Design of the Main Spans, Second Gateway Bridge,

Brisbane
John Connal, Industry Director Structures, Maunsell Australia Pty Ltd
Ken Wheeler, Technical Director, Maunsell Australia Pty Ltd
Andrew Pau, Principal, Cardno (NSW) Pty Ltd
Miho Mihov, Principal, Cardno (NSW) Pty Ltd
SYNOPSIS
The Second Gateway Bridge crosses the Brisbane River to the east of the city of
Brisbane and is of overall length of 1627m with a main span of 260m and adjacent
spans of 162m. The superstructure for the three main river spans comprises a twin
cell prestressed concrete box girder constructed by the cast in-situ balanced
cantilever method with segments varying from 15.57m to 5.2m deep and 3m to 5m
long. The main spans are supported on large groups of 1.8m diameter bored piles
and twin blade piers. The design of the Second Gateway Bridge was delivered by an
integrated design team of Maunsell AECOM and Cardno for the Maunsell SMEC
Joint Venture, with Coffey Geosciences performing foundation design.
This paper describes the design development process for the main river spans
including the determination of the design criteria, selection of the superstructure
cross section and structural form for the substructure. Details are provided of the
investigations and testing carried out to confirm the design assumptions and how the
design was prepared to facilitate the subsequent construction phase.
1.0

GATEWAY UPGRADE PROJECT

In September 2005, the Queensland Motorways Limited called for tenders for the
upgrade of approximately 25km of the Gateway Motorway including a second bridge
over the Brisbane River. The successful consortium was required to design, build
and maintain the upgraded facility for 10 years. The upgrading required duplication of
roadway and bridges south of the Brisbane River, a new bypass north of the
Brisbane River, and a new second crossing of the Brisbane River to allow traffic to be
converted to one-way northbound on the existing Gateway Bridge, and one-way
southbound on the new, Second Gateway Bridge.
A consortium comprising Leighton Contractors Pty Limited and Abigroup Contractors
Pty Limited formed the Leighton Abigroup Joint Venture (LAJV) to successfully bid
the project which was awarded in October 2006. The design and construction of the
Second Gateway Bridge is under the control of the Gateway Bridge Alliance
comprising the Leighton Abigroup Joint Venture and VSL Australia. The design of
the Second Gateway Bridge was delivered by an integrated design team of Maunsell
AECOM and Cardno for the Maunsell SMEC Joint Venture, with Coffey Geosciences
performing foundation design. The design commenced in October 2006, was
completed during 2007 and the construction of the works is scheduled to be
completed by August 2010, following which the existing Gateway Bridge will be
refurbished to allow completion of the motorway upgrade.

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2.0

DESIGN REQUIREMENTS AND DESIGN PHILOSOPHY

The Second Gateway Bridge is constructed just 50m to the east of the existing
crossing and was required to have substantially the same appearance as the existing
bridge, except it was also required to have:
The same structural form.
Provide for six lanes of southbound traffic and a 4.5m wide shared
pedestrian/cycle path on its eastern edge.
Piers of the same form but the width could match the width of the box girder.
All piers in the same positions.
A river navigational clearance that matches the existing bridge,
A bridge height that sits below the obstacle limitation surface of the nearby
Brisbane Airport, and
A 300 year design life for durability.
These requirements were specified in the Project Scope and Technical
Requirements (PSTR). This document implied the new bridge would be constructed
using similar techniques to the existing bridge and was directed at achieving a
continuous prestressed concrete bridge with no joints and continuity of
reinforcement, thought to be the most durable bridge form and one more able to
achieve the long life targeted by the 300 year design life. Although the existing
bridge remains in excellent condition and is a credit to its designer and constructor,
replicating this form was not considered appropriate for the Second Gateway Bridge
because of the abundance of bearings and the resulting maintenance and
replacement issues, and also because the structural form has a fundamental lack of
redundancy. Advances in technology that reduce construction cost and time as well
as increasing durability were sought. This approach was embraced by the client and
some of the requirements in the PSTR were modified accordingly.
The design philosophy of the Second Gateway Bridge has sought to maximise
service life and address the issue of the lack of redundancy of the existing bridge.
Current technologies and materials are applied to achieve these goals and also
reduce construction time and cost.
3.0

BRIDGE DESCRIPTION

3.1

Form and Articulation

The Second Gateway Bridge comprises 18 spans and is of overall length 1627m
between abutments. The span configuration has been chosen to match that of the
existing Gateway Bridge, with piers and abutments at the same alignments, except
for the two piers that flank the main span. The structure is divided longitudinally into
four modules of lengths 287m in the southern approaches, 701m for the main river
crossing and two modules of lengths 352m and 287m in the northern approaches
(refer to Figure 1).
Each module comprises a continuous frame with the
superstructure integral with the piers. The deck modules are separated by three
superstructure halving joints incorporating deck expansion joints and guided sliding
pot-type bearings. At both abutments, guided sliding pot bearings provide for
longitudinal expansion. The halving joints provide a means of accommodating
relative longitudinal displacements and have been positioned away from the piers, as

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in the existing bridge. This is structurally efficient, minimizes the number of bearings,
and keeps the tops of the piers compact in width.

Figure 1 Bridge Elevation

The vertical alignment of the main span superstructure is dictated by the road
geometry and the geometric restrictions of the shipping navigational and aircraft
clearance envelopes, and matches the existing bridge. The superstructure each side
of the main span is on a constant grade of 5.3% with a vertical circular curve
provided between the main river piers.
3.2

Superstructure

The main span bridge superstructure comprises a two cell prestressed concrete box
girder with vertical webs, supporting six 3.5m wide traffic lanes, two 0.5m wide
shoulders and a 4.25m wide shared path separated from the traffic lanes by an
intermediate concrete traffic barrier. The overall deck width is 27.5m which is wider
than the existing bridge because of the presence of the shared path. The deck has
two-way 2% deck crossfall with the crown at the middle of the box (refer to Figure 2).
The depth of the box girder varies from 15.57m at the two main river piers, to 5.20m
deep at mid-span of the main span, and transitioning to a constant depth of 3.3m in
the approach spans.

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Figure 2 Main Span Cross Section

At the two main piers, two-cell pier-boxes integrate the superstructure with the piers.
Each pier-box comprises 2.5m thick transverse diaphragms which are extensions of
the twin blade pier columns. In the longitudinal direction, each of the webs has been
increased in thickness to 1.0m wide. This arrangement of webs and transverse
diaphragms provides a direct transfer of girder shear forces into the substructure.
3.3

Substructure

Each river pier consists of twin, box section, reinforced concrete columns of outer
dimensions 15.0m by 2.5m, supported by a reinforced concrete pilecap, which in turn
is supported by 24 No. 1.8m diameter bored piles. The twin blade piers are similar in
appearance to the existing bridge twin blades, and provide the necessary longitudinal
flexibility in-service, whilst providing adequate stiffness and stability for the
construction stage cantilevering (refer to Figure 3). The design of the twin blade
piers was governed by in-service conditions, and the effects of second-order
geometric non-linearity were considered.

Figure 3 River Piers

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The main pier pile caps have a plan area of 19.5m by 17.6m and are 3.2m thick. The
pile caps support an outer pier protection system (refer to Figure 4). The tops of the
pile caps have been set at Reduced Level 4.2m, which allows construction of the pile
caps to be carried out above water using land-based methods. Considerable savings
in construction time resulted from this cost effective solution, which also eliminates
the additional risks associated with marine-based work and construction below river
water level. Precast skirt units attached to the pile caps for pier protection from low
impacts also provide the necessary enclosure of the piles during low tidal levels.

Figure 4 River Pier Pilecaps

The pile layout comprises two groups of piles directly located under each pier blade
at 3.3m centres transversely and 3.6m centres along the axis of the bridge. The piles
are 1.8 m diameter bored piles, cased through the upper soft alluvial and clay layers.
The casing achieved a seal in the upper zone of a carbonaceous siltstone layer and
coring continued through that layer and into the inter-bedded siltstone and sandstone
layers below, to form the rock sockets that found at depths ranging between 25-55m
depth and provide vertical load capacity via a combination of side friction and end
bearing.
4.0

DESIGN CRITERIA

4.1

General

The design criteria were based on the Australian Bridge Design Code AS5100 (Ref
1), supplemented by additional requirements nominated by the Main Roads
Department, Queensland for Queensland Motorways. Reference was also made to
the AASHTO Load Resistance Factor Design (LRFD) Code (Ref 2) for inclusion of
additional loadings applicable to the construction stage for this form of bridge
structure and to the design requirements for shear and torsion appropriate to large
box girders.

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4.2

Design Life

The PSTR specified that the Second Gateway Bridge primary structural elements
shall have a design life of 300 years for durability. As established guidelines, and the
guidance in AS 5100 in particular, address a design life of "only" 100 years, it was
therefore necessary to take a first principles approach based on building in the
required durability at the outset, where feasible, and minimising the need to take
measures later in the life of the bridge to achieve 300 years of service. Integral with
this philosophy was the appropriate selection of high quality materials chosen to
address the particular durability issues that are posed by the range of exposure
classifications experienced by the bridge elements. The selection of high quality
concrete (generally 50MPa), cover of 55mm (generally) and the typical use of black
steel reinforcing with the selective use of stainless steel in splash zone elements,
was the general philosophy adopted. This approach is described in more detail in a
companion paper.
4.3

Dead Loads

The dead load condition was determined by stepping through the proposed
construction sequence for the main spans and cumulating stress resultants. The
time dependent effects of creep and shrinkage were incorporated in the
determination of the long-term dead load condition. The creep and shrinkage
behaviour of the concrete for the superstructure was the subject of testing to ensure
these properties were realistically predicted in the design. This testing was seen as a
good safeguard against unexpected long term creep behaviour of the bridge. This
testing was conducted on larger size test samples and led to a modification of the
shrinkage strains that would otherwise be determined from the Australian design
code, but the code values for creep strains were retained.
4.4

Live Loads

The bridge was designed for six lanes of the SM1600 traffic design load in
accordance with AS5100. In addition, consideration was given to the HLP400 heavy
load platform, which was restricted to travel between the outer box girder webs. The
HLP400 loading was applied concurrently with half of four design lanes of SM1600
traffic loading.
The above design live loads were applied in combination with a global shared path
loading of 3kPa. A 7kPa localised shared path live load was also applied on any
30m length of the footway.
Notwithstanding the fact that the shared path is located on the eastern side of the
Second Gateway Bridge, each box girder cantilever was designed for the same traffic
loading to give flexibility for possible future traffic use.
4.5

Balanced Cantilever Construction Loads

Under the assumed balanced cantilever construction of the superstructure, the

following additional loadings were considered:
Asymmetric concreting of one full segment on the river side of the main piers;
Formwork travellers for the main spans, assumed as 265 tonnes;
Construction live load of 0.25kPa;

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Density variation from one cantilever to the other of 2% to account for systematic
errors in concrete dimensions;
Differential wind loading from one cantilever to the other of 100% and 60% of the
construction wind load;
Allowance for 50mm vertical jacking to correct potential misalignment of the main
span cantilever arms prior to casting closure segments;
Loss of traveller case.
Load factors and load combinations for the above construction loads at the ultimate
limit state were taken from the AASHTO LRFD design code.
4.6

Wind Loads

A 2000 year average return interval was adopted for the in-service ultimate limit state
together with a 20 year average return interval for the in-service serviceability limit
state. The average return interval for the construction ultimate limit state was
adopted as 40 years. The corresponding regional basic wind design wind speeds
were determined using AS5100.
Drag coefficients for the superstructure were determined from AS5100 and compared
to the coefficients adopted in the design of the Existing Gateway Bridge These were
found to correlate closely.
The effect of wind turbulence and buffeting due to the close proximity of the new
bridge to the existing bridge was investigated via sectional models in a wind tunnel.
No instability of either structure (depending on the direction of the wind) was found
up to wind speeds beyond the ultimate limit state wind speed (63m/sec).
4.7

Earthquake Loads

A design acceleration response spectrum for the bridge was developed assuming an
acceleration coefficient of 0.10 (100 year design life), an Importance Factor of 1.25
(structure essential to post-earthquake recovery) and a site factor specifically
determined for the site. A three dimensional response spectrum analysis was carried
out using a structural response factor of 2.8. The use of a low structural response
factor allowed simplified reinforcement detailing to be used.
4.8

Ship Impact Protection

The first line of defence against ship impact on the main river piers is the submerged
arrestor islands that are created by reducing the construction platforms from which
the piers are built. The width of the arrestor islands provide sufficient river width for
navigation and have a clearance to the face of the pier columns to prevent the bow of
an Oriana Class passenger liner from piercing the pier columns. The length of the
islands was determined from the recommended best practice indicated in the
AASHTO Guide Specification for Vessel Collision Design (Ref 3) that indicates that
the use of arrestor islands has been proven as the safest and most effective means
of providing pier protection (refer to Figures 5 and 7). The extent of the arrestor
island is different at each pier, in recognition of the different levels of risk presented
by the different pier locations relative to the shipping channel

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In addition, the pile caps and pile groups at the main river piers have been designed
to resist an ultimate lateral impact force of 20,000kN applied either perpendicular or
parallel to the longitudinal axis of the bridge. The pile caps are finished with precast
panels to provide a clean finish and to ensure the pile cap and piles are not exposed
at low water levels. These panels also provide resistance to small impacts from
pleasure vessels and provide some energy absorption characteristics by way of
crushing, should the pier experience an impact from a medium sized vessel of low
draft that may be able to pass over the top of the arrestor island at high tide.
Existing Bridge

New Bridge

Figure 5 Cross Section through Pier 6 Ship Arrestor Island

5.0

DEVELOPMENT OF CONCEPTS

Within the constraints of the PSTR, a number of options were considered in the
development of the final tender design for the main spans. These included the
following:
Variation in lengths for the spans adjacent to the main span. The Existing
Gateway Bridge has span lengths of 145m, 260m and 145m for the three main
spans, with 88m transition spans between the three main spans and the regular
71m approach spans.
The extension of the 71m span lengths (with
corresponding increase in side spans to 162m) best suited the construction
methods proposed;
Use of a single cell box girder. Consideration was given to retaining the same
width of bottom flange as the Existing Gateway Bridge (12m) and in adopting a
wider 15m bottom flange. For the overall width of the new bridge required, it was
determined that a two cell box girder was more economic.
Vertical and inclined webs. Preliminary designs indicated that the use of a two
cell box girder with either inclined or vertical webs resulted in similar construction
costs. It was considered that vertical webs provided a structure of closer
appearance to the Existing Gateway Bridge and simplified the design of the form
travellers.
Pile size. Preliminary designs for pile groups comprising 1.5m, 1.8m, 2.0m and
2.5m diameters were carried out. A linked pile caisson option was also
considered using similar technology to offshore oil platform construction. The
1.8m diameter solution was determined to be the most cost effective solution.

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6.0

MAIN PIER FOUNDATIONS

6.1

Pile Design

Stress resultants to the main river pilegroups under in-service and construction
loadings were determined using 3D global analysis models of the bridge structure. A
CLAP analysis (derivative of DEFPIG) was carried out to consider pile group effects
and determine design axial loads and bending moments in the individual piles. Pile
axial capacity was assessed using the method of Rowe and Armitage, as required by
the PSTR, which is specifically intended for socketed piles in relatively weak rock.
Conservative design parameters were adopted including the use of a geotechnical
strength reduction factor of 0.45, a further reduction factor for end bearing of 0.5 and
limiting the vertical displacement to 25mm. Maximum ultimate design axial load S*
on the 1.8m diameter piles was 35.2MN and all piles were designed with a factor of
safety on working loads in excess of 3.0. During construction, bore logs were taken
at each individual pile, two large scale test piles were used and a load test of 1.2S*
was applied to one pile in each pile group to confirm the pile axial capacity.
6.2

Pilecap Design

The two main river pilecaps were designed using strut and tie analysis methods. The
multiple layers of bottom reinforcement were detailed using bundles of three bars to
minimise potential congestion from the vertical pile reinforcement, which was also
detailed using a combination of 75mm diameter stressbars and conventional
reinforcement (refer to Figure 6).
To alleviate adverse thermal differentials and high peak temperatures during the
casting and hydration of the large pile caps, a system of cooling pipes was installed.
Ambient air was pumped through these pipes at relatively high velocity to extract
heat from the pilecap core. The pile cap was instrumented to show that this system
effectively maintained temperature differentials to 25C, and peak temperatures of
81C were reached at isolated points in the concrete mass.

Figure 6 Part Cross Section Through River Pier Pilecaps

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Figure 7 View of Pier 6 Showing Construction Platform (Future Ship Arrestor), Pile Cap and
Column with Pier Head Complete, Ready for Balanced Cantilevering

7.0

SUPERSTRUCTURE

The twin cell box girder has a 15m wide constant-width bottom flange. Bottom flange
thickness in the main spans varies parabolically from 1040mm at the main pier faces
to 300mm at mid-span of the main span and 375mm at the side span closure
segments. The thickness of the three webs steps from 500mm wide at the main
piers to 400mm at the ends of the cantilevers in steps of 50mm. A stability check
was carried out on the tall thin webs.
At the two main piers, two-cell pier-boxes integrate the superstructure with the piers.
Each pier-box comprises 2.5m thick transverse diaphragms which are extensions of
the twin blade pier columns. In the longitudinal direction, each of the webs has been
increased in thickness to 1.0m wide. This arrangement of webs and transverse
diaphragms provides a direct transfer of girder shear forces into the substructure.
Access openings 2.1m high by 2m wide are provided through the transverse
diaphragms in each box girder cell.
A partial prestress design was used for the main span superstructure. The in-service
stress limit cases were:
Maximum concrete compressive stress of 0.4fc under permanent effects and
0.5fc under peak serviceability limit state load cases;

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Zero tension under permanent loads and half live load (as per AS5100 for B2
Exposure Classification)
The typical distribution of permanent stresses in the superstructure are indicated in
Figure 8. for the long term case following all losses and creep redistribution.

Figure 8 Permanent Dead Load Stresses in the Superstructure

The bottom flange thicknesses were determined from consideration of serviceability

stress limits. The web thicknesses were determined from consideration of combined
shear, torsion and transverse bending effects. The top slab dimensions were
dictated by the requirement to accommodate the large number of top cantilever
prestressing tendons and to achieve the required transverse bending strength of the
outer cantilevers.
Segment lengths were initially set at 3.0m long for the first 10 segments out from the
main piers, then 10 segments of length 4.0m and 10 segments of length 5.0m. This
segmentation was chosen to limit the wet concrete weight being carried by the form
travellers and the out-of-balance moment carried by the piers. At the end of the land
side cantilevers, the twin cell box girder transitions to the twin box girders of the
approach spans within the closure segment, which incorporates a transverse
diaphragm. To achieve a balanced condition at the end of cantilever construction of
the approach spans about the transition pier, the last 5m cantilevered segment in the
main spans on the land side is reduced in length to 2.75m. The resulting imbalance
of the main cantilevers was restored by increasing the web and bottom slab
thicknesses of the land cantilever over the last ten segments.
Longitudinal prestressing tendons in the top slab each comprise 17 and 19/15.2mm
diameter extra high tensile strands. Typically at each segment end, six tendons are
anchored in face anchors adjacent to the web. The tendons are located in three
layers, symmetrically positioned about each web. There are 60 cable ducts per web
with two as spares. Bottom flange longitudinal tendons each comprise 19/15.2mm
diameter extra high tensile strands, which are anchored in internal blisters located

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adjacent to the webs.

The plan layouts of the tendons were detailed to
accommodate the penetrations in top and bottom flanges which are required to
support the form traveller. Transversely, the top slab is post-tensioned using
5/15.2mm diameter extra high tensile strands spaced at 1.0m centres along the main
spans. The transverse prestress is provided to balance permanent loads and control
cracking under traffic loading.

Figure 9 Typical Box Girder Top Flange Details

Vertical post-tensioning of the webs has been restricted to areas of the webs
adjacent to the bottom flange prestressing anchorages, to control potential cracking
in the webs. At these locations, 32mm diameter stressbars on the web centreline at
600mm centres have been adopted.
The pier boxes are post-tensioned longitudinally at regular intervals down the webs,
vertically in the webs and diaphragms and transversely in the bottom slab.
The design for shear and torsion involved recourse to the AASHTO LRFD code
because the Australian code does not deal well with this issue. The AASHTO LRFD
Ed 3 rules for combined shear and torsion at the ultimate limit state were used with a
resistance factor of
=0.7. A review of a number of international codes was
undertaken to confirm this was an appropriate design approach. In order to ensure
satisfactory performance at service loads a principal tensile stress limit of 0.289 fc at
a serviceability load case of permanent loads plus half live load was adopted.
The transverse reinforcement was detailed as slices at 200mm centres to facilitate
prefabrication of reinforcement cages. The longitudinal reinforcement in the top
flange was also designed to allow launching of the form traveller prior to stressing the
cantilever tendons for the newly cast segment.
8.0

SUMMARY

The Second Gateway Bridge is a prestressed concrete bridge with the three main
river spans of the same structural form and of similar appearance to the Existing
Gateway Bridge. The design philosophy has been to closely match the geometry of
the existing bridge to achieve a similar visual appearance, to satisfy the design
requirements of the PSTR, to incorporate current worlds best practice in the

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technical design and methods of construction and to provide a robust, durable

structure appropriate to the bridges significance in the Brisbane landscape.
ACKNOWLEDGEMENT
The authors wish to thank Queensland Motorways Limited, Leighton Abigroup Joint
Venture, and the Maunsell-SMEC Joint Venture for permission to publish this paper.
The views expressed in this paper are those of the authors
REFERENCES
1.
Standards Australia, Australian Standard Bridge Design Code (AS5100).
Standards Australia International, Sydney (2004)
2.
AASHTO, LRFD Bridge Design Specifications, Edition 3, American
Association of State Highway and Transportation Officials, Washington DC, (2006).
3.
AASHTO, Guide Specification and Commentary for Vessel Collision Design of
Highway Bridges, American Association of State Highway and Transportation
Officials, Washington DC, (1991).

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