
According to Tam et al. [11], there are great benefits in using this technology to
remediate ash ponds.

2 Methods

The primary objective of this research study was to conduct a life-cycle cost analysis
of the entire Nu-Rock block manufacturing process: block manufacturing, trans-
portation, operational stage of the building and the demolition. However, here we
present the results for initial stages up to construction.
486 V. W. Y. Tam et al.

2.1 Scope of the Analysis

The scope of the analysis was the entire process of blocks from cradle to cradle,
calculated in four stages:
Stage 1—Initial stage including setting up the plant.
Stage 2—Raw material extraction, manufacturing, transportation to construction
site.
Stage 3—Construction, operation, and maintenance.
Stage 4—Demolition/re-use phase.
The data are for stages 1 and 2 up to construction. The calculations included
sensible assumptions and limitations, which are given with the relevant calculations.

2.2 Life-Cycle Cost

Key parameters of a life-cycle analysis are determined to a large extent by its purpose
and objectives [15]. The main parameters of this project included costs, period of anal-
ysis, method of economic evaluation, extent of environmental input and sensitivity
analysis. Each of these parameters is discussed in detail.
Costs included those incurred by Nu-Rock Technology Pty Ltd and by the end-
users. The life-cycle cost calculation followed ISO15686-5:2017: Building and
construction assets—service life planning—Part 5: Life cycle costing” as a guide-
line. When a decision required including any cost or income in the analysis, the ISO
standard was followed. However, additional notes illustrate if any exception was
made. All the assumptions (if any) relevant to each calculation are given with it. This
project selected an analysis period of 60 years to reflect the anticipated total lifespan
of Nu-Rock Technology Pty Ltd.
All costs incurred within the life cycle must be captured and discounted into
present day values to calculate the life-cycle cost. Net present value (NPV) was the
economic evaluation method used for life-cycle cost calculation, as shown in Eq. 1
(adapted from Dell’Isola and Kirk [15])..

Σ
N
Rt
NPV (i, N ) = (1)
t=0
(1 + i)t

In Eq. 1, i denotes the discount rate; t denotes the time of cash flow; Rt denotes the
net cash flow, and N is the total number of periods. The discount rate considers the
time value of money and the associated risk. The return on equity (RoE) of Nu-Rock
Technology Pty Ltd was used as the discounting rate in the life-cycle cost calculation
to reflect the capital used by the company (RoE was provided by Nu-Rock).
Economic Impacts of Environmentally Friendly Blocks: The Case … 487

2.2.1 Life-Cycle Cost Analysis

The initial investment cost for Nu-Rock was AUD12,000,000, which included land
cost, specialized design costs, construction cost, cost for initial approvals, electricity
connection charges, water connection charge, cost of machinery, cost of specialized
equipment and other professional fees. The site establishment cost of AUD200,000
was not included in the investment cost. Therefore, the total initial investment was
AUD12,200,000 including site establishment.
Life- cycle cost calculations were based on the following sensible assumptions.
• Transportation costs included loading and unloading and bulk discounts for blocks
were not considered. Investment costs provided by Nu-Rock were for the Mt. Piper
plant.
• On-site labor costs include one site manager, four factory staff, including two for
lift drivers and accounts manager. Salary and associated costs were provided by
Nu-Rock.
• Repair and maintenance costs were 5% of the plant costs. The initial investment
cost of AUD12,000,000.00 was taken as the plant costs.
• Distribution cost included AUD30 per tonne as provided by Nu-Rock.
• Operational costs included other miscellaneous operations and energy costs. The
cost per kWh was considered to be 66 cents/kWh.
Table 1 summarizes the life-cycle costs for Nu-Rock during the raw material
extraction and manufacturing stage. The life-cycle cost for the 60-year period at
15% discount rate per tonne of blocks was AUD321.

Table 1 Life-cycle costs for Nu-Rock during the raw material extraction and manufacturing stage
Description Cost (AUD)
Factory costs
Site manager 174,750
Outsourced factory operations 360,000
Accounts Manager/plant administration 79,220
Miscellaneous costs 300,000
Selling and distribution costs
Advertising 220,000
Distribution costs 167,893
Other costs
Repairs and maintenance cost p.a 600,000
Other operational costs 40,032
Total cost per annum 1,941,895
Total cost per tonne of blocks 48
Life-cycle costs for 60 years 15% discount rate per tonne of blocks 321
488 V. W. Y. Tam et al.

Table 2 Construction cost for Nu-Rock block types

Type of block
Mighty block 200 series 100 series
Size 90 × 245 × 190 × 190 × 90 × 190 ×
470 mm 390 mm 390 mm
No. of blocks per m2 8 12.5 12.5
Cost per block (AUD) 2.40 1.80 1.50
Cost of blocks per m2 (AUD) 19.20 22.50 18.75
Other costs including mortar, tools etc. Labor 4.20/ Labor 4.20/block Labor 3.20/
(AUD) block block
Total cost of construction (per m2 ) 52.80 75.00 58.50
Total cost of construction (per block) 6.6 6 4.68

The next phase was the construction, operations and maintenance stage. This cost
calculation included a wall construction using Nu-Rock blocks. “Wall” was assumed
as a face brick without any finishing. The size and further details of the three types
of Nu-Rock block are given in Table 2.
Stages 3 and 4 are not discussed here but are expected to derive savings by using
Nu-Rock.
The selling price of a Nu-Rock block varies between AUD 1.50 and 2.40 which is
within the range of common bricks. When setting up the factory Nu-Rock incurred
an initial cost of AUD12,200,000. The cost attributed during material extraction was
AUD48 per tonne of Nu-Roc blocks (refer to Table 1). During this phase Nu-Rock
blocks absorb industrial byproducts such as waste from coal-fired power stations,
steel mills, non-ferrous smelters and alumina smelters. This is a non-quantifiable
benefit of using Nu-Rock blocks. According to Tam et al., the Nu-Rock manufac-
turing plant provides almost AUD20 million worth of jobs [11], which is a significant
social benefit.

3 Conclusions

The life-cycle cost impacts of Nu-Rock blocks for the first two stages of the manu-
facturing process, starting from setting up the factory up to the actual construction,
were calculated. Nu-Rock technology uses pond ash to manufacture blocks. Ash
ponds in discontinued power plants pose serious environmental and social threats.
According to our life-cycle calculations, the cost of setting up the factory was ≈
AUD12.2 million. The life-cycle cost for the 60-year period at 15% discount rate
per tonne of blocks was AUD321. However, it is interesting to note that the cost
of Nu-Rock blocks was within the range of conventional bricks. Although there are
many research studies of environmentally friendly materials using industrial byprod-
ucts, the commercialization of these products is very slow. It is necessary to conduct
Economic Impacts of Environmentally Friendly Blocks: The Case … 489

similar economic and environmental impact analyses of these products to ensure end-
users of their importance. Although this research study was limited to the first two
stages, we are planning to extend the study to include stage 3 and 4 in the life-cycle
cost calculation.

References

1. Waheed A, Azam R, Riaz MR, Zawam M (2021) Mechanical and durability properties of
fly-ash cement sand composite bricks: an alternative to conventional burnt clay bricks. Innov
Infrastruct Solut 7(1):24. https://doi.org/10.1007/s41062-021-00630-w
2. Zhang L (2013) Production of bricks from waste materials—a review. Constr Build Mater
47:643–655. https://doi.org/10.1016/j.conbuildmat.2013.05.043
3. Chen Y, Zhang Y, Chen T, Zhao Y, Bao S (2011) Preparation of eco-friendly construction bricks
from hematite tailings. Constr Build Mater 25(4):2107–2111. https://doi.org/10.1016/j.conbui
ldmat.2010.11.025
4. Cicek T, Tanrıverdi M (2007) Lime based steam autoclaved fly ash bricks. Constr Build Mater
21(6):1295–1300. https://doi.org/10.1016/j.conbuildmat.2006.01.005
5. Wang L, Sun H, Sun Z, Ma E (2016) New technology and application of brick making with
coal fly ash. J Mater Cycles Waste Manage 18(4):763–770. https://doi.org/10.1007/s10163-
015-0368-9
6. Wahab AA, Arshad MF, Ahmad Z, Ridzuan ARM, Ibrahim MHW (2018) Potential of bottom
ash as sand replacement material to produce sand cement brick. Int J Integr Eng 10(8)
7. Naganathan S, Mohamed AYO, Mustapha KN (2015) Performance of bricks made using fly
ash and bottom ash. Constr Build Mater 96:576–580. https://doi.org/10.1016/j.conbuildmat.
2015.08.068
8. Sahu S, Ravi Teja PR, Sarkar P, Davis R (2019) Variability in the compressive strength of fly
ash bricks. J Mater Civ Eng 31(2):06018024
9. Department of Industry Science Energy and Resources (2021) Australian electricity genera-
tion–fuel mix. https://www.energy.gov.au/data/australian-electricity-generation-fuel-mix
10. Ghosh P, Goel S (2017) Leaching behaviour of pond ash. In: Goel S (ed) Advances in solid
and hazardous waste management. Springer International Publishing, Cham, pp 171–204
11. Tam VWY, Rahme D, Illankoon IMCS, Le KN, Yu J (2021) Effective remediation strategies for
ash dam sites in coal power plants. Proc Instit Civil Eng Eng Sustain 174(2):94–105. https://doi.
org/10.1680/jensu.20.0005212Nu-Rock. Technology Pty Ltd. (2017). Nu-Rock technology.
http://www.nu-rock.com/
12. Nu-Rock Technology Pty Ltd. (2017) Nu-Rock technology. http://www.nu-rock.com/
13. Allen M (2018) Nu-Rock: the product which rose from the ashes. Western Advocate. https://
www.westernadvocate.com.au/story/5658840/product-rises-from-ashes/
14. Nu-Rock (2020) Nu Rock–greener, smarter, stronger. https://nu-rock.com/
15. Davis and Langdon (2007) Life cycle costing (LCC) as a contribution to sustainable
construction: a common methodology
490 V. W. Y. Tam et al.

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Life Cycle Assessment (LCA) of Recycled
Concrete Incorporating Recycled
Aggregate and Nanomaterials

W. Xing, V. Tam, K. Le, J. L. Hao, J. Wang, and P. Yang

1 Extended Abstract

Nanotechnology is widely used in construction to advance high performance and

durability materials, as represented by concrete, in which nanomaterials such as
nanosilica (SiO2 ), nano-alumina (Al2 O3 ), and nanotitanium oxide (TiO2 ) are used
to replace cement as a partial binder. Suitable amounts of adding these nanoma-
terials are beneficial with regard to the mechanical, economic, and environmental
aspects of concrete. For example, the pores in the cement matrix can be filled by
nano-SiO2 , which has been demonstrated to enhance by ≈10% the comprehen-
sive strength and by 25% flexural strength of the concrete [1]. Nano-TiO2 concrete
shows outstanding effective self-cleaning because TiO2 nanoparticles can trigger a
photocatalytic degradation of gas emissions (e.g. NOx, CO and VOCs), reducing the
environmental impact of such concretes in a promising way [2].
To seek better performance of the concrete with the least amount of virgin material
and address the global issues of construction and demolition waste and emissions,
some researchers have introduced both recycled aggregate and nanomaterials into
concrete, but environmental evaluations of such concrete products are scarce. There-
fore, we assessed a broad range of environmental impacts of recycled aggregate
concrete in a comprehensive literature review of sustainable concretes containing
recycled aggregate and nanomaterials to establish an experimental database. Only
concretes with recycled aggregate and nano-SiO2 were included after the review

W. Xing · V. Tam (B) · K. Le · J. Wang · P. Yang

School of Engineering, Design and Built Environment, Western Sydney University, Sydney,
NSW, Australia
e-mail: vivianwytam@gmail.com
J. L. Hao
Department of Civil Engineering, Xi’an Jiaotong-Liverpool University, Sydney, China

© The Author(s) 2023 491

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_50
492 W. Xing et al.

Fig. 1 Carbon emissions and energy consumption of concretes

process. These concrete products were evaluated by life cycle assessment (LCA),
considering both volume and compressive strength.
The LCA model fitting sustainable concretes incorporating recycled aggregate
and nanomaterials was modified from the framework of Xing et al. [3], based on
the normal conditions and current technology applied to the Australian concrete
industry. The functional unit was 1 m3 of ready-mix concrete, considering its 28-day
compressive strength, and the system boundary was from cradle to gate. The LCA
methodology was CML 2002, but focused primarily on carbon emissions and energy
consumption.
The literature review process resulted in 92 mix designs containing recycled aggre-
gate and nano-SiO2 and their carbon emissions and energy consumption were quan-
tified by the LCA model, as shown in Fig. 1. Generally speaking, the incorporation
ratio of nano-SiO2 was limited to 15% whereas virgin aggregate could be fully
replaced by the recycled aggregate. The carbon emission and energy consumption of
the concretes examined were 91.6–607.1 kg CO2 eq., and 611.6–2453.6 MJ, respec-
tively. With the development of compressive strength, both the carbon emissions and
energy consumption of the concrete increase. When other supplementary cementi-
tious materials are simultaneously used in the concrete mix, the concrete products
perform much better.
Compared with the effect of nano-SiO2 on the strength and environmental
impact of the concrete, substituting recycled aggregate was less effective. We there-
fore suggest introducing nanomaterials into recycled aggregate concrete for higher
potential regarding the sustainability of the concrete.
Our study quantified the effects of nanomaterials and recycled aggregate to the
overall environmental impacts of designed concrete products and estimated the
ecological performance based on the mix design and compressive strength.
Life Cycle Assessment (LCA) of Recycled Concrete Incorporating … 493

References

1. Qing Y, Zenan Z, Deyu K, Rongshen C (2007) Influence of nano-SiO2 addition on properties

of hardened cement paste as compared with silica fume. Constr Build Mater 21(3):539–545
2. Zhu W, Bartos PJ, Porro A (2004) Application of nanotechnology in construction. Mater Struct
37(9):649–658
3. Xing W, Tam VW, Le KN, Butera A, Hao JL, Wang J (2022) Effects of mix design and functional
unit on life cycle assessment of recycled aggregate concrete: Evidence from CO2 concrete.
Constr Build Mater 348:128712

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Analysis of the Compressive Strength
of CO2 Concrete While Eliminating
Overshadowing Concrete Variables

V. W. Y. Tam, A. Butera, K. N. Le, and L. Liu

1 Extended Abstract

Although other technologies for the improvement of recycled aggregate, such as

removal of adhered cement, improvement of aggregate by methods including nano-
materials, sequential mixing and precoating, the ideology of sequestration of CO2 into
cementitious materials provides great potential for both the enhancement and utilisa-
tion of recycled concrete materials and sequestration of CO2 . However, a great deal of
study has focused on the complete carbonation of recycled aggregate without consid-
eration of a practical timeframe for use with ready-mix concrete. Furthermore, the
carbonation of recycled aggregate concrete does provide strength increases that can
be overshadowed by other concrete mix design variables such as the water to cement
(w/c) ratio and aggregate replacement percentage. Consequently, here, we focus on
practical carbonation timeframes with a maximum of 2 h where the carbonation reac-
tion is most prevalent, and keep other major concrete variables consistent to ensure
the effect of carbonation can clearly be observed. The injection of CO2 into recy-
cled aggregate sequesters the undesirable gas while reducing the porosity and water
absorption of the recycled aggregate. When CO2 is injected into recycled aggregate, a
chemical reaction occurs, converting the CO2 into a mineral, calcium carbonate. The
converted calcium carbonate fills air voids within the aggregate and consequently
improves its mechanical performance. The carbonated aggregate concrete is also
known as CO2 concrete. Many researchers have investigated this mechanism, but on
many occasions the true effect of carbonation variables (e.g., pressure and duration)
on the compressive strength of the concrete cannot be observed due to other concrete
variables, such as the w/c ratio and aggregate replacement percentage, which can

V. W. Y. Tam (B) · A. Butera · K. N. Le · L. Liu

School of Engineering, Design and Built Environment, Western Sydney University, Penrith,
NSW, Australia
e-mail: vivianwytam@gmail.com

© The Author(s) 2023 495

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_51
496 V. W. Y. Tam et al.

overshadow the performance compressive increase provided by the carbonated recy-

cled aggregate. In our study, we provided a constant w/c ratio and separate recycled
aggregate percentages so that the effect of the carbonation variables could be truly
observed. The longer carbonation duration of 120 min surpasses the lesser durations
of 60 and 30 min. Furthermore, generally the lower carbonation pressure of 25 kPa
provided greater results over the 75 and 200 kPa treated samples. Our results show
that practical carbonation variables can help to increase the mechanical properties
of recycled aggregate concrete. Furthermore, without other mix design variables
a longer carbonation duration is of great importance to improving the properties of
recycled aggregate concrete. The carbonation of recycled aggregate can help to make
the greener product mainstream.

Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Harmonic Vibration of Inclined Porous
Nanocomposite Beams

D. Chen and L. Zhang

Abstract This work investigated the linear harmonic vibration responses of inclined
beams featured by closed-cell porous geometries where the bulk matrix materials
were reinforced by graphene platelets as nanofillers. Graded and uniform porosity
distributions combined with different nanofiller dispersion patterns were applied in
the establishment of the constitutive relations, in order to identify their effects on
beam behavior under various harmonic loading conditions. The inclined beam model
comprised of multiple layers and its displacement field was constructed using Timo-
shenko theory. Forced vibration analysis was conducted to predict the time histories
of mid-span deflections, considering varying geometrical and material characteriza-
tions. The findings may provide insights into the development of advanced inclined
nanocomposite structural components under periodic excitations.

Keywords Functionally graded porosity · Graphene platelets · Harmonic

vibration · Inclined beams

1 Introduction

The inclined beam problem has attracted many researchers and engineers due to its
application potential in various fields (bridges and skytrain rail [1], for instance).
The axial force induced by the inclined angle leads to a different deformation pattern
compared with horizontal beams [2]. Meanwhile, in order to achieve enhanced perfor-
mance, inclined structures with novel material compositions have been proposed,
such as functionally graded (FG) inclined pipes [3] and inclined FG sandwich beams
[4]. The recent development in this field using graded distributions of porosity and
graphene platelets (GPLs) is demonstrated to be promising. Chen et al. [5] pointed

D. Chen (B) · L. Zhang

Department of Infrastructure Engineering, University of Melbourne, VIC, Australia
e-mail: da.chen1@unimelb.edu.au
D. Chen
School of Civil Engineering, University of Queensland, St. Lucia, QLD, Australia

© The Author(s) 2023 497

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_52
498 D. Chen and L. Zhang

out that strategic arrangements of internal pore size/density and GPL weight frac-
tions significantly benefit the low-velocity impact properties of inclined beams under
various impulses. The combination of FG porosities and graphene nanofillers can
boost the mechanical performance of lightweight structural components [6, 7].
In this work, we aimed to further reveal the behavior of this novel porous nanocom-
posite by focusing on the responses of corresponding inclined beams under harmonic
excitations, which represent steady wind, unbalanced rotating machine force, or
vehicle loadings, etc. The theoretical formulations were first briefed and validated,
then the mid-span deflections of fully clamped beams with changing inclined angles,
porosity coefficients, GPL weight fractions and slenderness ratios were examined,
considering the typical harmonic force sitting on the top mid-span surface.

2 Formulation Briefing

Figure 1 is a schematic illustration of the examined FG beam, of which the elastic

properties varied along the height direction with E(z) = E ∗ [1 − e0 α(z)] for
Young’s modulus, G(z) = E(z)/[2(1 + v(z))] for shear modulus, and ρ(z) =
ρ ∗ [1 − e∗ α(z)] for mass density, where Young’s modulus of non-porous nanocom-
posites E ∗ = φ(E Al , ΔG P L ) were determined using the Halpin–Tsai microme-
chanics model, e0 reads the porosity coefficient related to internal pore size/density,
α(z) = cos(π z/ h) and α(z) = constant correspond to the graded and uniform
porosities, respectively, and ρ(z) and G(z) root in the closed-cell morphologies. The
GPL weight fraction was ΔG P L = ζ1 [1 − cos(π z/ h)] for non-uniform dispersions
and ΔG P L = ζ2 for the uniform one, while ζ1 and ζ2 were related to material distri-
butions. The mass density and Poisson’s ratio of non-porous nanocomposites were
computed via the rule of the mixture with the corresponding values of the matrix
materials (ρ Al , v Al ) and GPLs (ρG P L , vG P L ). Note that the FG variations of porosity
and GPL dispersion were both symmetric about the mid-height plane. Based on
the established constitutive relations, the beam governing equation systems were
derived within the framework of Timoshenko theory, and solved with the aid of the
Ritz method and Newmark method. Consequently, the time history of beam mid-
span deflections was estimated, in which the initial deformation status caused by
self-weight was also embedded [5].

3 Results and Discussion

Based on aluminum composites, the parameters adopted in this study included E Al =

70 GPa, ρ Al = 2700 kg/m3 , v Al = 0.34, wG P L = 1.5 µm, l G P L = 2.5 µm, tG P L =
1.5 nm, E G P L = 1.01 TPa, ρG P L = 1062.5 kg/m3 , vG P L = 0.186. The beam section
dimension was 0.1 × 0.1 m and the layer number was 14 [8]. The time length (2 s) was
Harmonic Vibration of Inclined Porous Nanocomposite Beams 499

200sin 2 t Loading A
Inclined Harmonic load (N) 400sin 3 t Loading B
angle 800sin 4πt Loading C
Perfectly bonded layers
Applied on the mid-span top surface

Fully clamped supports

at both ends

Symmetrically graded along height

Graded/uniform porosities
Non-uniform/uniform graphene
dispersions

Fig. 1 Inclined functionally graded beam

the same for all calculation cases to obtain the maximum absolute value of the mid-
span deflections. The beam boundary condition was taken as the clamped–clamped
end supports. The present results were validated by being compared with ANSYS
simulations (see Fig. 2). Figure 3 shows the mid-span deflection time histories of
the examined beams under three harmonic loading scenarios. Results suggested that
dispersing both internal pores and GPLs non-uniformly enhanced the inclined beam’s
stiffness due to reduced maximum deflections (~15%, ~18%, ~30%, as marked in
Fig. 3a). It is also noticeable that the minimum deflections for the beams remained
almost the same in loading case A but differed in loading cases B and C, because of
the influence of gravity and the increased peak force in B and C.
For the purpose of simplification, the assessment given below is limited to
responses subjected to harmonic loading case A. Figure 4a compares the harmonic
vibration deflections of beams inclined at various angles. It is obvious that a larger
inclined angle resulted in smaller deflections with lower harmonic force components
along the height direction. The variations of porosity coefficient and GPL weight
fraction are displayed in Figs. 4b, c. We can see that when the porosity coefficient
increased from 0.2 to 0.6, the maximum deflection increased by ~10%. Meanwhile,
an improved level of graphene weight fraction from 0.2% to 1.0% significantly

Fig. 2 Mid-span deflection Present ANSYS

time history of an inclined 0.04
Mid-span deflection (mm)

(45°) beam with graded

porosity and non-uniform
GPL dispersion under 0.02
loading case C (e0 = 0.5,
ΔG P L = 1.0 wt.%, L / h =
20) 0.00

-0.02

0.0 0.5 1.0 1.5 2.0

Time (s)
500 D. Chen and L. Zhang

0.030 0.05 0.08 Graded Porosity & Non-uniform GPL

Graded Porosity & Non-uniform GPL Graded Porosity & Non-uniform GPL
Graded Porosity & Uniform GPL Graded Porosity & Uniform GPL
Graded Porosity & Uniform GPL Uniform Porosity & Non-uniform GPL

Mid-span deflection (mm)

0.025 Uniform Porosity & Non-uniform GPL
Mid-span deflection (mm)

0.04 0.06

Mid-span deflection (mm)

Uniform Porosity & Non-uniform GPL Uniform Porosity & Uniform GPL
Uniform Porosity & Uniform GPL Reduced by Uniform Porosity & Uniform GPL
0.020 ~15% 0.03 0.04
~18%
0.015 0.02 0.02
~30%
0.010 0.01
0.00

0.005 0.00
-0.02
0.000 -0.01
0.0 0.5 1.0 1.5 2.0 0.0 0.5 1.0 1.5 2.0 0.0 0.5 1.0 1.5 2.0
Time (s) Time (s) Time (s)

(a) Loading A (b) Loading B (c) Loading C

Fig. 3 a–c Mid-span deflection time histories of inclined functionally graded beams under
harmonic loadings (inclined angle 45°, e0 = 0.5, ΔG P L = 1.0 wt.%, L / h = 20)

stiffened the inclined beam of which the maximum mid-span deflection evidently
decreased (~29%).
Table 1 details the influence of the slenderness ratio and inclined angle on the
maximum mid-span deflections of inclined beams and still considers loading case
A. Compared with two extreme cases (60° for L/h = 20; 0° for L/h = 40), a wide
gap of 0.2476 mm was identified between their deflections. Maximum mid-span deflection (mm)

0.025
Maximum mid-span deflection (mm)

o 0.0250 0.04
0 e0= 0.2
o
15 e0= 0.3
Mid-span deflection (mm)

0.020 o
0.0225
30 e0= 0.4
o
45 0.0200 e0= 0.5 0.03
0.015 60
o
e0= 0.6
0.0175
0.010
0.0150 0.02

0.005
0.0125

0.000 0.0100 0.01

0.0 0.5 1.0 1.5 2.0 0 10 20 30 40 50 60 0 10 20 30 40 50 60
Time (s) Inclined angle ( )
o
Inclined angle ( o )

(a) Inclined angle (b) Porosity coefficient (c) Graphene weight fraction

Fig. 4 a–c Mid-span deflections of inclined functionally graded beams under harmonic loading A
(graded porosities & non-uniform graphene dispersions, e0 = 0.5 and ΔG P L = 1.0 wt.% for a,
ΔG P L = 1.0 wt.% for b, e0 = 0.5 for c, L/h = 20)

Table 1 Maximum mid-span deflections (mm) of inclined beams under harmonic loading A
(graded porosity & non-uniform graphene dispersion, e0 = 0.5, ΔG P L = 1.0 wt.%)
Inclined angle L/h = 20 L/h = 30 L/h = 40
0° 0.0219 0.0911 0.2586
20° 0.0206 0.0856 0.2430
40° 0.0168 0.0698 0.1981
60° 0.0110 0.0455 0.1293
Harmonic Vibration of Inclined Porous Nanocomposite Beams 501

4 Conclusions

We have discussed the influence of inclined angle, porosity, and graphene on beam
harmonic vibration responses. We conclude that all three parameters are closely
related to beam deflections under different harmonic loadings. The examined porous
nanocomposites may be used to develop novel inclined structural components with
reduced weight and enhanced stiffness subjected to periodic excitations.

Acknowledgements This research was supported by Australian Research Council (Chen: ARC
DECRA DE220100876).

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3. Reddy RS, Panda S, Gupta A (2020) Nonlinear dynamics of an inclined FG pipe conveying
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beam with bidirectional functionally graded face sheets under a moving mass. Eur J Mech A
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graded graphene reinforced porous nanocomposite beams. Compos Sci Technol 142:235–245
7. Yang J, Chen D, Kitipornchai S (2018) Buckling and free vibration analyses of functionally
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the copyright holder.
Influence of Carbon Nanotubes
on the Fracture Surface Characteristics
of Cementitious Composites Under
the Brazilian Split Test

Y. Gao, J. Xiang, Z. Yu, G. Han, and H. Jing

Abstract To better analyze the reinforcing mechanisms of carbon nanotubes (CNTs)

in cementitious composites, the micromorphological characteristics of the fracture
surface of cement-based specimens under the Brazilian split test were investigated.
The results demonstrated that the addition of CNTs promoted nucleation and pore
filling effects that optimized the pore structure and inhibited the development of
microcracks in the cement matrix during loading. Due to the ultra-high specific
surface area and bridging effects of CNTs, the formation of hydration products
was promoted, internal microcracks and micropores were effectively reduced, and
the regularity and integrity of the samples also improved. Therefore, under tensile
loading, CNT-reinforced cementitious composites absorbed more energy, resulting
in more complex stress paths and rougher fracture surfaces.

Keywords Cement · Fracture characteristics · Nano-reinforcement · Reinforcing

mechanisms

1 Introduction

In the past decades, carbon nanomaterials, including carbon nanotubes (CNTs),

graphene, graphene oxide and reduced graphene oxide, have attracted great atten-
tion as additives to reinforce the physical and mechanical properties of cementitious
composites [1]. Due to their ultra-high specific surface area, superior mechanical

Y. Gao (B)
School of Transportation and Civil Engineering, Nantong University, Nantong, China
e-mail: Y.Gao@ntu.edu.cn
J. Xiang · G. Han
Key Laboratory of Rock Mechanics and Geohazards of Zhejiang Province, Shaoxing University,
Shaoxing, China
Z. Yu · H. Jing
State Key Laboratory for Geomechanics and Deep Underground Engineering, China University of
Mining and Technology, Xuzhou, China

© The Author(s) 2023 503

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_53
504 Y. Gao et al.

properties, and the densification effect on the microstructure of the cement, CNTs
have been proved to be capable of enhancing the tensile strength [2] and dura-
bility [3] of cement-based materials with a minimal mixing ratio (0.01–0.05% by
weight of cement) [4]. Moreover, the latest research suggests that carbon nanomate-
rial reinforcement is a cost-effective strategy for reducing the environmental impact
of cement usage and maintenance, providing an effective way to reduce CO2 emis-
sions and energy consumption in the cement production and construction industries
[5]. Hence, it is significant to carry out research on the development and applica-
tion of carbon nanomaterial-modified cementitious composites and investigate the
corresponding reinforcing mechanisms.
The micromorphological characteristics of the fracture surface of cementitious
composites are closely related to the micromechanical damage evolution process
of hardened cement-based materials. The micro-geometric features of the fracture
surface record the irrecoverable deformation of the cementitious sample when it was
broken, as well as micro-information from the initiation, propagation, penetration
and nucleation of cracks to the final fracture and instability of the overall cement
structure [6]. In scientific research and engineering practice, the relevant information
recorded during the fracture damage process of cement materials can be obtained
through the study of the micromorphology of the typical fracture surface of cement
samples. Afterwards, mathematical statistics, induction and analysis of geometric
features can be applied and then the mechanical mechanism of fracture could be
reversed [7].
In the present study, we analyzed the reinforcing mechanisms of CNTs in cemen-
titious composites by investigating the micromorphological features of the fracture
surface of cement-based specimens under the Brazilian split test. First, we used 3D
scanning technology to reconstruct the micromorphology of the cement-based spec-
imens after the Brazil split test. Next, the micro-roughness characteristic parameters
of the specimens were calculated and quantitatively analyzed. Considering the influ-
ential mechanism of micro-roughness on material fracture, a micro-dilatancy micro-
element model of the fracture surface of the cement samples was constructed. The
micro-dilation angle and the normal dilatation deformation of the CNT composite
cement-based material after the Brazilian split test were calculated using the micro-
roughness characteristic parameters. Finally, the influence of the CNTs on the fracture
surfaces characteristics of cementitious composites under the Brazilian split test was
discussed.

2 Methods

Ordinary Portland cement (PO. 32.5) was applied as the binder material and fly ash
(FA), with a density of 2.4 g/cm3 , was used as a partial replacement for cement powder
in the cementitious composites. Multi-wall carbon nanotubes (MWCNTs), fabricated
via chemical vapor deposition, were selected as the nano-reinforcing material to
enhance the composites. Polycarboxylic acid water reducer (PC), an ionic surfactant,
Influence of Carbon Nanotubes on the Fracture Surface Characteristics … 505

non-toxic and harmless, containing both hydrophobic and hydrophilic groups in the
molecules, was used to improve the dispersion of CNTs in the suspensions. The
CNTs dispersion and cement-based material pouring process were consistent with
our previous report [8]. Three groups, Ref-group, FA-group and CNT-group, were
marked in this study. A 0.4 water-to-cement ratio was applied in all three groups.
For the FA-group, 20 wt% FA was used as a substitute to reduce cement usage and
enhance the workability of the paste to generate a cost-effective material. For the
CNT-group, 0.08 wt% CNTs was mixed into the cement–FA-based slurry to further
optimize the pore structure of the hardened matrix and enhance the mechanical
properties of the cementitious composite.
Then, the 28-day-old standard cured disc samples with a diameter of 50 mm and
a thickness of 25 mm were selected for the Brazilian split tests. The loading rate
was 0.10 mm/min. In order to reduce the influence of errors in the test and improve
accuracy, three samples were selected for each group for testing. Afterwards, a VR
3000 3D contour scanner with high precision was used to scan the fracture surface
of the tested specimens. The manual stitching mode was selected for the scanning
process. The maximum scanning size was 90 × 160 mm, and the highest resolution
was 1 μm.

3 Results

3.1 Tensile Properties of the Cementitious Composites

The mean and standard deviation of the tensile strength of the cementitious compos-
ites are shown in Fig. 1. The tensile strength of the plain cement specimen was
≈3.6 MPa. After mixing in FA, the tensile strength of the cementitious specimens
was significantly reduced to 3.17 MPa, ≈12% lower than that of the Ref-group.
The deterioration in tensile strength was mainly due to the decreased cement powder
content, which led to a decrease in the hydration degree of the composite. The hydra-
tion in the sample was uneven and incomplete, and there were more microcracks and
micropores in the specimens. In contrast, the mean tensile strength of the CNT-
group specimen was significantly improved to 3.77 MPa with only mixing 0.08wt%
CNTs into cementitious composites, which was 4.6% and 18.9% higher than the
Ref-group and FA-group specimens, respectively. The reinforcing rate of the tensile
properties of cement-based materials after mixing CNTs was highly consistent with
previous studies [9]. According to our previous report [8], by their nucleation and
crack bridging effects, CNTs effectively enhance the microstructure of cementitious
composites, thereby strengthening their mechanical properties.
506 Y. Gao et al.

Fig. 1 Tensile strength of 5

different groups of
cementitious composite
materials 4

c MPa
2

0
Ref-group FA-group CNT-group
Sample

3.2 Analysis of the Mesoscopic Asperities of the Fracture

Surface

The fracture method describes the surface of the fracture through three aspects: shape,
undulation and roughness [10]. Typically, macro-geometric features are described
by shape and relief, and micro-geometric features are described by roughness. The
micro-roughness is the degree of unevenness of the asperities attached to the macro-
relief surface in a small-scale range, as shown in Fig. 2a. Previous studies have
proposed several qualitative and quantitative methods for evaluating the microscopic
roughness of rock fracture surfaces [7]. For example, Barton et al. [11] proposed
evaluating the roughness of fracture surfaces by visual inspection, comparing and
referencing 10 standard joint roughness coefficient curves. Based on their theory,
many researchers have further explored and developed related statistical and fractal
methods to more accurately characterize the roughness of fracture surfaces [12, 13].
In this study, the statistical method for the microscopic roughness of the fracture
surface of cement-based materials was based on the method proposed by Belem et al.
[14]. As demonstrated in Fig. 2b, the 3D fracture point cloud data were discretized
into equally spaced points and gridded into a series of fine-scale planes. For each
microscopic plane, the local inclination angle (αij) is defined as its angle with the
horizontal plane, which is the angle between the normal vector n and z-axis. The
αij and the height of each minutiae plane are counted to obtain the maximum,
minimum, mean and standard deviation. For microstructure characterization, a 3D
contour scanner was employed to image the typical fracture surface of the cementi-
tious specimen after tensile damage to obtain the fracture surface point cloud data.
The fracture surface point cloud image was cropped by Geomagic Wrap software,
with a cropping size of 40 mm in length and 20 mm in width. The cropped 3D mesh
data were exported by PolyWorks software and 3D reconstruction was performed
using MATLAB. Representative 3D reconstructed images are exhibited in Fig. 4a–c.
Influence of Carbon Nanotubes on the Fracture Surface Characteristics … 507

Fig. 2 Schematic of microscopic fracture surface: a microscopic roughness of fracture; b slope

angle of reference surface in different dimensions

The fracture surface undulation height distribution for the three groups of cementi-
tious specimens after the Brazilian split tests are shown in Fig. 3. It can be seen that the
standard deviations of the undulation heights of the Ref-group, FA-group and CNT-
group specimens are 0.83, 1.01 and 0.76 mm, respectively. The standard deviation
of the undulation height of the FA-group specimens’ fracture surface was ≈21.7%
larger than that of the Ref-group, because the incorporation of FA decreased the
hydration reaction rate and reduced the hydration products, resulting in more micro-
cracks and micropores in the specimens. The fundamental cause of the failure of
the sample was the original microcracks and micropores inside the sample gradually
extending outward and deeper, causing the damage. Hence, the more microcracks
and micropores inside the FA-group specimens caused a more complex and tortuous
extension of the primary fracture. The undulation of the fracture surface was larger,
which eventually led to a more significant standard deviation of the undulation height
of the fracture surface of the FA-group than that of the Ref-group. With the addition
of CNTs, the micropores inside the cementitious composites were optimized due
to the nucleation and pore infilling effects [8], and porosity was also reduced. As a
result, the extension path of the primary fracture developed from fewer micropores
and microcracks became single and smooth, with fewer undulations on the fracture
surface. Finally, the standard deviation of the undulation height of the fracture surface
of the CNT-group specimens was smaller than that of the Ref- and FA-groups.
The measured results of the microscopic relief angle of the fracture surface are
presented in Fig. 4. The standard deviations of the microscopic relief angles of
the CNT-group and FA-group specimens were more extensive than those in the Ref-
group, indicating that the plane slope changed more in space, and the fracture surface
was correspondingly rougher. Cementitious composites continuously absorb energy
during the loading process. For the FA-group specimens, after FA replaced part of
the cement powder, the internal hydration reaction of the slurry was not uniform,
resulting in nonuniform energy distribution inside the loading samples. The crack
propagation path was more complicated when the unevenly distributed energy began
508 Y. Gao et al.

Fig. 3 Three-dimensional reconstruction images of fracture surface of typical cementitious

materials under tensile damage: a Ref-group, b FA-group and c CNT-group
Influence of Carbon Nanotubes on the Fracture Surface Characteristics … 509

Fig. 4 Microscopic undulation height distribution of the fracture surface of cement-based samples
under tensile failure condition: a Ref-group, b FA-group and c CNT-group

to be released into the stress concentration area at the tip of the microcrack. For CNT-
group specimens, with the incorporation of CNTs, the nucleation effects promoted
the hydration reaction, contributing to more and denser hydration products. The
integrity of the specimens became higher and could absorb more energy during
tensile loading. When the energy accumulated to a specific value, a large amount
of energy was released instantaneously, and the stress was concentrated around the
fracture surface, making the crack propagation path more complicated. In addition,
the bridging and pull-out effects [9] of the CNTs effectively inhibited the development
of cracks. More secondary microcracks were generated inside the sample, and the
crack propagation path became more complicated, resulting in higher fracture surface
roughness (Fig. 5).
510 Y. Gao et al.

Fig. 5 Microscopic slope angle distribution of the fracture surface of cement-based samples under
tensile failure condition: a Ref-group, b FA-group and c CNT-group

4 Discussion and Conclusions

In this work, CNTs were used as the additive in cementitious composites to assist FA
in reducing cement usage and generating cost-effective, environmentally friendly,
high-workability cement-based materials. The tensile properties of three groups of
specimens were tested by the Brazilian split test. The fracture surfaces after tensile
loading were scanned by a high-precision 3D scanning system. Based on fracture
theory, the micro-enhancing mechanism of CNTs on cementitious composites was
revealed.
Compared with plain cement, replacing the same cement mass in the slurry with
20 wt% FA resulted in a 12% deterioration in the tensile strength of the cement-
based material. Nevertheless, adding only 0.08 wt% of CNTs into the cement–FA
hybrid composites significantly increased the tensile strength by 18.9%. Compared
with plain cement-based materials, the tensile strength of the hardened matrix was
enhanced by 4.6%. This finding indicated that 0.08 wt% CNTs combined with 20
wt% FA could be a good substitute for cement in cementitious composites without
affecting its mechanical properties.
Influence of Carbon Nanotubes on the Fracture Surface Characteristics … 511

The 3D scanning results of the fracture surfaces of the tested specimens demon-
strated that the FA-group samples had a higher standard deviation of the macroscopic
undulation height than the Ref-group at the macroscopic level. After mixing 20 wt%
FA, more microcracks and micropores appeared in the cementitious matrix. As a
result, the extension path of the primary fracture became tortuous and complicated
in the FA-group samples. By contrast, due to the ultra-high specific surface area
and bridging effects of CNTs, the formation of hydration products was promoted,
internal microcracks and micropores were effectively reduced, and the regularity and
integrity of the samples were also improved. Therefore, under tensile loading, the
extension path of the final primary fracture was simple, and the standard deviation
of the corresponding macro-fluctuation height was low.
At the microscopic level, energy was continuously absorbed during tensile loading
of the specimens. Compared with the Ref-group specimens, the uneven hydra-
tion reaction inside the FA-group specimens made the energy distribution inside
the sample uneven, resulting in a rougher fracture surface. The addition of CNTs
promoted nucleation and had pore filling effects that optimized the pore structure
and inhibited the development of microcracks in the cement matrix during loading.
The CNT-group specimens could absorb more energy, resulting in more complex
stress paths and rougher fracture surfaces. In conclusion, using 0.08 wt% CNTs
combined with 20 wt% FA to replace the same cement mass in the slurry will reduce
the microcracks and micropores in the hardened matrix, making the change rate of
the macro-undulation height smaller. CNT-reinforced cementitious composites can
absorb more energy during tensile loading. As the energy is released to the stress
concentration zone at the tip of the microcrack, the stress path of the CNT-group
specimens became more complicated, with a more extensive change rate of the
microscopic inclination angle and rougher fracture surface.

Acknowledgements This study was supported by the Natural Science Foundation of the Jiangsu
Higher Education Institutions of China under Grant (22KJB560010) and Nantong Basic Science
Research Program of China under Grant (No. JC12022098).

Author Contributions The manuscript was written through contributions of all authors. All
authors have given approval to the final version of the manuscript.

Declaration of Competing Interests The authors declare that they have no known competing
financial interests or personal relationships that could have appeared to influence the work reported
in this paper.

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Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Effect of Carbon Nanotubes
on the Acoustic Emission Characteristics
of Cemented Rockfill

Z. Yu, H. Jing, Y. Gao, X. Wei, and A. Wang

Abstract The use of carbon nanotubes (CNTs) to reinforce cemented rockfill is

attracting considerable interest due to the remarkable improvement in performance
and the extremely low dose of the added nanomaterial. To reveal the enhancement
mechanism of the CNTs on cemented rockfill, the acoustic emission (AE) charac-
teristics of cemented rockfill specimens during the Brazilian split test were investi-
gated. The results demonstrated that CNTs improved tensile strength by 17.2% and
decreased the AE count. The nucleation and micropore-filling effects of the CNTs
promoted the cement hydration reaction and formation of a denser structure, thereby
improving resistance to loads. Meanwhile, finer pores avoid stress concentration,
resulting in AE activity becoming more sparse. Finally, the AE b-value increased
by 14.8%, which further indicated that the overall failure process was at a lower
intensity.

Keywords Acoustic emissions · Carbon nanotubes · Cemented rockfill · Failure

patterns

1 Introduction

Carbon nanomaterials, including carbon nanotubes (CNTs), graphene oxide and

carbon nanosheets, have been widely recognized for their enhancement of cement-
based composites [1, 2]. The strengthening mechanism of CNTs in cement-based
composites has been the focus in recent decades [3, 4], and it is though physical
properties such as extremely high mechanical strength and large specific surface

Z. Yu · H. Jing (B) · X. Wei · A. Wang

State Key Laboratory for Geomechanics and Deep Underground Engineering, China University of
Mining and Technology, Xuzhou, China
e-mail: hongwenjingcumt@126.com
Y. Gao
School of Transportation and Civil Engineering, Nantong University, Nantong, China

© The Author(s) 2023 513

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_54
514 Z. Yu et al.

area that CNTs fill microscopic pores in cement matrix and promote the develop-
ment of hydration products [5]. The resulting compact microstructure greatly opti-
mizes the macroscopic mechanical properties of cement-based composites, including
compressive strength, impermeability and corrosion resistance [6, 7]. At the same
time, the addition of CNTs is potentially a way to reduce CO2 emissions from cement
production [8].
In construction and mining projects, cemented rockfill with higher mechanical
properties is significant for structural stability and production safety. CNTs improve
the mechanical strength of cemented rockfill materials [9], but due to the lack of
sensitivity of current monitoring systems, the effect of CNTs on failure patterns
during the loading process needs further research. Acoustic emission (AE) tech-
nology is a nondestructive testing method widely used in the study of the mechanical
properties of rock materials [10]. During the failure process, the gradual develop-
ment of microfractures will cause acoustic signals, which can be monitored by the
AE system. The waveform of the AE signal reflects the microscopic failure pattern.
For example, the RA (Rise time to Amplitude ratio) and AF (AE counts to Duration
ratio) can be used as criteria for the type of failure [11]. Therefore, the AE technique
is effective for studying the destruction mode of the cemented rockfill.
In this study, we add a very low content of CNTs with fly ash (FA) to partially
replace cement in the cemented rockfill. The tensile failure strength of the specimens
was measured by the Brazilian split test and simultaneously the AE system collected
signals during the destruction process. Combing the AE activity, and the stress–-
strain curve, the distribution and intensity of the failure events in each loading stage
of the CNT-enhanced specimens were analyzed. The failure event accumulation of
the specimen is discussed according to the peak and growth rate of the cumulative
AE counts curve. The b-value in AE technology is calculated to characterize the
severity of the overall failure process. The findings in this study promote the engi-
neering application of CNTs and we discuss the enhancement mechanism of CNTs
in cemented rockfill materials.

2 Methods

Waste coal gangue extracted from a coal mine in Shanxi province, China, was selected
as the solid particles in cement rockfill. The composition of coal gangue does not
react chemically with the cement. Ordinary Portland cement (PO. 42.5), conforming
to the Chinese Standard GB-175-2007 [12], was the binder material. Multiwall
CNTs (MWCNTs; manufactured by Nanjing XFNANO Materials Tec. Co., Nanjing
City, Jiangsu Province, China) were added to the rockfill materials. Polycarboxylate
superplasticizer was added to promote CNT dispersion in the suspension.
The preparation of CNT-enhanced cement rockfill samples mainly comprised the
preparation of CNT–cement slurry and mixing of the slurry with solid particles. The
process is detailed in our previous report [13]. A Ref-group and CNT-group were
set in this study, and a 0.4 water-to-cement ratio was applied in both groups. In the
Effect of Carbon Nanotubes on the Acoustic Emission Characteristics … 515

Fig. 1 Brazilian split test and acoustic emission (AE) signal monitoring

CNT-group mortar, there was 20% FA with 0.05 wt% CNTs to replace 20% cement.
The other preparation processes of the two groups were consistent.
We chose 28-day-old standard-cured disc samples with a diameter of 50 mm and
a thickness of 25 mm for the Brazilian split tests. The loading rate was 0.10 mm/min.
To avoid the influence of errors in the test and improve accuracy, three samples were
selected for each group. During the loading process, the AE signals were monitored
by a Micro-II AE system (developed by the American Physical Acoustic Corpora-
tion), as shown in Fig. 1. The AE threshold and frequency were set at 30 dB and 10
Msps, respectively.

3 Results

3.1 Mechanical Properties of Cemented Rockfill

After curing for 28 days, the peak tensile strength of the specimens is shown in Fig. 2.
The peak strength of the CNT-group was 3.4 MPa, an increase of 17.2% compared
with the Ref-group, indicating that CNTs significantly enhanced the ability of the
sample to resist tensile stress, which was consistent with previous studies [14]. Higher
tensile strength can reduce the risk of sudden failure and improve engineering safety.
The increase in tensile strength was attributed to the strengthening effect of CNTs
in the cement matrix. CNTs show a bridging effect between microfractures in the
cement matrix [15]. When the sample is under tensile stress, the high tensile strength
of CNTs can help the cement matrix resist external stress and avoid microcrack devel-
opment [16]. As the sample’s hydration reaction continued, the hydration products
516 Z. Yu et al.

Fig. 2 Tensile strength of

two groups of cemented
rockfill specimens. CNTs,
carbon nanotubes

became more closely connected with the CNTs and the resultant compact internal
structure enabled the specimen to resist failure and deformation.

3.2 Analysis of AE Activity During the Failure Process

The stress–strain curve and real-time AE signals of the samples are shown in Fig. 3.
As the stress gradually increased, the stress–strain curve divided into four failure
stages. In the initial compaction stage and elastic deformation stage, there was almost
no fracture inside the sample, accompanied by sparse AE signals. During the plastic
deformation stage, cracks gradually developed and the AE activity gradually became
denser. When peak strength was reached, the samples suddenly produced a lot of AE
counts, corresponding to the development of major fractures. In the post-peak stage,
cracks continue to expand.
With the addition of CNTs to the cemented rockfill, the AE activity became sparse
and AE counts decreased significantly during the failure process. The cumulative

Fig. 3 Acoustic emission counts of the Ref-group (a) and CNT-group (b) during the Brazilian split
test. CNTs, carbon nanotubes
Effect of Carbon Nanotubes on the Acoustic Emission Characteristics … 517

Fig. 4 The b-value of two

groups of cemented rockfill
specimens. CNTs, carbon
nanotubes

counts of the Ref-group reached 20(×103 ), much higher than in the CNT-group. At
the peak strength point, the Ref-group showed a single and high-count AE event,
whereas the CNT-group showed multiple peaks of AE activity. The reason for this
phenomenon is the promotion of a dense microstructure by CNTs. The defect area is
smaller, reducing the severity of failure, resulting in fewer counts and lower frequency
of AE events, which leads to more gradual macro disruption.
The b-value reflects the ratio of the number of lower and higher grade AE events
during the rock failure process, and can be calculated by the following equation [17]:

lg N( A/20) = a − b × (A/20) (1)

where A is the amplitude of the AE event, N(A/20) is the cumulative frequency of AE

events with an amplitude ≥ A, and b represents the b-value.
According to Eq. 1, the b-value of the CNT-group reached 1.27, increasing 14.8%
more than the Ref-group (Fig. 4), which reflects the AE activity being mostly concen-
trated at lower energy levels during the failure of the sample. The reason behind this
phenomenon is that CNTs promote the development of hydration products to reduce
the size of micropores, further avoiding the high energy release of cracks during
loading.

4 Discussion and Conclusions

In this study, CNTs and FA were added to a cemented rockfill mix to investigate the
production of environmentally friendly and high-workability materials. The tensile
properties of the three specimens in each group were tested by the Brazilian split test
and AE signals were monitored by a PCI-2 system. The failure mode pattern of the
518 Z. Yu et al.

CNT-reinforced cement rockfill material was studied by the stress–strain curve and
AE waveform.
In the Brazilian split experiment, the CNT-reinforced sample showed increased
peak tensile strength by 17.2%. At the same time, the AE events became sparser and
the count decreased, indicating that the intensity of destruction decreased. Finally,
the addition of CNTs increased the b-value by 14.8%, which indicated that CNT-
reinforced cemented rockfill material will have fewer high energy release destruction
events, thereby making it a safer option.

Acknowledgements This study was supported by Yunlong Lake Laboratory of Deep Underground
Science and Engineering Project (No. 104023002), National Natural Science Foundation of China
(No. 52074259), Assistance Program for Future Outstanding Talents of China University of Mining
and Technology (No. 2022WLJCRCZL050) and Postgraduate Research & Practice Innovation
Program of Jiangsu Province (No. KYCX22_2580).

Author Contributions The manuscript was written through contributions of all authors. All
authors have given approval to the final version of the manuscript.

Declaration of Competing Interest The authors declare that they have no known competing
financial interests or personal relationships that could have appeared to influence the work reported
in this paper.

References

1. Gao Y, Jing HW, Chen SJ, Du MR, Chen WQ, Duan WH (2019) Influence of ultrasoni-
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Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Effects of Graphene Oxide Content
on the Reinforcing Efficiency of C–S–H
Composites: A Molecular Dynamics
Study

W. Chen, J. Xiang, Y. Gao, and Z. Zhang

Abstract Determining the graphene oxide (GO) content is the key to applying GO
to reinforce the mechanical performance and durability of cementitious compos-
ites. However, most of the previous studies are conducted from the perspective of
experiments and lack elaboration on the mechanism of the GO-reinforced cementi-
tious composite under different GO content. Hence, we investigated the effect of GO
content on the reinforcing efficiency of calcium–silicate–hydrate (C–S–H) to trade off
the enhancement of GO in cementitious composites and the corresponding economic
benefits. The results demonstrated that an appropriate number of GO nanosheets can
reinforce the cementitious composite with simultaneous high enhancing efficiency
and economic benefits. The microdamage evolution of GO/C–S–H composites and
the GO reinforcing mechanisms are reported. Our findings deepen the understanding
of the enhancing mechanisms of GO embedded in C–S–H nanocomposites and help
to determine the suitable GO content in practical engineering.

Keywords Cementitious composites · Graphene oxide · Mixing content ·

Molecular dynamics simulations

W. Chen (B)
Department of Mechanical, Aerospace and Civil Engineering, The University of Manchester,
Manchester, UK
e-mail: weiqiang.chen@manchester.ac.uk
J. Xiang
Key Laboratory of Rock Mechanics and Geohazards of Zhejiang Province, Shaoxing University,
Shaoxing, China
Y. Gao (B)
School of Transportation and Civil Engineering, Nantong University, Nantong, China
e-mail: Y.Gao@ntu.edu.cn
Z. Zhang
Department of Chemical Engineering and Analytical Sciences, The University of Manchester,
Manchester, UK

© The Author(s) 2023 521

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_55
522 W. Chen et al.

1 Introduction

By virtue of superior mechanical properties, ultra-high specific area and abundant

oxygen-containing functional groups, graphene oxide (GO) exhibits outstanding
performance in optimizing the microstructure [1], improving mechanical properties
[2], enhancing permeability resistance [3], increasing durability [4], and improving
the thermal and electrical conductivity [5] of cement-based materials through nucle-
ation effects [6] and pore-infilling effects [7]. Furthermore, some latest studies indi-
cate that graphene-based enhancement of cementitious composites could be a cost-
effective strategy for decreasing CO2 emissions and energy consumption in practical
engineering [8].
The concentration of GO sheets significantly influences the mechanical perfor-
mance of cement-based grouting materials when they are incorporated into the
cementitious material. In the past study, researchers have primarily determined the
proportion of GO with the help of experiments and experience [2]. There are few
studies on the mechanism of composite enhancement by different ratios of GO.
Hence, to better understand the regulatory effect of GO on calcium–silicate–hydrate
(C–S–H) gel, it is necessary to understand the effect of different mixing concen-
trations of GO sheets on the mechanical properties of C–S–H gels. Therefore, in
this study, we built a molecular dynamics (MD) model with different ratios of GO
nanolayers incorporated into the C–S–H nanocomposite, and performed tensile tests
to investigate the effects of the GO content on the reinforcing efficiency of C–S–H
composites and the enhancing mechanism.

2 Methods

Following a previous study [9] and using the same model parameters of GO and
C–S–H gel, GO/C–S–H nanocomposites were built in which one or two parallel GO
nanolayers were embedded in the C–S–H matrix and the spatial orientation of the
sheets was aligned with the x direction (Fig. 1). The whole system was firstly relaxed
for 2 ns in the NPT ensemble at 300 K and 0 atmosphere to achieve equilibrium.
Next, tensile deformation was applied to the nanocomposite with a stretching rate of
1 m/s. The NPT ensemble was applied in the y and z directions during the tensile
deformation at 300 K and 0 atmosphere. The tensile process was completed when
the tensile strain in the x direction reached 110%. The C–S–H gel was described
by the CLAYFF force field [10] and GO was described by the consistent-valence
forcefield [11], with a cutoff radius of 1.5 nm for both Lennard–Jones 12–6 potential
and Coulomb electrostatic potential. The intermolecular force between the C–S–
H gel and GO nanolayer was described by the Lennard–Jones 12–6 potential and
Coulomb electrostatic potential, and the corresponding parameters were derived by
Effects of Graphene Oxide Content on the Reinforcing Efficiency … 523

Fig. 1 Atomic model of

GO/C–S–H nanocomposite
with two GO nanolayers.
The model with only one GO
nanolayer can be found in
Fig. 3a and is not shown here

the Lorentz–Berthelot combining rules. The time step was set as 0.5 fs and the long-
range electrostatic interactions were resolved by the particle–particle–particle-mesh
method.

3 Results and Discussion

3.1 Tensile Mechanical Properties of the GO/C–S–H Model

The ultimate tensile strength of the MD system did not change significantly after
embedding different numbers of GO nanosheets, and the tensile stress–strain curves
are presented in Fig. 2. Similar to the visible test results [2], when the MD model
reached its peak strength, some damage was generated inside the C–S–H model, but
the C–S–H model still had a certain amount of residual strength until its complete
failure after further stretching. The monolayer GO/C–S–H model underwent damage
at a tensile strain of 0.5. By contrast, the tensile ductility of the bilayer GO/C–S–H
model was significantly enhanced to 0.9, with an enhancement ratio of 80%.

3.2 Failure Process of GO/C–S–H Models Under Stretching

In order to reveal the enhancement mechanism of GO incorporated into C–S–H,

typical atomic model images (Fig. 3) during stretching were selected to display the
microdamage evolution of the C–S–H models with different GO layers embedded.
Figure 3 shows that the deformation of the GO/C–S–H model can be divided into
the following stages: structural loosening stage, pore development stage, and failure
stage.
524 W. Chen et al.

Fig. 2 Mechanical responses of GO/C–S–H models with different GO content under stretching

Fig. 3 Microdamage evolution of C–S–H models with different GO content: a single layer and
b two layers
Effects of Graphene Oxide Content on the Reinforcing Efficiency … 525

In the monolayer GO/C–S–H model, the C–S–H structure loosened as the tensile
strain reached 0.2. As the tensile strain increased to 0.5, noticeable pores began to
appear in the structure, but there was still a certain amount of bridging C–S–H, so
the model was still a monolithic structure. As the tensile strain continually increased
to 0.6, the C–S–H completely separated and fractured, consistent with Fig. 2. At the
same time, the molecular structure of the bilayer GO/C–S–H model slowly started
to be loosen. Pores gradually appeared and expanded when the strain increased from
0.6 to 1.1. At this time, the C–S–H gel still had residual strength and ability to
resist tensile deformation. In addition, it was not hard to find that with increasing
deformation, the molecular structure around the GO nanosheets became more closely
arranged, especially the C–S–H molecules sandwiched between the two GO layers.
It can be concluded that the ductility of the C–S–H was greatly improved with the
addition of two GO nanolayers compared with the intercalation of a single layer
of GO. Our results also proved that an appropriate amount of GO can effectively
enhance the stretching mechanical performance of cement-based materials.

4 Conclusions

The effects of incorporating different numbers of layers of GO nanosheets on the

tensile performance of C–S–H gels were studied by MD simulations, the micro-
damage evolution of GO/S–H models in the tensile process were investigated, and
the enhancement mechanism of GO on the tensile properties of hydrated calcium sili-
cate composites was analyzed at the atomic scale. Our main conclusions are listed
below.
(1) Incorporating an appropriate amount of GO was able to effectively improve the
ductility of hydrated calcium silicate composites. MD simulations showed that
the bilayer GO/C–S–H model increased the post-peak plastic strain by 80%
compared with a monolayer GO/C–S–H model.
(2) The deformation of the GO/C–S–H model can be divided into structural loos-
ening stage, pore development stage, and failure stage. Compared with the
monolayer GO/S–H model, the strain values corresponding to the bilayer GO
model entering the third stage were delayed.
(3) At the atomic scale, the structure of C–S–H around the GO nanosheets was more
compact, showing a higher ability to resist tensile deformation, and also reflected
the enhancement effect of an appropriate amount of GO on the mechanical
properties of C–S–H composites.

Acknowledgements This study was supported by the Zhejiang Provincial Natural Science
Foundation of China (No. LQ21E040003).

Author Contributions The manuscript was written through contributions of all authors. All
authors have given approval to the final version of the manuscript.
526 W. Chen et al.

Declaration of Competing Interest The authors declare that they have no known competing
financial interests or personal relationships that could have appeared to influence the work reported
in this paper.

References

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cation on the dispersion and enhancing effect of graphene oxide–carbon nanotube hybrid
nanoreinforcement in cementitious composite. Compos B Eng 164:45–53
8. de Souza FB, Yao X, Gao W, Duan W (2022) Graphene opens pathways to a carbon-neutral
cement industry. Sci Bull 67:5–8
9. Gao Y, Jing H, Wu J, Fu G, Feng C, Chen W (2022) Molecular dynamics study on the influence
of graphene oxide on the tensile behavior of calcium silicate hydrate composites. Mater Chem
Phys:126881
10. Cygan RT, Liang J-J, Kalinichev AG (2004) Molecular models of hydroxide, oxyhydroxide,
and clay phases and the development of a general force field. J Phys Chem B 108:1255–1266
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Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.
Graphene-Induced Nano- and Microscale
Modification of Polymer Structures
in Cement Composite Systems

Z. Naseem, K. Sagoe-Crentsil, and W. Duan

Abstract Redispersible polymers such as ethylene–vinyl acetate copolymer (EVA)

have attracted attention in construction due to their enhanced flexural strength, adhe-
sion, flexibility and resistance against water penetration. However, EVA particles
cluster in a highly alkaline cementitious matrix and exhibit poor interaction with the
cement matrix. The underlying mechanism of poor dispersibility of EVA is attributed
to hydrophobic groups of polymers, a variation in the adsorption rate and molec-
ular diffusion to the interface where they cluster together. This phenomenon can
negatively affect the fresh properties of cement and produce a weak microstruc-
ture, adversely affecting the resulting composites’ performance. This study high-
lights how graphene oxide (GO) nanomaterial alters the nano- and microscale struc-
tural characteristics of EVA to minimize the negative effects. Transmission electron
microscopy (TEM) revealed that the GO sheets modify EVA’s clustered nanostructure
and disperse it through electrostatic and steric interactions. Furthermore, scanning
electron microscopy (SEM) confirmed altered microscale structural characteristics
(viz. surface features) by GO. The altered and enhanced material scale engineering
performance, such as the compressive strength of the resulting cement composite,
was notable.

Keywords Cement composites · Dispersion · Graphene oxide · Ethylene-vinyl

acetate copolymer · Interactions

1 Introduction

Concrete is the most widely used construction material, but its quasi-brittle nature and
low durability can affect its performance. Recently, polymer-modified concrete has
become popular in response to the durability issue, because redispersible polymers

Z. Naseem (B) · K. Sagoe-Crentsil · W. Duan

Department of Civil Engineering, Monash University, Clayton, VIC, Australia
e-mail: Zunaira.naseem@monash.edu

© The Author(s) 2023 527

W. Duan et al. (eds.), Nanotechnology in Construction for Circular Economy,
Lecture Notes in Civil Engineering 356,
https://doi.org/10.1007/978-981-99-3330-3_56
528 Z. Naseem et al.

such as ethylene–vinyl acetate (EVA) can modify and enhance flexural strength, adhe-
sion, flexibility and resistance against water penetration [1, 2]. However, EVA shows
poor interaction with the highly alkaline cement matrix, which adversely affects
the material scale performance, such as the compressive strength of the resulting
composite [3, 4].
Polymer additives can cluster in a highly alkaline cement composite due to their
poor interaction with the cement matrix [5]. The resultant weak microstructure dete-
riorates engineering performance, such as the compressive strength of the cement
composite [6, 7]. The underlying mechanism is the hydrophobic groups present in
the macro-molecular long chain polymer [8]. Additionally, there is variation in the
adsorption rate and molecular diffusion to the interface, where they cluster together
and interact poorly with the cementitious environment [8, 9].
In this regard, a two-dimensional (2D) nanomaterial such as graphene oxide (GO)
can potentially modify the nano- and microscale characteristics by its unique phys-
ical and chemical properties and larger surface area [10, 11]. In addition, outstanding
mechanical properties of GO have been widely reported to enhance the compres-
sive and tensile properties of resulting cement composites [12, 13]. In addition, the
abundant oxygenated groups of GO can strongly interact with cement particles and
modify their microstructure [14, 15]. As reported previously, adding a low dosage
of 0.05% GO can significantly enhance the engineering performance of cement
composites [16]. Muhammad et al. also reported that GO incorporation could alter
the microstructure and enhance the transport properties of cement composites [17].
In the present work, we investigated the effect of 2D GO sheets on the nano-
and microscale characteristics of the EVA polymer. Scanning electron microscopy
(SEM) and transmission electron microscopy (TEM) were utilized to determine
the effects of GO on the polymer’s structural features. Engineering performance,
such as compressive strength, was evaluated using the Instron 4204 50KN. Our
experimental results revealed that the addition of GO altered the nano- and microscale
characteristics of EVA, resulting in uniform dispersion of the EVA polymer and
enhanced performance in the alkaline cement environment.

2 Methods

2.1 Nano- and Microscale Structural Characterization

Nanoscale structural characterization of samples was performed with TEM using a

FEI Tecnai T20 instrument operated at 200 kV. Comprehensive nanoscale charac-
terization of EVA polymer, as well as evaluation of the effect of the 2D GO sheets
on the polymer, was performed in comparison with the reference EVA polymer. The
microscale surface characteristics were investigated by SEM (FEI Nova NanoSEM
450 FEG SEM) under an accelerating voltage of 5 kV and the effect of incorporating
GO was compared with reference EVA polymer.
Graphene-Induced Nano- and Microscale Modification of Polymer … 529

2.2 Specimen Preparation for Performance Evaluation

in Alkaline Cement Matrix

EVA polymer powder was first dispersed in water, followed by the addition of GO
solution before ultrasonication for 10 min. Next, the solution was mixed with ordinary
Portland cement (OPC) according to the procedures specified in ASTM C1738-11a
to prepare cube-shaped specimens that were vibrated for the 30 s, covered with
polyethylene sheets and demolded after 24 h, followed by the curing method. The
engineering performance was specifically investigated as the compressive strengths
of the cubic specimens. A universal loading machine, the Instron 4204 50KN, was
used to test the cement specimens at 28 days of age. For each batch, a minimum
of five specimens was tested and the average values were taken as the compressive
strength.

3 Results and Discussion

3.1 Modification of Nano- and Microstructural

Characteristics by GO

The nanostructure of the EVA polymer as examined by TEM is shown in Fig. 1. EVA
polymer exhibited a clustered structure of polymer particles (Fig. 1a), which could
be attributed to the rate of change of adsorption and diffusion, causing aggregation
of the polymer particles. One of the critical reasons for their poor interaction with
the highly alkaline cementitious environment is their aggregation, which deteriorates
the material scale performance of the resultant cement composites. Remarkably, as
presented in Fig. 1 (b, c), the incorporation of GO sheets disperses the polymer parti-
cles uniformly compared with the reference EVA polymer. The underlying reason
could be the unique physical and chemical properties and larger surface area of GO,
which alters the EVA polymer’s nanostructure. In addition, effective electrostatic
and steric interactions between the GO sheets and polymer molecules could hamper
their aggregation [18].
Furthermore, SEM also showed the microscale characteristics, which confirmed
the changes in the the nanoscale characteristics. Aggregated surface features were
present in the reference EVA polymer (Fig. 2a). The addition of GO altered the
microscale structure characteristics of the polymer (Fig. 2b), which was consistent
with TEM results and confirmed the potential of GO to modify the EVA polymer
characteristics at the nano- and microscale due to its unique physical and chemical
properties and abundant functional groups on the basal plane that improved the
interaction and material scale performance of the resultant cement composites [10].
530 Z. Naseem et al.

Fig. 1 Nanoscale structural characteristics by TEM. a EVA polymer shows clustering of molecules
and aggregation, which causes poor interaction with the cement matrix. (b, c) GO dispersion
of clustered polymer structure through electrostatic and steric interactions and alteration of the
nanostructure

Fig. 2 Microscale surface features by SEM. a EVA polymer shows aggregated surface features at
the microscale. b Altered aggregated surface features of the polymer with GO addition compared
with the reference EVA polymer in (a)

3.2 Engineering Performance in the Alkaline Cement

Environment

The material scale performance of the EVA polymer with and without GO in the
highly alkaline cementitious environment was further assessed. Specifically, the
compressive strength of the prepared cement composite specimens was investigated
at 28 days of hydration age. As shown in Fig. 3, the EVA-incorporated cement
composite (PMC) exhibited a lower compressive strength than OPC composites.
The underlying reason is the clustered polymer structure and the poor interaction
with the highly alkaline cement matrix, resulting in a deteriorated performance of
the PMC. Remarkably, the addition of GO improved and enhanced the compres-
sive strength by ~40% higher than the reference PMC samples. The underlying
phenomena could be attributed to altered nano- and microscale characteristics of
Graphene-Induced Nano- and Microscale Modification of Polymer … 531

Fig. 3 Compressive strength of cement samples showing significantly enhanced compressive

strength compared with the reference PMC sample

EVA by the GO sheets. In addition, the presence of GO can disperse the polymer
particles through electrostatic and steric interactions and hamper their aggregation in
the alkaline cementitious environment. As a result, significantly enhanced material
scale performance was observed in the resultant cement composites (GO-PMC).

4 Conclusion

In this study, we used the nanomaterial GO as a novel approach to altering and

improving the structural characteristics and the material scale performance of poly-
mers in cementitious environments. TEM showed that GO altered the aggregated
surface characteristics of EVA at the nanoscale, which was further confirmed by SEM
displaying the modified surface features compared with the reference EVA polymer.
The compressive strength of the composite specimens was quantified to signify their
engineering performance at the material scale. Compared with the reference EVA
composite, the GO–polymer-modified composite (GO-PMC) achieved ~ 40% higher
compressive strength. This improvements in engineering performance is attributed
to the altered nano- and microscale characteristics of the polymer by GO causing
electrostatic and steric interactions and hampering particle aggregation in the cement
matrix.

Acknowledgements The authors are grateful for the financial support of the Australian Research
Council in conducting this study. They acknowledge the use of facilities within the Monash Center
of Electron Microscopy.
532 Z. Naseem et al.

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Open Access This chapter is licensed under the terms of the Creative Commons Attribution 4.0
International License (http://creativecommons.org/licenses/by/4.0/), which permits use, sharing,
adaptation, distribution and reproduction in any medium or format, as long as you give appropriate
credit to the original author(s) and the source, provide a link to the Creative Commons license and
indicate if changes were made.
The images or other third party material in this chapter are included in the chapter’s Creative
Commons license, unless indicated otherwise in a credit line to the material. If material is not
included in the chapter’s Creative Commons license and your intended use is not permitted by
statutory regulation or exceeds the permitted use, you will need to obtain permission directly from
the copyright holder.