FIBRE LASER WELDING OF

DISSIMILAR MATERIALS

A thesis submitted to The University of Manchester for the degree of

PhD
in the Faculty of Engineering and Physical Sciences

2010

HUI-CHI CHEN

SCHOOL OF MECHANICAL, AEROSPACE AND CIVIL ENGINEERING

 Table of contents

Table of contents

Table of contents 2

List of figures 8

List of tables 19

Abstract 21

Declaration 22

Copyright statement 23

Lists of publications 24

Acknowledgements 25

Nomenclature 26

Acronyms 28

Chapter 1 Introduction 30
1.1 Overview 30
1.2 Major challenges and objectives of this work 31
1.3 Thesis structure 32

Chapter 2 A literature review on welding technology 35

2.1 Introduction 35
2.2 Fusion welding 37
2.2.1 Tungsten inert-gas welding 38
2.2.2 Electron-beam welding 39
2.2.3 Laser welding 40

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2.2.4 Comparison between fusion welding technologies 41

2.3 Brazing and soldering 42
2.4 Solid state welding 43
2.4.1 Resistance welding 43
2.4.2 Diffusion welding 44
2.4.3 Friction stir welding 45
2.5 Laser welding mechanisms 46
2.5.1 Conduction mode welding 48
2.5.2 Keyhole mode welding 48
2.5.3 Plasma 50
2.5.4 Energy transfer efficiency 51
2.6 Laser welding parameters 53
2.6.1 Power 53
2.6.2 Welding speed 55
2.6.3 Beam quality 56
2.6.4 Shielding gas 58
2.6.5 Material properties 59
2.7 Laser configurations 59
2.7.1 Butt weld 60
2.7.2 Lap weld 61
2.8 Metallurgical aspects of laser welding 62
2.8.1 The fusion zone (FZ) 62
2.8.2 The heat affected zone (HAZ) 63
2.8.3 Thermal cycle 63
2.8.4 Inhomogenities 64
2.9 Laser weld defects 64
2.9.1 Cracking 65
2.9.2 Spatter 65
2.9.3 Porosity 67
2.10 Residual stresses generated in laser welding 69

Chapter 3 A literature review on welding of dissimilar

materials 71
3.1 Introduction 71

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3.2 The state-of-the-art in welding dissimilar materials 72

3.2.1 Friction stir welding of dissimilar materials 72
3.2.2 Laser welding of dissimilar materials 73
3.2.3 Comparison between LBW and FSW 75
3.3 Mechanisms in laser welding of dissimilar materials 76
3.3.1 Melting 77
3.3.2 Mixing 77
3.3.3 Solidification 77
3.4 Technology for laser welding of dissimilar materials 79
3.5 Summary 80

Chapter 4 Materials, equipments and experimental

procedures 81
4.1 Introduction 81
4.2 Materials 81
4.2.1 Titanium alloy 82
4.2.2 Nickel alloy 84
4.2.3 Zn-coated steel 85
4.2.4 Aluminium alloy 86
4.3 Fibre laser welding system 88
4.3.1 Introduction to fibre lasers 88
4.3.2 Set-up for fibre laser welding of dissimilar materials 90
4.4 Analytical facilities 93
4.4.1 Macrostructure and microstructure observation 93
4.4.2 Microhardness test 95
4.4.3 Shear force test 96
4.4.4 Chemical composition analysis 97
4.4.5 Corrosion test 97

Chapter 5 Fibre laser welding of Ti-6Al-4V titanium alloy

to Inconel 718 nickel alloy 99
5.1 Introduction 99
5.2 Experimental investigation 101
5.3 Experimental results 104

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5.3.1 The weld geometry 104

5.3.2 The weld defects 107
5.3.3 Hardness distribution of the weld 109
5.3.4 Micro-segregation of the weld 111
5.4 Discussion 114
5.4.1 Development of analytical model for welding dissimilar
materials 115
5.4.2 Discussion on experimental results 121
5.5 Conclusions 122

Chapter 6 Pulsed wave fibre laser welding of Zn-coated

steel to Al alloy 124
6.1 Introduction 124
6.2 Experimental procedure 128
6.3 Results 131
6.3.1 The weld morphology and beam geometry 131
6.3.2 Hardness distribution 135
6.3.3 Micro-segregation 137
6.3.4 Shear force 139
6.4 Discussion 140
6.5 Conclusions 142

Chapter 7 Continuous wave fibre laser welding of Zn-

coated steel to Al alloy 143
7.1 Introduction 143
7.2 Materials and methods 144
7.3 Results 147
7.3.1 The weld beam geometry 147
7.3.2 Weld defects 151
7.3.3 Hardness distribution 152
7.3.4 Micro-segregation and Microstructure 155
7.3.5 Shear force 159
7.4 Discussion 160
7.5 Conclusions 161

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Chapter 8 Corrosion performances in fibre laser welding

of dissimilar materials 163
8.1 Introduction 163
8.2 Experimental work 166
8.2.1 Electrochemical polarisation technique 166
8.2.2 Corrosion experiment 167
8.2.3 Results 168
8.2.3.1 Effect of the shielding gas type 169
8.2.3.2 Effect of number of welding pass 175
8.2.4 Discussion 178
8.3 Conclusions 179

Chapter 9 General discussion on the basic characteristics

of fibre laser welding of dissimilar materials 180
9.1 Introduction 180
9.2 The main factors in fibre laser welding of dissimilar materials 181

Chapter 10 Conclusions and recommendation for future

work 184
10.1 Conclusions 184
10.2 Recommendation for future work 185

References 187

Appendix 1 203

Appendix 2 208

Appendix 3 Fibre laser welding of un-coated steels 210

A3.1 Introduction 210
A3.2 Experiments 210
A3.2.1 Geometry and microstructure 211
A3.2.2 Hardness 213
A3.2.3 Micro-segregation 215

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A3.3 Discussion 216

Appendix 4 Introduction of corrosion mechanisms 217

Word count: 47,557 words.

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 List of figures

List of figures

Figure 2-1 Outline of joining processes in manufacturing [25]. 36

Figure 2-2 Basic types of joints in fusion welding [29]. 37
Figure 2-3 TIG welding of commercial pure titanium [35]. 38
Figure 2-4 Schematic illustration diagram of TIG welding equipment [27,
32]. 39
Figure 2-5 Schematic illustration of electron beam welding system [27]. 40
Figure 2-6 Relationship between energy density and heat input in fusion
welding [29]. 41
Figure 2-7 Penetration profiles of different welding methods [32]. 41
Figure 2-8 Schematic illustration of resistance welding [27]. 44
Figure 2-9 Schematic illustration of mechanisms occurred in diffusion
welding: (a) the initial stage (before welding); (b) the first stage
(the deformation of grains at interfacial boundaries; (c) the
second stage (grain boundary migration and elimination); (d)
the third stage (pore elimination in volume diffusion) [27]. 45
Figure 2-10 Diffusion welding of Fe3Al intermetallic to 18-8 austenitic
stainless steel at (a) 980 °C; (b) 1020 °C [50]. 45
Figure 2-11 Illustration of: (a) FSW; (b) the affected zones in FSW [29]. 46
Figure 2-12 Illustration of: (a) the conduction mode welding; (b) the
keyhole mode welding [58]. 47
Figure 2-13 The rapid sequences in a keyhole mode welding process [27].
(To, TMP and TBP are the room temperature, the melting and
boiling points of the welding material, respectively.) 49
Figure 2-14 Schematic illustration of deep penetration effects in laser
welding [63]. 49
Figure 2-15 The formation of plume and spatter in fibre laser welding of
304 stainless steel [67]. 50

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 List of figures

Figure 2-16 Influence of the reflectivity with different surface treatments on

steel surfaces when the laser beam wavelength was 10.6 µm
[74]. 52
Figure 2-17 Relationship between laser beam wavelength and the absorption
for different materials at 20 °C [76]. 53
Figure 2-18 Schematic of laser power delivery methods: (a) the continuous
wave mode; (b) the pulsed wave mode. 54
Figure 2-19 Relationship between the weld geometry and welding speed: (a)
at proper welding speed; (b) at fast welding speed; (c) at low
welding speed [46]. 56
Figure 2-20 Influences of the laser beam quality on minimise beam waist:
(a) with the same focal length; (b) with the same focus diameter
[86]. 56
Figure 2-21 Characteristics of the laser beam quality [86]. 57
Figure 2-22 Beam parameter products of conventional laser systems [88]. 57
Figure 2-23 Joint types used in laser welding [12]. 60
Figure 2-24 Schematic illustration of laser butt welding: (a) the initial stage;
(b) melting occurs at the point of impingement of the laser
beam; (c) a keyhole forms; (d) the keyhole and the molten area
have penetrated the welding materials; (e) the weld formed after
solidification [27]. 60
Figure 2-25 Schematic diagram of the interfacial reaction layer in laser lap
welding: (a) the liquid/liquid state; (b) the solid/liquid state; (c)
the solid/solid state [48]. 61
Figure 2-26 Relationship between microstructure and thermal cycle in laser
welding of structural steel [97]. 62
Figure 2-27 Interaction between the heat source and the base metal: (a) three
distinct regions in the weld; (b) The convective flow within the
weld pool [23]. 63
Figure 2-28 Interaction between process parameters affecting weld
solidification cracking [101]. 65
Figure 2-29 Relationship between the size of molten zone and the presence
of spatter: (a) view of the cross-sectioned weld; (b) view of the
longitudinal-sectioned weld [102]. 66

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 List of figures

Figure 2-30 Schematic illustration of the melt pool dynamic in fibre laser
welding at: (a) a lower welding speed; (b) a higher welding
speed [67]. 67
Figure 2-31 Schematic illustration of the formation of porosity at different
welding speeds in continuous wave laser welding [67]. 68
Figure 2-32 Schematic illustration of the relationship between the formation
of porosity and the weld shape in laser-spot welding of A1050
Al alloy [100]. 69
Figure 2-33 Distributions of residual stresses in the butt weld: (a)
longitudinal residual stresses (σx); (b) transverse residual
stresses (σy) [110]. 70
Figure 3-1 Welds of aluminium alloy to magnesium alloy obtained from: (a)
laser welding (Thickness of Mg alloy was 1.2 mm) [125]; (b)
FSW (Thickness of both Al alloy and Mg alloy was 2 mm)
[121]. 73
Figure 3-2 Laser welding of dissimilar materials: (a) Cu to Al alloy [126];
(b) AISI 304 stainless steel to AISI 420 stainless steel [79]; (c)
steel to Al alloy [127]. 74
Figure 3-3 Butt welding of AA6013-T6 aluminium alloy obtained from: (a)
CO2 laser welding; (b) FSW [128]. 76
Figure 3-4 Illustration of laser welding of dissimilar materials with backing
blocks [130]. 79
Figure 4-1 Allotropic crystal structures of pure Ti. 82
Figure 4-2 Microstructure of fibre laser welding of Ti-6Al-4V Ti alloy: (a)
the fusion and heat affected zones, 100X; (b) the fusion zone,
200X [146]. 84
Figure 4-3 Microstructure of Inconel 718 Ni alloy: (a) the parent material;
(b) the fusion zone obtained from the electron beam welding
process [152]. 85
Figure 4-4 Schematic illustration of a painting sample with the Zn-coated
steel in the automotive application [159]. 86
Figure 4-5 The formation of spatter on the weld surface in Nd:YAG PW
laser lap welding of DX56 Zn-coated steels [166]. 86
Figure 4-6 Industrial sectors of Al alloy in 2007 [167]. 87

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 List of figures

Figure 4-7 The formation of cracks in Nd:YAG laser welding of AA2024 Al

alloy [171]. 88
Figure 4-8 Arrangement of an Ytterbium-doped large-core fibre laser with
two pump sources. (HR: high reflectivity; HT: high
transmission) [175]. 89
Figure 4-9 Schematic illustration of a double-clad fibre laser [176]. 89
Figure 4-10 Operating costs of commercial laser systems [179]. 90
Figure 4-11 Experimental setup of fibre laser welding system. 91
Figure 4-12 Used laser head, jig and single axle stage. 92
Figure 4-13 Illustration of the custom-made jig. 92
Figure 4-14 Illustrations of welding fixture for: (a) laser butt welding; (b)
laser lap welding. 93
Figure 4-15 PolyVar – MET optical microscope. 94
Figure 4-16 Hitachi S-3400N scanning electron microscope. 94
Figure 4-17 Veeco Wyko NT1000 optical profiling system. 95
Figure 4-18 MICROMET 5114 microhardness test machine. 96
Figure 4-19 INSTRON 4507 universal testing machine. 96
Figure 4-20 Zeiss EVO 50 scanning electron microscope equipped with an
Oxford instruments EDS detector. 97
Figure 4-21 Experimental setup of electrochemical polarisation corrosion
test. 98
Figure 4-22 Schematic illustration of the corrosion test sample. 98
Figure 5-1 Ti-Ni equilibrium diagram [194]. 100
Figure 5-2 Thermal properties of Ti-6Al-4V and Inconel 718 [202-204]. (Cp
and k mean specific heat and thermal conductivity, respectively.)
102
Figure 5-3 Schematic diagram of the weld dimension and hardness tests. 104
Figure 5-4 Macrostructure of the cross sectional welds when laser power
and welding speed were kept at 800 W and 100 mm/s,
respectively, for the different laser beam offset positions : (a) on
the bond; (b) offset 35 μm on the Ti-6Al-4V side; (c) offset 35
μm on the Inconel 718 side. 105

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 List of figures

Figure 5-5 The weld width of full penetration welds with: (a) different laser
powers; (b) different welding speeds at a constant laser power
of 1000 W. (The laser beam was positioned on the interface of
welding plates.) 106
Figure 5-6 The weld widths with three different laser beam offset positions
at the constant welding speed of 80 mm/s. 107
Figure 5-7 Relationship between the formation of porosity, laser power,
welding speed and the laser beam offset position. (“Centre”, “Ti
side” and “Ni side” mean the laser beam was positioned on the
interface, offset to the Ti-6Al-4V side and offset to the Inconel
718 side, respectively.) 108
Figure 5-8 Relationship between the formation of crack, laser power,
welding speed and the laser beam offset position. (“Centre”, “Ti
side” and “Ni side” mean the laser beam was positioned on the
interface, offset to the Ti-6Al-4V side and offset to the Inconel
718 side, respectively.) 108
Figure 5-9 Hardness distributions of welds in fibre laser welding of Ti-6Al-
4V to Inconel 718 at the constant welding speed of 80 mm/s. 109
Figure 5-10 Hardness distributions of welds in fibre laser welding of Ti-
6Al-4V to Inconel 718 at the constant laser power of 700 W. 110
Figure 5-11 Hardness distributions in the fibre laser welding of Ti-6Al-4V
to Inconel 718 welds with different laser beam offset positions
when laser power and welding speed were 900 W and 80 mm/s,
respectively: (a) on the interface; (b) offset to the Ti-6Al-4V
side; (c) offset to the Inconel 718 side. 110
Figure 5-12 Micro-segregation in Ti-6Al-4V/Inconel 718 weld at 800 W, 60
mm/s and with the laser beam offset to the Inconel 718 side
(Optical microscope image). 112
Figure 5-13 Micro-segregation in Ti-6Al-4V/Inconel 718 weld at 700 W, 80
mm/s and the laser beam was offset to the Ti-6Al-4V side
(SEM-Backscattered electron image). 113
Figure 5-14 Schematic diagram of the melt pool calculated according to
Rosenthal’s equation with characteristic dimensions shown: (a)

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 List of figures

the Ti-6Al-4V side; (b) the Inconel 718 side; (c) the melt pool
curve in laser dissimilar materials welding. 116
Figure 5-15 The melt pool curves in fibre laser welding of Ti-6Al-4V to
Inconel 718 obtained at three different laser beam offset
positions with: (a) 1000 W and 80 mm/s; (b) 1000 W and 100
mm/s. 119
Figure 5-16 Relationships between the formation of crack, the melt pool
area, melt ratio and cooling rate in analytical modelling. 121
Figure 6-1 Relationship between the vehicle weight and fuel efficiency
[212]. 125
Figure 6-2 The use of lightweight materials on car bodies in: (a) 1977; (b)
2006. (Glass, rubber and ceramic materials are included in the
“Other” category.) [210]. 125
Figure 6-3 The relationship between vapour pressure of pure metals and
temperature [219]. 126
Figure 6-4 Fe-Al equilibrium diagram [194] 127
Figure 6-5 Schematic diagram of hardness test. 131
Figure 6-6 Schematic diagram of shear force testing sample. 131
Figure 6-7 Surface weld appearance of long pulsed-wave fibre laser lap
welding of Zn-coated steel on Al alloy: (a) 250 W, 70 mm/s and
6.25 Hz; (b) 300 W, 100 mm/s and 1.64 Hz. 132
Figure 6-8 Macrostructure of fibre laser welding of Zn-coated steel on Al
alloy: (a) 250 W, 50 mm/s and 6.25 Hz; (b) 250 W, 50 mm/s
and 4.76 Hz; (c) 250 W, 50 mm/s and 3.85 Hz; (d) 250 W, 50
mm/s and 3.23 Hz; (e) 250 W, 70 mm/s and 3.85 Hz; (f) 300 W,
50 mm/s and 3.85 Hz 133
Figure 6-9 The weld depth at: (a) welding speed of 50 mm/s; (b) laser power
of 300 W. 134
Figure 6-10 Microhardness distribution at three penetration levels in the
weld at 50 mm/s and 3.85 Hz with: (a) 250W; (b) 300 W. 135
Figure 6-11 Microhardness distribution at three penetration levels in the
weld at 300 W and 70 mm/s with: (a) 3.85 Hz; (b) 3.23 Hz. 137
Figure 6-12 Backscattered electron images and EDS results of the weld
produced at 250 W, 50 mm/s and 3.23 Hz: (a) microstructure

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 List of figures

(X400); (b) the magnified image of Section 1 in Figure 9(a)

(X1000); (c) the distribution of Fe element; (b) the distribution
of Al element. 138
Figure 6-13 Shear force of fibre laser lap welding of Zn-coated steel to Al
alloy. 140
Figure 6-14 Fracture surfaces of the shear force testing sample at 300 W, 70
mm/s and 3.23 Hz: (a) on the Al alloy side; (b) on the Zn-
coated steel side. 140
Figure 7-1 Schematic illustration of the hardness test points and the weld
dimension. 147
Figure 7-2 Schematic illustration of shear force testing sample. 147
Figure 7-3 The weld dimension obtained from the single pass welding at
100 mm/s with different shielding gases: (a) the weld width on
the Zn-coated steel side (WTop-Steel); (b) the weld depth on the
Al alloy side (DAl). 148
Figure 7-4 The weld width of the double pass welding with different
shielding gases: (a) Ar gas; (b) N2 gas. 149
Figure 7-5 Macrostructure of double pass welding when the first pass
welding was 650 W,100 mm/s, f.p.p. of 0 mm, and
accompanied with the second pass welding and shielding gases
were: (a) 150 W, 75 mm/s, f.p.p. of +2 mm and Ar gas; (b) 150
W, 75 mm/s, f.p.p. of +2 mm and N2 gas; (c) 200 W, 75 mm/s,
f.p.p. of +2 mm and Ar gas; (d) 200 W, 75 mm/s, f.p.p. of +2
mm and N2 gas; (e) 250 W, 75 mm/s, f.p.p. of +2 mm and Ar
gas; (f) 250 W, 75 mm/s, f.p.p. of +2 mm and N2 gas. 150
Figure 7-6 Top appearance of welds produced from: (a) single pass welding
with Ar gas; (b) single pass welding with N2 gas; (c) double
pass welding with Ar gas; (d) double pass welding with N2 gas.
(the first pass welding parameters: 650 W, 100 mm/s, f.p.p. of 0
mm; the second pass welding parameters: 200 W, 75 mm/s,
f.p.p. of +2 mm.) 151
Figure 7-7 Backscattered electron image of weld cross sections obtained for
the double pass welding with: (a) Ar gas; (b) N2 gas. (the first
pass welding parameters: 650 W, 100 mm/s, f.p.p. of 0 mm; the

14
 List of figures

second pass welding parameters: 200 W, 75 mm/s, f.p.p. of +2

mm.) 152
Figure 7-8 Hardness distribution obtained from single and double passes
welding with Ar gas (the first pass welding parameters: 600 W,
100 mm/s, f.p.p. of 0 mm; the second pass welding parameters:
250 W, 75 mm/s, f.p.p. of +2 mm): (a) single pass weld; (b)
double pass weld. 153
Figure 7-9 Hardness distribution obtained from single and double passes
welding with N2 gas (the first pass welding parameters: 600 W,
100 mm/s, f.p.p. of 0 mm; the second pass welding
parameters: 250 W, 75 mm/s, f.p.p. of +2 mm): (a) single pass
weld; (b) double pass weld. 154
Figure 7-10 Backscatter electron images and EDS mapping analysis of
double pass welding (the first pass welding parameters: 650 W,
100 mm/s, f.p.p. of 0 mm; the second pass welding parameters:
150 W, 75 mm/s, f.p.p. of +2 mm) with different shielding
gases: (a) Ar gas; (b) N2 gas. 156
Figure 7-11 Relative chemical compositions at the 1100 μm penetration
level obtained from double pass welding (the first pass welding
parameters: 600 W, 100 mm/s, f.p.p. of 0 mm; the second pass
welding parameters: 250 W, 75 mm/s, f.p.p. of +2 mm) with
different shielding gases: (a) Ar gas; (b) N2 gas. 157
Figure 7-12 Welds produced at double pass welding (the first pass welding
parameters: 600 W, 100 mm/s, f.p.p. of 0 mm, and the second
pass welding parameters: 200 W, 75 mm/s, f.p.p. of +2 mm)
with different shielding gases: (a) with Ar gas; (b) with N2 gas. 158
Figure 7-13 Shear force of single pass welds (600 W, 100 mm/s, f.p.p. of 0
mm) and double pass welds (the first pass welding parameters:
600 W, 80 mm/s, f.p.p. of 0 mm; the second pass welding
parameters: 200 W, 75 mm/s, f.p.p. of +2 mm) with different
shielding gases. 160
Figure 8-1 Schematic corrosion processes of Zn-coated steel [234]. 164

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 List of figures

Figure 8-2 Corrosion in laser welding of:(a) Zn-coated steel [16]; (b) AZ31
Mg alloy [246]. 165
Figure 8-3 Illustration of an anodic polarisation plot [183]. 167
Figure 8-4 Top surface appearance of the un-welded Zn-coated steel: (a)
uncorroded area; (b) corroded area. 169
Figure 8-5 Top surface appearances between corroded single and double
pass welds(Single pass welding parameters were 650 W, 100
mm/s, f.p.p. of 0 mm. Double pass: the second pass welding
parameters: 200 W, 75 mm/s, f.p.p. of +2 mm.): (a) the
corroded single pass weld with Ar gas; (b) the corroded single
pass weld with N2 gas; (c) the corroded double pass weld with
Ar gas; (d) the corroded double pass weld with N2 gas. 170
Figure 8-6 Polarisation curves with different shielding gases(Single pass
welding parameters were 600 W, 100 mm/s and f.p.p. of 0 mm.
Double pass welding: the second pass welding parameters: 200
W, 75 mm/s, f.p.p. of +2 mm).: (a) single pass welding; (b)
double pass welding. 171
Figure 8-7 Chemical composition profiles on different pass welds with Ar
gas(Single pass welding parameters were 600 W, 100 mm/s and
f.p.p. of 0 mm. Double pass welding:the second pass welding
parameters: 200 W, 75 mm/s, f.p.p. of +2 mm).: (a) the
uncorroded single pass weld; (b) the corroded single pass weld;
(c) the uncorroded double pass weld. 172
Figure 8-8 Chemical composition profiles on different pass welds with N2
gas (Single pass welding parameters were 600 W, 100 mm/s
and f.p.p. of 0 mm. Double pass welding: the second pass
welding parameters: 200 W, 75 mm/s, f.p.p. of +2 mm).: (a) the
uncorroded single pass weld; (b) the uncorroded double pass
weld. 174
Figure 8-9 Polarisation curves between single and double pass welding with
Ar gas. (The first pass welding parameters: 600 W, 100 mm/s,
f.p.p. of 0 mm; the second pass welding parameters: 200 W, 75
mm/s, f.p.p. of +2 mm). 175

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 List of figures

Figure 8-10 The surface roughness of corroded welds obtained from

different number of welding passes with Ar shielding gas (The
first pass welding parameters: 600 W, 75 mm/s, f.p.p. of 0 mm;
the second pass welding parameters: 200 W, 75 mm/s, f.p.p. of
+2 mm.): (a) single pass weld; (b) double pass weld. 176
Figure 8-11 The polarisation curves of single and double pass welds with N2
gas. (Single pass welding parameters were 600 W, 100 mm/s
and f.p.p. of 0 mm. Double pass welding: the second welding
pass parameters: 200 W, 75 mm/s, f.p.p. of +2 mm. Circles І
and П highlight the peak of maximum and minimum current
density, respectively.) 177
Figure 9-1 The interacting effects in fibre laser welding of dissimilar
materials. 181
Figure 9-2 Influenced factors in fibre laser welding of dissimilar materials. 182
Figure 9-3 Laser welding of Ti alloy to Ni alloy using: (a) CW single mode
fibre laser; (b) CW CO2 laser (White arrows highlighting the
fusion interfaces.) [95]. 183

Figure A2-1 Drawing of the top part of custom-mode jig. 208

Figure A2-2 Drawing of the bottom part of custom-mode jig. 209

Figure A3-1 The relationship between laser power and the weld penetration
depth in single pass welding with a constant welding speed of
100 mm/s. 212
Figure A3-2 Cross sections of single pass welds at 450 W and 100 mm/s in
single pass welding of: (a) Zn-coated steel to Al alloy; (b) un-
coated steels. 213
Figure A3-3 Hardness distributions in fibre laser double pass welding of: (a)
Zn-coated steel to Al alloy; (b) un-coated steels. (The first pass
welding parameters: 600 W, 100 mm/s, f.p.p. of 0 mm; the
second pass welding parameters: 200 W, 75 mm/s, f.p.p. of +2
mm). 214

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 List of figures

Figure A3-4 Backscattered electron image of fibre laser single pass welding
of Zn-coated steel to Al alloy with Ar gas shrouding. (650 W,
100 mm/s and f.p.p. of 0 mm.) 215
Figure A3-5 Backscattered electron image of fibre laser double pass
welding of un-coated steels with Ar gas shrouding. (The first
pass welding parameters: 600 W, 100 mm/s, f.p.p. of 0 mm; the
second pass welding parameters: 200 W, 75 mm/s, f.p.p. of +2
mm). 215

Figure A4-1 Illustration of an electrochemical process [254]. 217

Figure A4-2 Schematic of corrosion types: (a) uniform corrosion; (b) pitting
corrosion; (c) galvanic corrosion; (d) intergranular corrosion;
(e) crevice corrosion; (d) stress corrosion [286]. 219

18
 List of tables

List of tables

Table 2-1Comparisons between different joining technologies [26]. 36

Table 2-2 The comparison between LBW, EBW and TIG [45, 46]. 42
Table 2-3 Relationship between absorption and laser welding mode [67]. 51
Table 2-4 Properties of shielding gases commonly used in the laser material
processing [91]. 58
Table 3-1 Work reported in FSW of dissimilar materials. 72
Table 3-2 Studies of laser welding of dissimilar materials. 75
Table 3-3 Comparisons between LBW and FSW processes. 76
Table 3-4 Weldability of binary metal in laser welding processes [133]. 78
Table 4-1 Material properties of pure metals [24]. 81
Table 4-2 Classification of commercial pure Ti [142]. 83
Table 4-3 Groups of Al alloys [140]. 87
Table 4-4 Comparisons between conventional laser systems [174, 177, 178]. 90
Table 4-5 Characteristics of IPG YLR-1000-SM fibre laser. 91
Table 4-6 Specification of Hitachi 3400N scanning electron microscope. 94
Table 4-7 Technical specification of Veeco Wyko NT1000 optical profiling
system. 95
Table 4-8 Specification of the Zeiss EVO 50 scanning electron microscope. 97
Table 5-1 Chemical composition (wt.%) of Ti-6Al-4V and Inconel 718
[196, 197]. 101
Table 5-2 Physical properties of Ti-6Al-4V and Inconel 718 [24, 198-201]. 102
Table 5-3 Experimental matrix. 103
Table 5-4 Hardness of Points A-G in Figure 5-12 and Points H-M Figure
5-13. 113
Table 5-5 Chemical composition (wt.%) of Points A-G in Figure 5-12. 114
Table 5-6 Detailed values from the analytical modeling with different
welding parameters. 120
Table 6-1 Classifications of intermetallic compounds in the Fe-Al binary
system [220]. 127

19
 List of tables

Table 6-2 Chemical composition (wt.%) of DX54 Zn-coated steel and EN-
AW-5754 Al alloy. 129
Table 6-3 Physical properties of Fe, Al and Zn elements [24]. 129
Table 6-4 Experimental parameter matrix. 130
Table 6-5 Hardness and chemical compositions of Points A-F in Figure
6-12(b). 139
Table 7-1 Experimental matrix of the single pass welding. 145
Table 7-2 Experimental design of the double pass welding. 146
Table 7-3 Hardness and chemical compositions of Points A-I in Figure 7-12. 159
Table 8-1 Experimental matrix and electrochemical parameters. 168

Table A1-1 Laser welding of Al alloys to other materials. 203

Table A1-2 Laser welding of Cu to other materials. 205
Table A1-3 Laser welding of Mg alloy to other materials. 205
Table A1-4 Laser welding of Ni alloy to other materials. 206
Table A1-5 Laser welding of stainless steels to other materials. 206
Table A1-6 Laser welding of Ti alloy to other materials. 207
Table A1-7 Laser welding of hard metals to other materials. 207

Table A3-1 Parameters used in fibre laser lap welding of un-coated steels
and Zn-coated steel to Al alloy. 211

Table A4-1 Electrochemical properties of pure metals [208]. 218

20
 Abstract

Abstract

The University of Manchester

Candidate: Hui-Chi Chen
Degree: PhD
Thesis title: Fibre laser welding of dissimilar materials
Date: 5th July 2010

Joining technology has played an important role in manufacturing since the industrial
revolution. Welding methods are under constant development in response to real
demands. Laser welding is considered an effective joining method that can provide high
quality and cost effective results to bring economical benefits to industry. Nowadays,
fibre lasers have the capability to fill some of the roles of the CO2 and Nd:YAG lasers
in industrial welding applications because of their excellent characteristics such as
higher energy density and superior beam quality. However, up to now, few quantitative
evaluations of its performance against more traditional lasers have been conducted in
laser material processing. This thesis presents an investigation into the fibre laser
welding of dissimilar materials processes. The challenges in the welding of dissimilar
materials are mainly related to the large differences in the physical and chemical
properties of the welding materials. These differences readily cause residual stresses,
intermetallic phases and chemical composition gradients. The aim of this work is to
understand and explain mechanisms occurring in single mode fibre laser welding of
dissimilar materials. The first part of this work addressed fibre laser butt welding of Ti-
6Al-4V titanium alloy to Inconel 718 nickel alloy. Here, the weld quality was evaluated
in terms of the weld geometry, microstructures, hardness distributions and the formation
of intermetallic phases. Results showed that the offset position of the laser beam was an
important factor affecting the weld quality. Furthermore, the thermal history of the weld
was simulated using analytical modelling analysis and this was used to identify a
parameter window for crack-free welding. The second part of this work focused on fibre
laser lap welding of Zn-coated steel to Al alloy with different laser power delivery
modes (pulsed wave and continuous wave). The relationship between the weld quality
and process parameters, such as: pulse frequency, laser power, welding speed, the
shielding gas type and number of welding passes, were investigated. The mechanical
properties, metallurgical effects and corrosion performances of welds were analysed.
Results showed that the shielding gas type and the number of welding passes were key
factors in controlling the weld quality in the fibre laser welding of Zn-coated steel to Al
alloy process. Finally, the common features, characteristics and the potential of fibre
laser welding of dissimilar materials are presented.

21
 Declaration

Declaration
I hereby declare that no portion of the work referred to in the thesis has been submitted
in support of an application for another degree or qualification of this or any other
university or other institute of learning.

22
 Copyright statement

Copyright statement
I. The author of this thesis (including any appendices and/or schedules to this thesis)
owns certain copyright or related rights in it and she has given The University of
Manchester certain rights to use such Copyright, including for administrative
purposes.

II. Copies of this thesis, either in full or in extracts and whether in hard or electronic
copy, may be made only in accordance with the Copyright, Designs and Patents
Act 1988 (as amended) and regulations issued under it or, where appropriate, in
accordance with licensing agreements which the University has from time to time.
This page must form part of any such copies made.

III. The ownership of certain Copyright, patents, designs, trade marks and other
intellectual property (the “Intellectual Property”) and any reproductions of
copyright works in the thesis, for example graphs and tables (“Reproductions”),
which may be described in this thesis, may not be owned by the author and may be
owned by third parties. Such Intellectual Property and Reproductions cannot and
must not be made available for use without the prior written permission of the
owner(s) of the relevant Intellectual Property and/or Reproductions.

IV. Further information on the conditions under which disclosure, publication and
commercialisation of this thesis, the Copyright and any Intellectual Property and/or
Reproductions described in it may take place is available in the University IP
Policy (see
http://www.campus.manchester.ac.uk/medialibrary/policies/intellectual-
property.pdf), in any relevant Thesis restriction declarations deposited in the
University Library, The University Library’s regulations (see
http://www.manchester.ac.uk/library/aboutus/regulations) and in The University’s
policy on presentation of Theses.

23
 List of publications

Lists of publications

Conference papers

1. Hui-Chi Chen, Andrew J Pinkerton, Lin Li, Single mode fibre laser welding of
dissimilar aerospace alloys, in the proceeding of the 3rd pacific international
conference on application of lasers and optics 2008, PICALO 2008, Bejing,
China, Paper P107.
2. Hui-Chi Chen, Jonathan Blackburn, Andrew J Pinkerton, Lin Li, Paul Hilton, A
comparative study of single mode fibre laser and Nd:YAG laser welding of Ti-
6Al-4V, in the proceeding of the 3rd pacific international conference on
application of lasers and optics 2008, PICALO 2008, Bejing, China, Paper P107.
3. Hui-Chi Chen, Andrew J Pinkerton, Lin Li, Fibre laser welding of Zn-coated
steel on Al alloy for next generation lightweight vehicles, in the proceeding of
the 28th international congress on applications of laser and electro-optics,
ICALEO 2009, Orlando, FL, USA, Paper P104.

Journal papers

1. Hui-Chi Chen, Andrew J Pinkerton, Lin Li, Fibre laser welding of dissimilar
alloys of Ti-6Al-4V and Inconel 718 for aerospace applications, International
Journal of Advanced Manufacturing Technology (In press).
2. Hui-Chi Chen, Andrew J Pinkerton, Lin Li, Zhu Liu, Anil T Mistry, Gap-free
fibre laser welding of Zn-coated steel on Al alloy for light-weight automotive
applications, Materials & Design (Submitted in 2010).
3. Hui-Chi Chen, Andrew J Pinkerton, Zhu Liu, Lin Li, Anil T Mistry, Mechanical
and corrosion performances in fibre laser lap welding of Zn-coated steel to Al
alloy for automotive applications, (To be submitted).

24
 Acknowledgements

Acknowledgements

First of all, I would like to show my gratitude to my parents, Yao-Ching Chen and Hsiu-
Mei Lo, for their endless love since I was born. Words can barely express my
appreciation. They always remind me to compete with myself and tackle every
challenge with an optimistic mind. I am also grateful to my elder sister, Ching-Wen
Chen, for being the best friend in all my life. Whenever I was running or falling, she
always tried to be the best captain in my cheer squad.

Secondly, many thank my supervisors, Professor Lin Li and Dr Andrew Pinkerton, for
their guiding and advising during pursuing PhD. I particularly appreciate Dr Andrew
Pinkerton for his practical teaching on cultivating a correct altitude and research skills
throughout the work. Thanks are also extended to Dr Marc Schmidt, Dr David
Whitehead, Dr Wei Wang and Dr Zhu Liu for their technical advisement.

I would also like to thank all my colleagues in Laser Processing Research Centre
(LPRC), especially for Dr Juan Carlos Hernández, Dr Mohamed Sobih, Dr Gareth
Littlewood, Dr Stephen Leigh, Dr Nazanin MIR-Hosseini, Dr Sohaib Khan, Mostafa
Okasha, Ana Peña, Robert Lloyd, Mohsen Rakhes, Norhafiza Muhammad, Reza
Negarestani, Kamran Shah, Ashfaq Khan, Michael Vogel and Jonathan Blackburn. This
journey became more joyful because they always accompanied me.

I deeply appreciate Mr Arthur Summer, Mr Dave Buckland and Mr David Mortimer in

the School of Mechanical, Aerospace and Civil Engineering; Mr Stephen Blatch and Ms
Xiangli Zhong in the School of Materials for their fully technical assistance.

Finally, I would like to thank all my friends in Taiwan and Manchester. Chatting with
them was the best remedy for recovery and going ahead.

This thesis is dedicated to my great parents. I would not be able to complete this work
without their boundless love, consistent encouragement and unconditional support.

25
 Nomenclature

Nomenclature

Cp Heat capacity (J/kg ±C)

D Spot diameter of the laser beam (µm)
DAl The weld penetration depth in the Al alloy side (µm)
E Energy density
Eabsoebed Energy absorbed by the welding materials (J)
Eeff Energy transfer efficiency
Etotal Laser output energy (J)
f The pulse frequency (Hz)
k The thermal conductivity (W/m ±C)
K0 The modified Bessel function of the second kind and zero order
L1 The mean length of the separate melt pools in the forward direction
(mm)
L1Ni The melt pool length in the forward direction on the Inconel 718
side (mm)
L1Ti The melt pool length in the forward direction on the Ti-6Al-4V side
(mm)
L2 The mean length of the separate melt pools in the rear direction
(mm)
L2Ni The melt pool length in the rear direction on the Inconel 718 side
(mm)
L2Ti The melt pool length in the rear direction on the Ti-6Al-4V side
(mm)
P Laser power (W)
PAvg The average laser power (W)
PPeak The peak laser power (W)
q′ The rate of heat per unit length (W/m)
r The distance from the heat source (m)
Ra The roughness average (µm)

26
 Nomenclature

Rq The root mean square roughness (µm)

Rt Average maximum height of the profile (µm)
T Temperature (±C)
t The duration in continuous wave laser welding (s)
t1 The duration of laser “power on “ in pulsed wave laser welding (s)
T2 The duration of laser “power off” in pulsed wave laser welding (s)
•
T Cooling rate (±C/mm)

T0 The original sample temperature (20 ±C)

tNi The thickness of the Inconel 718 plate (mm)
tTi The thickness of the Ti-6Al-4V plate (mm)
v Welding speed (mm/s)
VNi The melt volume in the Inconel 718 side (mm3)
VTi The melt volume in the Ti-6Al-4V side (mm3)
VSCE Voltage of satured calomel electrode
wt.% Weight percentage (%)
WNi The melt pool width on the Inconel 718 side (mm)
WTi The melt pool width on the Ti-6Al-4V side (mm)
WAl The weld width on the Al alloy side (µm)
WTop-steel The weld width on the top surface of steel (µm)
WBottom-steel The weld width on the bottom surface of steel (µm)
θ Half of the laser beam divergence angle (mrad)
σx Longitudinal residual stresses (MPa)
σy Transverse residual stresses (MPa)
l Thermal diffusivity (m2/s)
ω0 Diameter of laser beam waist (mm)

27
 Acronyms

Acronyms

AE Auxiliary electrode
ASTM The American Society for Testing and Materials
BCC Body centre cubic
BSI Backscattered electron imaging
BPP Beam parameter product
BS British Standard
BSE Back-scattered electrons
CFD Computational fluid dynamics
CP Ti Commercial pure titanium
CTE Coefficient of thermal expansion
CW Continuous wave
EBW Electron beam welding
EDS Energy dispersive spectrometer
EG Electrogalvanized
EN European standard
FCC Face centred cubic
f.p.p. Focal point position
FSW Friction stir welding
FZ Fusion zone
GA Galvannealed
HAZ Heat affected zone
HCP Hexagonal close-packed
HPDL High power diode laser
ISO International Organisation for Standardisation
LBW Laser beam welding
GMA Gas metal arc welding
NDT Non-destructive testing
Nd:YAG laser Neodymium-doped yttrium aluminium garnet laser

28
 Acronyms

OVAT The one variable at a time

PMZ Partially melt zone
PRR Pulse repeat rate
PW Pulsed wave
RE Reference electrode
SCE Saturated calomel electrode
SE Secondary electrons
SEM Scanning electron microscope
TIG Tungsten inert gas welding
TWI The Welding Institute
WE Working electrode

29
 Chapter 1 – Introduction

Chapter 1
Introduction

1.1 Overview

Today, welding is one of the most important joining technologies in manufacture

industry. Laser welding has been widely used in the aerospace, automotive, electronic,
medical, shipbuilding and military defence industries due to excellent characteristics,
such as small fusion and heat affected zones, low distortion, short cycle time and
flexibility. For these reasons, laser welding processes have been gradually improved and
developed in the past two decades. Recent research includes developments of laser
systems [1-3], such as single- and multi-mode fibre lasers; investigations of the
interactions of processing parameters [4]; laser welding of light metals and plastic
generation materials [5-8]; improvements of the welding process by means of
combining a laser with another type of welding process at the same time [9-11] , such as
the hybrid laser – gas metal arc (GMA) welding.

Because of increasing demands from the industry, laser welding of dissimilar materials
has received more attention recently. Not only can the cost of materials be reduced, but
the production designs can become more flexible by using the technique [12]. However,
due to the differences in physical and chemical properties between welding materials,
some challenges still exist for further development in laser welding of dissimilar
materials. For instance, effectively suppressing the formation of intermetallic brittle
phases in the weld is a well-known issue.

Because of the increasing issues in environmental protection and energy saving, which
are seriously considered in most of the developed countries, materials currently used in
industry are slightly different from those used ten years ago. Light materials have been
gradually applied for transportation applications [13, 14]. For example, traditional steels

30
 Chapter 1 – Introduction

have been successfully replaced by titanium alloys; whereas aluminium alloys and
magnesium alloys have been applied in automotive and aerospace components [15-17].
Therefore, systematic studies in laser welding of similar and dissimilar materials
welding (particularly for joining light materials) have become important. In fact, those
studies are urgently required.

1.2 Major challenges and objectives of this work

Although works on laser welding have been reported widely, few studies have focused
on laser welding of dissimilar materials [18-21]. Accordingly, mechanisms and
phenomena occurring in laser welding of dissimilar materials have not been made clear.

One of the challenges in laser welding of dissimilar materials is to overcome the

different physical and chemical properties between the welding materials. Finding
optimal parameters to suppress the formation of intermetallic brittle phases in the weld
is another considerable task in laser welding of dissimilar materials. The weld quality is
mainly determined by the formation of intermetallic phases because they can affect the
susceptibility to solidification cracking in the weld. For these reasons, deeply
understanding and effectively controlling the formation of intermetallic brittle phases
become the priority in laser welding of dissimilar materials.

The aim of this research work is to understand and explain the phenomena and
mechanisms involved in single mode fibre laser welding of dissimilar materials (i.e. Ti-
6Al-4V titanium alloy to Inconel 718 nickel alloy and Zn-coated steel to Al alloy). The
key factors in controlling the weld quality are investigated. Analytical modelling is used
to verify and assist the understanding of experimental results. The objects of this
research are summarised below:

List of Objectives

• To identify the feasibility of fibre lasers welding of dissimilar materials.

• To investigate the effects of processing parameters and their interactions in the
fibre laser welding of dissimilar materials process.
• To model the thermal history in fibre laser welding of dissimilar materials.

31
 Chapter 1 – Introduction

• To clarify the influence of shielding gas type in the fibre laser keyhole welding
process.
• To validate and optimise process.
• To develop a scientific understanding in this process, and use this knowledge for
further welding applications.
• To contribute to the development of laser welding of dissimilar materials in
aerospace and automotive industries.

1.3 Thesis structure

This thesis addresses theoretical and experimental investigations of fibre laser welding
of dissimilar materials. Ten chapters are included. These are described sequentially as
follows.

Chapter 2

A literature review related to welding technology is described in this chapter. The first
part of this chapter covers welding history and the classification of welding methods.
Welding methods, which are commonly used in industry, are individually described
here. The second part of the review addresses laser welding. It contains the phenomena,
mechanisms, and metallurgical effects developed in the laser welding processes.

Chapter 3

A review of welding of dissimilar materials is presented in this chapter. The welding of

dissimilar materials, its features and further potential of the laser welding process are
compared with the friction stir welding process. Three main mechanisms – melting,
mixing and solidification – that occur in laser welding are discussed. The methods
currently used for improving the weld quality are mentioned and summarized here.

Chapter 4

This chapter describes the materials, equipments in welding processes and analysis
methods used throughout this thesis.

32
 Chapter 1 – Introduction

Chapter 5

The work on fibre laser butt welding of Ti-6Al-4V titanium alloy to Inconel 718 nickel
alloy is presented in this chapter. Analytical modelling is also included to support the
findings from the experimental work. The relationship between the weld quality and
processing parameters, particularly with the offset position of the laser beam, is
investigated and discussed.

Chapter 6

A preliminary study on fibre laser welding of Zn-coated steel to Al alloy using the
pulsed wave mode is presented in this chapter. Phenomena occurring in this process,
such as the formation of spatter, porosity and intermetallic brittle phases, are
investigated.

Chapter 7

This chapter presents work on continuous wave fibre laser lap welding of Zn-coated
steel to Al alloy. The influence of the shielding gas type and number of welding passes
are investigated. An explanation of the physical phenomena involved is attempted after
carrying out a series of metallurgical and mechanical analyses.

Chapter 8

This chapter extends the work from chapter seven. The corrosion resistance of fibre
laser welded of Zn-coated steel and Al alloy is evaluated using the electrochemical
polarisation method. The relationship between processing parameters and the corrosion
performance of welds are discussed in terms of chemical composition, the weld
appearance, surface roughness and the polarisation curve.

Chapter 9

This chapter presents the common features and characteristics of fibre laser welding of
dissimilar materials. It also presents a list of the general phenomena that should be
observed in the fibre laser welding process for dissimilar metallic materials. The

33
 Chapter 1 – Introduction

potential benefits of fibre laser welding of dissimilar materials in comparison with the
CO2 laser welding process are highlighted.

Chapter 10

This chapter presents the conclusions reached in this research project as well as
recommendations for future work.

34
 Chapter 2 – A literature review on welding technology

Chapter 2
A literature review on welding
technology

2.1 Introduction

Welding is to a metallurgical joining process [22]. The existence of welding can be

traced to the prehistoric period, when people jointed copper-gold and lead-tin alloys. As
electricity became readily available, the development of modern welding technology
accelerated in the latter half of the 19th century. In a broad sense, the formation of
metallurgical bonds in welded, brazed and soldered joints all include in the family of
welding processes. However, methods of adhesive bonding and the mechanical
fastening are excluded from welding, as Figure 2-1 shows. Table 2-1 briefly compares
advantages and disadvantages between different joining methods. Basically, the welding
mechanisms can be divided into three types: solid-state welding, brazing/soldering and
fusion welding. Today, according to applications, each welding technology has its own
role in manufacturing. For example, adhesive bonding is used to seal materials in the
very low temperature process (< 100 °C); soldering and brazing are generally used in
electronic applications; resistance welding is broadly employed for welding nuts and
bolts; tungsten inert gas welding is usually applied for architectural fabrications.
Meanwhile, the application of electron beam welding and laser welding are diverse as it
can be used to weld elements in the micrometer scale to the metre scale in industry. In
general, the weld quality is evaluated in terms of its geometry, microstructure, the
formation of defects and the gradient of residual stresses in the weld [23, 24]. In this
chapter, selective welding methods from fusion welding, solid state welding to
brazing/soldering are briefly introduced in the first part. The second part of this chapter
explains the characteristics, the effects of processing parameters and the physical
phenomena developing in laser welding.

35
 Chapter 2 – A literature review on welding technology

Figure 2-1 Outline of joining processes in manufacturing [25].

Table 2-1Comparisons between different joining technologiesa [26].

Temperature Advantages /
Technology Application
range Disadvantages
• Low temperature
Adhesive + low cost
< 100 °C sealing
bonding - selective process
• Optoelectronics
+ low cost
• Electronic
Soldering < 200 °C + large quantity
interconnection
- long time instability
+ fast • Chip
Brazing < 600 °C - restricted material choice interconnection
- selective
+ fast • Connector
Resistance
< 1500 °C + all materials packaging
welding
- selective
Electron + fast • Joining different
beam > 1500 °C + all materials material
welding -vacuum enviroment • Medical device
• Packaging
Laser + fast • Joining different
> 1500 °C
welding + all materials material

a
‘+’ and ‘-’ mean advantages and disadvantages, respectively.

36
 Chapter 2 – A literature review on welding technology

2.2 Fusion welding

The principle of fusion welding uses a heat source, the chemical or electrical method, to
cause the heating and cooling processes resulting in the welding material to melt and
solidify. Depending on the thickness of welding materials and applications, fusion
welding can be carried out with or without a filler material. Multi passes welding is
another common method for welding a thick material which cannot readily obtain an
acceptable result from single pass welding [23, 25, 27]. Regarding the weld quality, the
metallurgical effect developed in fusion welding plays a very important role in
determining the welding results. Due to the fact that fusion welding is a non-equilibrium
process, which ininvloves rapid heating and cooling processes, microstructure
transformations resulting from a rapid solidification process can occur in fusion welding
[27, 28]. Today, fusion welding has been widely used in welding different types of
joints in industry as shown in Figure 2-2.

Figure 2-2 Basic types of joints in fusion welding [29].

In fusion welding, several parameters (including laser power and welding speed,
physical and chemical properties of the welding material, the solidification rate and the
thermal gradient) can determine the weld quality. During solidification, variations of
chemical composition in the weld result from the micro-segregation and precipitation of
intermetallic phases. Therefore, mechanical properties and corrosion resistance of the
weld are thought to be different in comparison with parent metals [30]. From the
viewpoint of metallurgy, the weld can usually be divided into the fusion zone (FZ), the
heat affected zone (HAZ) and the un-affected zone according to their microstructure in

37
 Chapter 2 – A literature review on welding technology

welded materials. A partially melt zone (PMZ) between the fusion and HAZ might be
observed in fusion welding of alloys. Besides, depending on whether a filler material is
used or not and chemical composition of the filler material, fusion welding can also be
classified into autogeneous welding, homogeneous welding and heterogeneous welding.
Brief explanations of them are listed as below [27]:

• Autogenous welding: a filler material is not applied in fusion welding.

• Homogenous welding: the chemical composition of the filler material applied is
identical with the welding material.
• Heterogenous welding: the chemical composition of the filler material applied is
different with the welding material.

2.2.1 Tungsten inert-gas welding

Tungsten inert gas welding (TIG) has been applied successfully for numerous materials,
such as aluminium alloys, magnesium alloys, titanium alloys (Figure 2-3) and refractory
metals, with a wide range of material thicknesses. An electric arc is applied as a heat
source during the TIG welding process. As Figure 2-4 shows, the arc is generated
between the end of a non-consumable tungsten electrode and the material at the joint
line. A shielding gas supplied by argon or helium usually is used to protect the melt
pool and the electrode. TIG welding can be carried out with or without a filler wire [27,
31, 32]. Because of its characteristics of high temperature and low welding speed, a
wider HAZ normally is produced [33]. Although TIG was widely used in industry for
welding of various materials, features of low welding speed and a wide HAZ are
considered as the main obstacles in developing this technology. Since some drawbacks
still exist in TIG welding, engineers have had their attention gradually drawn to the
laser welding technology [34].

Figure 2-3 TIG welding of commercial pure titanium [35].

38
 Chapter 2 – A literature review on welding technology

Figure 2-4 Schematic illustration diagram of TIG welding equipment [27, 32].

2.2.2 Electron-beam welding

Similar to laser welding, electron beam welding (EBW) with a high energy density can
easily produce a deep and narrow weld. When the electrons strike the material, as
shown in Figure 2-5, their kinetic energy transfers to heat and therefore causes the
welding material melt. Because the heat in EBW is obtained from the kinetic energy of
the electrons, a high vacuum environment (10-3 to 10-5 atm) is normally required in
order to avoid the influence of air or gases that affect the electrons flow to the welding
materials during the process. The higher vacuum the welding environment has, a deeper
and narrower weld can effectively be produced. Not only a wide range of material
thicknesses, from a foil to a plate, can generally be welded by EBW, but the shielding
gas and a filler material are not required in EBW. In short, EBW can produce a weld
with a near parallel-sided, deep and narrow, small HAZ, low distortion and less

39
 Chapter 2 – A literature review on welding technology

shrinkage. For these reasons, EBW has been widely utilised in the manufacture of
aircraft, missiles, nuclear and electric power plants [25, 32].

Figure 2-5 Schematic illustration of electron beam welding system [27].

2.2.3 Laser welding

In laser welding (laser beam welding; LBW), a laser beam is used as the heat source.
Due to its high energy density, which is similar to electron beam welding, it has been
widely used in industry for high precision and high quality welds, such as the aerospace,
automobile, microelectronics and medical industries [36]. Laser welding has been
reported to have many benefits, such as high energy density, high welding speed, high
precision, reliability, high efficiency and productivity as well as short cycle time [37-
40]. In comparison with other fusion welding technologies, a weld with better
metallurgical and mechanical properties can be obtained in laser welding because of its
high energy density (Figure 2-6) and the ability of high welding speed [41, 42]. LBW
currently plays an important role and offers comparable quality to electron beam
welding in industry [43].

40
 Chapter 2 – A literature review on welding technology

Figure 2-6 Relationship between energy density and heat input in fusion welding [29].

2.2.4 Comparison between fusion welding technologies

When comparing welds made by TIG, EBW and LBW separately, laser beam welding
can produce a narrow and deep penetration weld; electron beam welding can protect the
weld from any contamination by gases since it operates in a vacuum environment;
tungsten inert-gas welding has the largest fusion and HAZs, with high temperature and
low welding speed [44]. The comparative melt areas are schematically illustrated and
compared in Figure 2-7. Table 2-2 outlines the detailed comparisons between these
three types of fusion welding methods.

Figure 2-7 Penetration profiles of different welding methods [32].

41
 Chapter 2 – A literature review on welding technology

Table 2-2 The comparison between LBW, EBW and TIG [45, 46]b.
LBW EBW TIG
Joining efficiency -
Low heat input -
High aspect ratio + + -
Small heat affected zone + + -
High processing speed + + -
Weld bead appearance + + -
Weld at atmospheric pressure + - +
Weld reflective materials - + +
Weld magnetic materials -
Weld heat sensitive materials -
Combine with filler - +
Automate process + - +
Equipment cost - - +
Operating cost +
Reliability + - +

2.3 Brazing and soldering

The principle of brazing and soldering is to use a molten filler with the metallurgy of
binary or ternary eutectic systems to wet the mating surface between the joining
materials of lower melting temperature [24]. During this process, metallurgical bonds
form between the filler material and the joining material [47]. Theoretically, the
difference between brazing (above 450°C) and soldering (below 450°C) is their
operating temperature. Due to the different operating temperature, the strength of the
brazing joint is higher than that of the soldering joint [25]. With regard to the
metallurgy, the bonding formed between adjacent surfaces in brazing or soldering is
caused by the action of the capillary. The diffusion rate of the interfacial reaction layer
between joining materials is strongly dependent on the melting time and the melting
temperature. In order to increase the heat transferred between joining components to

b
‘+’ and ‘-’ mean advantages and disadvantages, respectively.

42
 Chapter 2 – A literature review on welding technology

modify or remove the surface oxide and contaminations, a flux is used during the
soldering or brazing process [24].

Because of the low operating temperature, soldering is widely used to join electronic
components today [25]. In comparison with the adhesive bonding process, advantages
of brazing and soldering are their permanent joint, good electrical and thermal
conductivity, and better corrosion resistance in organic solvents [24].

2.4 Solid state welding

The main difference between solid state welding and fusion welding is the absence of
melting or liquid phases in the solid state welding process. The principle of solid state
welding is to place two clean surfaces into atomic contact under sufficient pressure and,
occasionally, apply heat or some movement into the mating surface by plastic
deformations to enhance the joining strength [25]. Comparing with other joining
technologies, solid state welding is especially suggested as one suitable method to join
dissimilar metals [48]. Three solid state welding technologies, resistance welding,
diffusion welding and friction stir welding, are briefly introduced in the following
sections.

2.4.1 Resistance welding

The principle of resistance welding is based on the resistance to flow of electrical

current, which produces the required heat for welding as shown in Figure 2-8. In
resistance welding, the main parameters are welding current, welding time, electrode
force and electrode shape. Other assisting equipments, such as consumable electrodes,
the shielding gas or flux are not necessary in resistance welding. In order to improve the
weld quality, the welding materials have to be compressed during welding. Due to the
higher compressive pressure applied in resistance welding, the formation of cracking
and oxidation of the weld can be avoided. Currently, resistance lap welding with
different thicknesses is still widely used in the automotive industry [27, 33, 49].

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 Chapter 2 – A literature review on welding technology

Figure 2-8 Schematic illustration of resistance welding [27].

2.4.2 Diffusion welding

The principle of diffusion welding is applying a pressure at the elevated temperature to

generate a weld between two materials. The main processes occurred in diffusion
welding are illustrated in Figure 2-9. The weld quality is closely related to the applied
pressure and temperature, duration and the surface condition. In diffusion welding, the
applied pressure must be high enough, usually near the yield strength of joining
materials, to cause plastic deformations of the faying surface. The total diffusion time is
strongly dependent on the welding temperature. Diffusion welding does not show
macroscopic deformations in the weld neither the relative motion between welding
materials. Furthermore, the weld has similar physical and mechanical properties with
the parent materials. The weld quality can be improved with increasing the diffusion
rate between the welding materials. Here, the welding temperature is usually required to
be higher than half of the melting point of the welding materials. Figure 2-10 compares
results produced at different processing temperatures in diffusion welding of Fe3Al
intermetallic to 18-8 austentic stainless steel [50]. Generally, a filler material can also be
used in diffusion welding. As the processing mechanism is related to the migration of
the atoms across welding materials, on the other hand, the processing time is usually

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longer than other conventional welding processes. Due to its superior performance in
welding ceramics and ceramics-to-metal without a HAZ, diffusion welding is
particularly suitable for welding of dissimilar materials in industry [24, 25, 27, 47].

Figure 2-9 Schematic illustration of mechanisms occurred in diffusion welding: (a) the
initial stage (before welding); (b) the first stage (the deformation of grains at interfacial
boundaries; (c) the second stage (grain boundary migration and elimination); (d) the
third stage (pore elimination in volume diffusion) [27].

Figure 2-10 Diffusion welding of Fe3Al intermetallic to 18-8 austenitic stainless steel at
(a) 980 °C; (b) 1020 °C [50].

2.4.3 Friction stir welding

In 1991, friction stir welding (FSW) was invented at The Welding Institute (TWI) in
England as an adaption of friction welding. In FSW, a cylindrical tool is rotated and
squeezed between two abutting materials, as shown in Figure 2-11(a). The friction heat

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which decreases the material resistance to plastic deformation can cause the softening of
materials behind the rotated tool. A solid-state weld is formed after the materials are
stirred. The influencing parameters in FSW are the depth of the rotating tool plunge into
joint, the rotational and translational speeds of the rotating tool, and the squeezing
pressure [27]. Since the maximum temperature in FSW is normally lower than the
melting point of welding materials, it is more suitable to weld non-heat-treatable
materials, such as aluminium alloys [51]. It is also able to weld materials which have a
wide range of mechanical and metallurgical properties. In this process, the heat input is
less, but the affected zone of FSW is wider than other fusion welding methods, such as
electron beam welding and laser welding [52]. Figure 2-11(b) illustrates the affected
zones obtained in FSW. It can be divided into the dynamically recrystallised zone,
thermomechanically affected zone and thermally affected zone. In the dynamically
recrystallised zone, new and smaller grains are produced to replace the original grain
structure. Meanwhile, severely twisted grains can be found in the thermomechanically
affected zone [29].

Figure 2-11 Illustration of: (a) FSW; (b) the affected zones in FSW [29].

2.5 Laser welding mechanisms

Due to the fact that laser welding is a non-equilibrium process, the mechanism of rapid
solidification can produce a weld with the finer microstructure and increased hardness
[53]. According to the weld geometry, laser beam welding can be divided into the
conduction mode (Figure 2-12(a)) and the keyhole mode (Figure 2-12(b)) [54]. In more
detail, the two modes can be distinguished according to their ratios of the weld depth to

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width [55]. On one hand, if the ratio is about 0.5 and the weld geometry is similar to a
semi-circle, then it can be classified as the conduction mode welding. If the ratio is
higher than 0.5 and the weld geometry is a deep penetration, on the other hand, it can be
called keyhole mode welding. To summarise, the clear difference between the
conduction mode and the keyhole mode is their penetration depths [56]. The welding
mode can be determined by the intensity distribution of the laser beam and thermo-
mechanical properties of welding materials [56]. Another difference between the
conduction and keyhole modes is their melting pool surfaces during welding. For the
conduction mode welding, its melting pool surface cannot be broken. In the keyhole
mode welding, the melt pool surface can be broken and produce a keyhole to allow the
laser beam to radiate in the melt pool [45]. Furthermore, the welding mode can be
classified according to the laser beam energy density. If the energy density is below 106
W/cm2, the welding process is classified as conduction mode welding while the energy
density of the keyhole mode welding is usually between 106 to 5×107 W/cm2 [12, 57].

Figure 2-12 Illustration of: (a) the conduction mode welding; (b) the keyhole mode
welding [58].

In the conduction mode welding, the weld will be wider and more heat will be
conducted to the surrounding material. It can be said that the heat at the irradiated area
can diffuse into the lower temperature areas. However, the material melt can only occur
at the irradiated region in the keyhole mode welding. When the amount of laser energy
is over the required for melting the material, the temperature within the melt pool can

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exceed its boiling point and therefore induce local vaporisation of molten metal. The
high pressure caused from the vaporisation can promote turbulences in the molten metal
and form a deep vapour cavity which is called keyhole. When a keyhole is formed, the
laser energy can be transferred into the welding materials more rapidly and deeply [12,
57].

2.5.1 Conduction mode welding

When the laser beam is defocused or the heat input is very low, conduction mode
welding can occur, which yields a semi-spherical shape of the weld, lower penetration,
higher heat losses and poor efficiency in welding [59]. The peak temperature of
conduction mode welding is generally lower than the boiling point of the welding
materials. In this case, heat is conducted from the surface [54]. The stable melt pools
and defect-free welds easily obtained with the conduction weld mode makes it a suitable
method for welding thin materials. These are commonly used in the aerospace,
automotive and electronics industries as well as in the manufacture of medical devices
[56].

2.5.2 Keyhole mode welding

In keyhole mode welding, the weld geometry is related to the focal spot size of the laser
beam and the laser power density. Characteristics of the keyhole mode welding are the
high energy density, low heat input, high penetration, high accuracy and high welding
speed, thus, most applications carried out by laser welding are keyhole mode welding
[60]. The weld can be divided into two parts in keyhole mode welding. One is a semi-
circular shape at the top of the weld. Another part is a near parallel-sided area, with a
gradual narrowing from the top surface to the bottom [61]. The full penetration welding
is one type of the keyhole welding and has been widely used in industry [62]. Figure
2-13 illustrates the sequences developing in a keyhole welding process.

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Figure 2-13 The rapid sequences in a keyhole mode welding process [27]. (To, TMP and
TBP are the room temperature, the melting and boiling points of the welding material,
respectively.)

During the keyhole mode welding, the surface of the welding material can be heated
beyond its boiling point and cause the vaporisation above and within the keyhole, as
shown in Figure 2-14. Meanwhile, the recoil pressure from vaporisation can depress the
surface of the molten material to generate a keyhole. The depth of the keyhole can be
increased by the multiple reflections of the laser beam within the keyhole. In general,
the efficiency of keyhole mode welding is higher than conduction mode welding since
most of the incident laser beam is reflected in the keyhole and the thermal conduction
losses are significantly reduced [54].

Figure 2-14 Schematic illustration of deep penetration effects in laser welding [63].

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2.5.3 Plasma

Plasma formation is one of the by-products in laser welding, as shown in Figure 2-15.
The existence of plasma helps with and accompanies the formation of a keyhole [64]. It
is produced from the ionization of metal vapour and the gas surrounding the melt area
[65]. The temperature of plasma can reach up to 20,000 °C [66]. During the welding
process, the plasma moves along the heat source through the melt pool, which increases
its width at the top surface of the weld.

Figure 2-15 The formation of plume and spatter in fibre laser welding of 304 stainless
steel [67].

Basically, the plasma developing in laser keyhole welding can be divided into the
external plasma and the keyhole plasma. The external plasma is produced when
vaporisations are induced from the molten metal of welding materials. The amount of
ionized chemical species coming from the metal vapours or the shielding gas can be
increased under continuous irradiated condition. Vaporised materials own similar
characteristics with the plasma being ejected. They move opposite and perpendicular to
the top surface of the welding materials, and towards the laser beam. The keyhole
plasma is in the form of an opaque plume. It can be found inside the keyhole and
contributes to remove vaporised materials. The difference between the external plasma
and the keyhole plasma in laser keyhole welding can be described as follows: the
external plasma reflects and defocuses the laser beam, which reduces the process
efficiency. On the other side, the keyhole plasma helps to transfer the laser energy into
the welding materials [68]. Helium was found to effectively reduce the formation of
plasma in laser welding as well as its high ionisation potential [69].

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2.5.4 Energy transfer efficiency

Energy transfer efficiency in laser welding is closely related to the absorption and
reflection of the laser beam on the melting surface of welding materials. The definition
of energy transfer efficiency is shown in Eq.(2-1) [70]:

E absorbed
E eff = (2-1)
Etotal

Where Eeff, Eabsoebed and Etotal are the energy transfer efficiency, the energy absorbed by
the welding materials and the total laser output energy.

In laser welding, several factors can affect the absorption of laser beam by the welding
materials. They include the top surface appearance (roughness, coating, surface
preparation, etc.), the chemical composition of welding materials, the type of shielding
gas, the formation of plasma, the joint geometry and the laser beam wavelength [70, 71].
Several methods are used to improve the absorption of energy in laser welding. For
instance, the use of the appropriate shielding gas (including type of gas and its flow rate)
can effectively avoid the strong plasma shielding effect which would reduce the
absorption of energy in the laser welding process [72]. Furthermore, the absorption of
energy can be significantly increased once the keyhole has been formed due to the
occurrence of multiple reflections within the keyhole, as compared in Table 2-3 [29, 70].
The energy absorption can also be increased by pre-heating the welding materials [71],
through the application of an absorbent coating layer on the surface of the welding
material and by increasing the surface roughness of the welding material [45].

Table 2-3 Relationship between absorption and laser welding mode [67].

Type of laser system Absorption of the laser beam

Conduction mode welding Keyhole mode welding
Nd:YAG laser 30 % 70%
Fibre laser 65% 85%

The reflectivity, surface roughness, chemical composition and temperature of the

welding material are factors which influence the laser welding process [73].
Theoretically, the reflectivity can be reduced by increasing the temperature of the

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welding material or the energy of the laser beam [72]. Depending on the material phase,
however, the reflectivity can be decreased from 95% in the solid state to 85% at the
liquid state in laser welding of steel [73]. The reflectivity becomes 60% when it is at the
vapour state. When the laser energy density is increased from 105 W/cm2 to 106 W/cm2,
the reflectivity in the solid, liquid and vapour states is reduced from 95% to 55%, 85%
to 52% and 70% to 45%, respectively. Figure 2-16 compares the reflectivity among
different surface treatments in laser welding of steel.

Figure 2-16 Influence of the reflectivity with different surface treatments on steel
surfaces when the laser beam wavelength was 10.6 µm [74].

The laser beam wavelength, which is another parameter characterising the laser beam
quality, also can influence the absorption. The shorter the wavelength of the laser beam,
the more energy theoretically is absorbed by materials [59, 75]. Figure 2-17 shows the
relationship between the laser beam wavelength and the absorption of different
materials. It can be said that using a laser beam with a shorter wavelength can produce a
relatively lower reflectivity from welding materials and increase the absorption at the
material surface [46].

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Figure 2-17 Relationship between laser beam wavelength and the absorption for
different materials at 20 °C [76].

2.6 Laser welding parameters

Generally speaking, processing parameters in laser welding include laser power,

welding speed, laser beam focal spot size, types of shielding gases, and welding
materials’ properties. These parameters can determine the weld quality, such as the weld
geometry, microstructure in the fusion and HAZs, the formation of defects and so on
[59, 77]. The appropriate selection of processing parameters and the understanding of
their interactions is a direct and effective way to obtain favourable results in laser
welding [53].

2.6.1 Power

The laser power delivery method can be classified into two modes: continuous wave
(CW) mode and pulsed wave (PW) mode, as shown in Figure 2-18(a) and Figure
2-18(b), respectively. Different results can be obtained between these two delivery
methods. In the CW laser welding process, the laser power should be increased when
the welding speed is increased for a given thickness of welding materials. It can be said
that a high welding speed can be achieved by operating with a high laser power,
especially for a laser system with high energy density [59]. The PW laser welding is

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characterised by applying intermittent laser beam powers, which would cause melting
and solidification at the same time in the welding process [78]. Due to the benefit of
precisely controlling the heating and cooling rates, the PW laser welding is commonly
used in welding heat-sensitive materials where a precise heat input are required [79, 80]
with low welding speed [81].

Figure 2-18 Schematic of laser power delivery methods: (a) the continuous wave mode;
(b) the pulsed wave mode.

In continuous wave laser welding, see Figure 2-18(a), laser power is kept as a constant
over the whole welding duration. Energy of continuous wave laser welding can be
defined as Eq.(2-2).

E = P×t (2-2)

Where E, P and t are laser energy (J), laser power (W) and the welding duration (s),
respectively.

In pulsed wave laser welding, the weld quality is influenced by the peak laser power
(PPeak), welding speed, laser frequency and laser beam diameter. Definitions of pulse
energy (E), pulse frequency (f) and average power (PAvg) are shown in Eq.(2-3) to
Eq.(2-5), respectively.

Pulse energy:
E = Ppeak × t1 (2-3)

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Duty cycle:
t1
f = (2-4)
t1 + t 2

Where t1 and t2 are the duration of laser beam pulse on and off, respectively.

Average power:
PAvg = E × f (2-5)

In pulsed wave laser welding, processing parameters, such as welding speed, pulse
frequency, the laser beam spot diameter and the pulse-on duration, can determine the
heating and cooling rates, the formation of spatter and cracks [78, 80, 82].

2.6.2 Welding speed

Laser power and welding speed are two main parameters which affect the weld
geometry including its depth and width. For example, a slower welding speed can cause
the melt pool to become wider and larger when other processing parameters are kept
constant. As Figure 2-19(c) shows, a weld with drop out is usually produced at slower
welding speed. A higher welding speed can induce a strong flow in the centre of the
melt pool and limit the degree of redistribution in the keyhole, as shown in Figure
2-19(b). Therefore, an undercut weld can be produced when welding speed is higher.
Additionally, welding speed affects the heat distributions in the weld [83]. Regarding to
plastic deformations, higher local deformations and lower angular distortions can be
easily caused by a lower welding speed. On the other hand, a higher welding speed can
cause lower local deformations but a higher angular distortion [63]. It can also result in
a higher welding speed can generally lead to the higher temperature gradient, the
formation of cracking and concentrated residual stresses are readily observed in the
weld at a higher welding speed [84]. In laser welding of dissimilar materials, however, a
higher welding speed could effectively restrict the degree of micro-segregation resulting
in the occurrence of intermetallic phases and degradation of mechanical properties of
welds [63].

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Figure 2-19 Relationship between the weld geometry and welding speed: (a) at proper
welding speed; (b) at fast welding speed; (c) at low welding speed [46].

2.6.3 Beam quality

Generally speaking, a higher beam quality means that a smaller focus diameter can be
obtained with the same focal length, Figure 2-20(a), and a greater standoff distance with
the same focus diameter, as shown in Figure 2-20(b). According to ISO standard
11146:2005 [85], as illustrated in Figure 2-21, the beam quality is a function of the laser
beam waist and its divergence angle. The beam quality can be presented in terms of the
beam parameter product (BPP) in Eq.(2-6).

θ
BPP = ω 0 × (mm-mrad) (2-6)
2

Where ω0, θ are the diameter of laser beam waist (mm) and the laser beam divergence
angle (mrad).

Figure 2-20 Influences of the laser beam quality on minimise beam waist: (a) with the
same focal length; (b) with the same focus diameter [86].

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Figure 2-21 Characteristics of the laser beam quality [86].

A low value of BPP means the laser beam has higher beam quality or has the ability to
focus to a small spot size. A smaller beam divergence and focal spot diameter can easily
produce a weld with higher penetration and smaller width for a given laser power and
welding speed [59]. In other words, a laser beam with a lower value of BPP has the
ability to produce a weld with the deeper penetration, due to its energy density being
increased by decreasing the laser beam spot size on the top surface of the weld materials
[87]. The beam parameter products between conventional laser systems are compared in
Figure 2-22.

Figure 2-22 Beam parameter products of conventional laser systems [88].

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2.6.4 Shielding gas

Shielding gas is another important factor which affects the weld quality in laser welding.
Shielding gas is used in the process to prevent the weld from oxidation, suppress the
formation of plasma to reduce the reflective of the laser beam [45] and protect the
optical lens from spatter. In general, the geometry of the nozzle, the physical and
chemical properties of the shielding gas and the gas flow rate were found to be factors
in determining the weld quality [46, 89]. Nowadays, helium and argon gases are
commonly used in laser welding due to their metallurgically neutral behaviour [32, 90].
Owing to the superior characteristics of high ionization potential and the capability to
breakdown the plasma, helium gas can contribute to increase the amount of heat
supplied into the welding material. Recently, argon gas is the most popular shielding
gas used in the laser welding process. Because of similar properties with helium gas,
sometimes, nitrogen gas is used in laser welding especially when the cost of shielding
gas is a serious concern [32]. Table 2-4 compares properties of different gases
commonly used in the laser material applications.

Table 2-4 Properties of shielding gases commonly used in the laser material processing
[91].
Gas Chemical Molecular Specific gravity Density Ionization
symbol weight with respect to air (g/L) potential
(g/mol) at 1 atm and 0 °C (eV)
Argon Ar 39.95 1.38 1.75 15.7
Carbon
CO2 44.01 1.53 1.98 14.4
dioxide
Helium He 4.00 0.14 0.18 24.5
Hydrogen H2 2.02 0.07 0.09 13.5
Nitrogen N2 28.01 0.97 1.25 14.5
Oxygen O2 32.00 1.11 1.43 13.2

To date, large amount of research has investigated the most appropriate shielding gases
needed in laser welding of non-ferrous alloys which are widely used in aerospace,
chemical and medical industries. However, these materials readily react with oxygen,
nitrogen and hydrogen under high temperature environments [92]. Using an appropriate

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shielding technology in the laser welding process can avoid the weld being seriously
oxidised or contaminated. For example, Caiazo et al. [93] investigated the effect of two
different shielding gases in CO2 laser welding of Ti-6Al-4V alloy. More internal defects
in the weld were obtained when argon was used as the shielding gas in comparison with
the use of helium. Moreover, deeper welds were produced when helium shielding gas
was used. Furthermore, numerical simulation results revealed that the flow rate of
shielding gas probably was not the main factor influencing the weld quality when a high
density gas was applied as the shielding gas [65]. Jorgensen [94] indicated that
comparing the use of pure argon gas, the weld depth could be increased by using a
mixture of shielding gases with 90% argon and 10% oxygen. His results also verified
that the flow rate of shielding gas was not an important factor influencing the weld
depth.

2.6.5 Material properties

Physical and chemical properties of welding materials dominate the weldability in laser
welding. Physical properties include thermal conductivity, diffusivity, reflectivity,
density, specific heat capacity, thermal expansion coefficient, the melting boiling points,
etc. [12]. For instance, thermal diffusivity and density of the liquid metal can affect the
degree of heat transfer and solute convection, respectively [95]. Regarding the
metallurgy, phenomena such as the formation of cracking, oxidation and intermetallic
phases are closely related to the chemical properties of the weld material. In laser
welding of dissimilar materials, the solubility between weld materials strongly controls
the presence of intermetallic brittle phases in the weld.

2.7 Laser configurations

In comparison with other fusion welding technologies, one extraordinary advantage of

laser welding is its flexibility. Laser welding is suitable to join pieces with different
geometries and easily produces an acceptable weld quality in industry. Figure 2-23
illustrates the joint geometry types that can be undertaken in laser welding.

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Figure 2-23 Joint types used in laser welding [12].

2.7.1 Butt weld

Laser butt welding is commonly used in the keyhole mode welding process. As Figure
2-24 shows, a keyhole is formed during the process. After the laser beam moves away,
the molten materials around the keyhole will collapse and fill the keyhole until
solidification occurs. The two weld materials can be joined completely after the welding
materials have undergone both melting and solidification phases. In fact, this type of
joint is widely used in manufacturing. The efficiency in laser butt welding relies on the
gap between weld materials. In laser butt welding of dissimilar materials, the
temperature profiles on both sides can be different according to their melting and
boiling points, thermal conductivity and so on.

Figure 2-24 Schematic illustration of laser butt welding: (a) the initial stage; (b) melting
occurs at the point of impingement of the laser beam; (c) a keyhole forms; (d) the
keyhole and the molten area have penetrated the welding materials; (e) the weld formed
after solidification [27].

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2.7.2 Lap weld

Laser lap welding is usually applied to thin plates, especially for applications in the
automotive industry. Laser lap welding can be classified into three types, i.e. the
liquid/liquid, solid/liquid and solid/solid states, as shown in Figure 2-25. Figure 2-25(a)
depicts the case of liquid/liquid state welding, where both weld materials are melted to
cause a thorough mixing process. The complex behaviour developing in the molten area
is similar to the conventional fusion welding process. In this type of lap welding, it is
difficult to control the formation of intermetallic phases in the weld. Solid/liquid state
welding is illustrated in Figure 2-25(b). Here, only the top material is melted and a
molten layer is formed between the mating faces of two materials by means of the
conductive heat transferred from the top material. The precise control of the heat input
and the diffusion rate are considered as the main benefits for the solid/liquid state
welding. Furthermore, the solid/liquid state welding is frequently used in laser lap
welding of two materials with significantly different melting points. Finally, solid/solid
state is shown in Figure 2-25(c). This lap welding type presents the same physical
principle of diffusion welding. In addition to the uniform pressure required during the
solid/solid state welding process, it is important to monitor the occurrence of an oxide
film produced between the mating faces [48].

Figure 2-25 Schematic diagram of the interfacial reaction layer in laser lap welding: (a)
the liquid/liquid state; (b) the solid/liquid state; (c) the solid/solid state [48].

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2.8 Metallurgical aspects of laser welding

As aforementioned, welding can be considered as a metallurgical process and the weld

quality is strongly affected by thermal, physical, chemical and mechanical properties of
the welding materials [96]. Microstructure in the weld generally is determined by the
heat transfer and the thermal cycle during laser welding, as shown in Figure 2-26 [23,
97]. The evolution of microstructures in the weld can also be influenced by micro-
segregation and non-equilibrium phase transformations that occur during solidification.
Due to the rapid heating and cooling rates, less micro-segregation is expected to occur
in laser welding than is induced during other fusion welding methods. In some cases,
the non-equilibrium phase transformation can cause intermetallic brittle phases in the
weld resulting in fractures of fatigue cracking [30].

Figure 2-26 Relationship between microstructure and thermal cycle in laser welding of
structural steel [97].

2.8.1 The fusion zone (FZ)

In laser welding, the peak temperature of the FZ is usually higher than the melting point
of the welding materials, for melting and solidification processes to occur. If the
temperature of welding materials reaches to their boiling points, the loss of alloy
elements will occur leading to variations of mechanical properties and the

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microstructure of the weld. Because the fusion zone will undergo a transformation from
the liquid phase to the solid phase, then the size and shape of grains and the distribution
of inclusions and defects (such as porosity and hot cracks) are controlled by the
solidification rate. The microstructures in the FZ are closely related to the temperature
gradient, growth rate, cooling rate and alloy constitution in the weld [23].

2.8.2 The heat affected zone (HAZ)

The temperature in the HAZ is lower than the melting point of materials in laser
welding. Although melting cannot occur in the HAZ, the heat is sufficient to cause
phase transformations. The thermal cycle and temperature gradient in the HAZ can
influence the degree of phase transformation, grain growth, microstructure, composition
gradients and residual stresses. A rapid cooling rate can produce finer microstructures,
limit grain growth and cause non-equilibrium phases in the weld [12]. Figure 2-27
schematically illustrates the HAZ formed during the interaction between the heat source
and the weld material.

Figure 2-27 Interaction between the heat source and the base metal: (a) three distinct
regions in the weld; (b) The convective flow within the weld pool [23].

2.8.3 Thermal cycle

In laser welding, the thermal cycle is composed of rapid heating and cooling phases.
Metallurgical effects in the fusion and HAZs are associated with the peak temperature
of the melt pool and the heating and cooling rates in the weld [98]. Furthermore, the
thermal cycle can be affected by the heat transfer and the fluid flow in the melt pool
resulting in variations of the temperature gradient, the cooling rate, microstructures,

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distortions and residual stresses in the weld [62, 69]. During the process of
solidification, the cooling rate is relatively slower because most the heat is transferred
by means of conduction. In laser welding, the heating and cooling phases almost exist
consistently [36]. For this reason, a weld with the smaller HAZ and the larger thermal
gradient is expected to be produced in laser welding than in any other fusion welding
methods.

2.8.4 Inhomogenities

The chemical inhomogenities in fusion welding have been considered to strongly

influence mechanical properties, cause cracking and reduce corrosion resistance in the
weld. The solute segregation, banding, inclusions and gas porosity in the weld are
generally classified as chemical inhomogenities. According to the scale size of the
solute segregation, it can be divided into micro-segregation and macro-segregation.
Micro-segregation is characterised by its varied composition in the weld across
structures of microscopic size. It closely relates with the cooling rate during
solidification and can affect the susceptibility to solidification cracking in the weld [29].
The welding process with a higher heat input results in a relatively lower cooling rate
which has been pointed out to readily cause micro-segregation in the weld [99].
Meanwhile, macro-segregation is defined as the phenomenon of variations in the local
average composition across structures of macroscopic size, such as the whole weld. It
usually occurs in the dissimilar materials welding, particularly, when the melt pool is
mixed incompletely [29].

2.9 Laser weld defects

In laser welding, the heating and cooling rates can lead to variations of chemical
composition, microstructures and residual stresses in the fusion and HAZs [23]. These
changes can cause the formation of defects and influence the weld quality [100]. In
other words, mechanical properties and corrosion resistance of the weld can be
deteriorated due to the occurrence of phase transformations and defects, such as
porosity, cracking, loss of elements and oxidation within the weld [30].

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2.9.1 Cracking

Cracking is usually caused when the strain in the weld exceeds the strain that can be
sustained by the tensile stress of the materials. The formation of cracking in the weld
can be affected by weld parameters, the weld fixture and chemical composition of
welding materials, as shown in Figure 2-28. The solidification temperature range, the
amount of liquid phases in the weld (at the final stage of solidification), the phases and
the grain structure in the weld determine the formation of cracking [29]. For example,
cracks can be formed when the residual stresses in the weld are higher than the yield
strength of welding materials or by the presence of impurities in the weld [55]. In
another point of view, solidification cracking can be suppressed by increasing the
heating rate in laser welding. This is because a faster heating rate usually leads to a
faster cooling rate particularly with a moving heat source. A higher cooling rate yields
finer microstructure which reduces the degree segregation in the weld [12] and therefore
contributes to limit the occurrence of cracking in the weld.

Figure 2-28 Interaction between process parameters affecting weld solidification

cracking [101].

2.9.2 Spatter

Spatter is molten droplets formed by melted materials being ejected from the keyhole in
laser keyhole welding, as shown in Figure 2-15. The presence of spatter can
contaminate the weld appearance and affect its corrosion resistance. The amount of
spatter is related to the boiling points of the weld materials, the vapour pressure and also

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 Chapter 2 – A literature review on welding technology

the processing parameters [77, 83]. In addition, the size of the melt pool can influence
the amount of spatter as shown in Figure 2-29 [102]. The bigger the melt pool is, the
less amount of spatter can be formed. Several methods have been reported for
suppressing the formation of spatter. They include: the reduction of heat input (by
reducing laser power, increasing welding speed, using a defocused laser beam or
increasing the pulse frequency), the selection of the proper shielding gas and adding a
filler material [77, 103, 104].

Figure 2-29 Relationship between the size of molten zone and the presence of spatter:
(a) view of the cross-sectioned weld; (b) view of the longitudinal-sectioned weld [102].

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 Chapter 2 – A literature review on welding technology

2.9.3 Porosity

Porosity is one common defect occurring in the welding process especially for welding
light metals such as aluminium alloys, magnesium alloy and titanium alloys. The
formation of porosity can degrade the mechanical properties of the weld and lead to its
fracture. In laser welding, the presence of bubbles in the weld could be attributed to the
keyhole collapse. If the interaction time between the incident laser beam and the
keyhole wall is longer, the low surface tension of the molten materials is readily to
induce the keyhole collapse [105]. Zhou and Tsai [106] modelled the mechanism of
porosity formation in pulsed wave laser welding of 304 stainless steel. They indicated
that the formation of porosity was related to the melt pool dynamics, keyhole collapse
and solidification processes. Their results also showed that porosity more readily
appeared in keyhole mode welding than in the conduction mode welding when the weld
width was kept constant. A optimal processing window of pulsed wave laser welding
can effectively suppress the formation of porosity in the weld. Regarding to the effect of
welding speed, Figure 2-30 shows that the melt flow in the tip of keyhole can be
stronger at a lower welding speed than that of a higher welding speed. For this reason,
porosity was easily observed at a lower welding speed in fibre laser welding [67].

Figure 2-30 Schematic illustration of the melt pool dynamic in fibre laser welding at: (a)
a lower welding speed; (b) a higher welding speed [67].

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 Chapter 2 – A literature review on welding technology

In Figure 2-31, Katayama et al. [67] explained the relationship between the formation of
porosity and welding speed in laser welding. They indicated that bubbles resulting from
the collapse of the front keyhole wall can be easily generated and observed at the tip of
keyhole at a lower welding speed. Conversely, at a higher welding speed, the formation
of porosity is less and smaller in the top area of weld due to the flow rate in the bottom
of keyhole is slower and the quicker solidification can restrict the collapse of the front
keyhole wall.

Figure 2-31 Schematic illustration of the formation of porosity at different welding

speeds in continuous wave laser welding [67].

In another point of view, the weld geometry is another factor to influence the formation
of porosity. In Figure 2-32, the formation of porosity can increase by increasing the
ratio of the weld depth to width in laser spot welding of aluminium alloy [100]. Besides,
porosity is readily present in the partial penetration weld compared with the full
penetration weld [12, 100, 107]. Generally speaking, processing parameters, welding
environments, welding materials’ properties and surface contaminations of the welding
materials have to be considered in reducing the formation of porosity [108].

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 Chapter 2 – A literature review on welding technology

Figure 2-32 Schematic illustration of the relationship between the formation of porosity
and the weld shape in laser-spot welding of A1050 Al alloy [100].

2.10 Residual stresses generated in laser welding

Residual stresses and distortions are common problems in fusion welding processes
[109]. First, the definition of residual stresses is the stress that exist in a material or
structure after all external loads are removed. Due to the non-uniform temperature
changes in laser welding, residual stresses can also be referred to thermal stresses [29].
The higher the temperature gradients occurring in the weld, the higher the thermal
stresses that can be induced. The three general situations causing thermal stresses in
fusion welding are summarised as follows [27]:

• Stresses are induced by a volumetric change, such as expansion or shrinkage.

These are associated with phase transformations in the material.
• Stresses are induced by the difference in coefficient of thermal expansion (CTE)
between two materials.
• Stresses are induced by a temperature gradient resulting in different rates of the
heat or contraction within the material.

Theoretically, when the weld cools down to room temperature, residual tensile stresses
can be found in the weld and residual compressive stresses will occur in the areas far
from the weld, as shown in Figure 2-33 [29].

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 Chapter 2 – A literature review on welding technology

Figure 2-33 Distributions of residual stresses in the butt weld: (a) longitudinal residual
stresses (σx); (b) transverse residual stresses (σy) [110].

The distribution of residual stresses in and surrounding the weld is complex and closely
relates to many factors, such as the chemical compositions and thickness of welding
material, the applied restraint and the welding direction [60]. The fast heating and
cooling rates in laser welding can induce metallurgical transformations, large
differences in the thermal expansion and severe thermal gradients. For these reasons,
significant residual stresses can easily form in the weld [12, 55, 111] and affect the
mechanical properties. Residual tensile stresses at the surface of the weld can induce
premature failure and, in contrast, residual compressive stresses at the surface of the
weld can improve its service life [111]. A numerical analysis comparison of the residual
stresses between laser welding of similar and dissimilar materials showed that the
residual stresses produced from similar steels welding is lower than that of dissimilar
steels welding [112]. The post-process heat treatment generally is used to release
residual stresses occurred in and near the weld.

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 Chapter 3 – A literature review on welding of dissimilar materials

Chapter 3
A literature review on welding of
dissimilar materials

3.1 Introduction

Welding of dissimilar materials is always performed in response to a specific demand

from industry [113]. The differences in physical and chemical properties between the
weld materials and their metallurgical imcompatibilities cause problems in the weld,
such as asymmetric welding shapes, the formation of intermetallic brittle phases,
segregation and residual stresses. The differences of thermo-physical properties between
the welding materials therefore play an important role in determining the weld quality;
hence, a review of these properties is needed before welding dissimilar materials.

As examples from the literature: segregation can result from different chemical
compositions; different thermal expansions can cause residual stresses; different heat
dissipation rates can result from the different thermal conductivities and different
densities can affect the degree of convective flow in the melt pool [95, 114]. For these
reasons, this work is highly important.

In FSW and DW, the parent materials are kept in the solid phase during welding;
therefore, it is recognised that these solid state welding processes have the potential to
minimise metallurgical effects when compared to fusion welding. However, solid state
welding is not suitable for welding refractory metals and for precision welding
applications. Furthermore, the HAZ and the degree of material mixing in the weld can
be minimised by using a high density heat source in fusion welding. Therefore, electron
beam welding and laser welding are also suitable for welding dissimilar materials and
producing a weld with less residual stresses and intermetallic brittle phases in

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 Chapter 3 – A literature review on welding of dissimilar materials

comparison with other fusion welding types. In the following section, the characteristics
of FSW and laser welding are briefly introduced and compared for their application in
welding dissimilar materials.

3.2 The state-of-the-art in welding dissimilar materials

3.2.1 Friction stir welding of dissimilar materials

In the past decade, both FSW and laser welding have been used extensively in joining
dissimilar materials. As mentioned in chapter 2, FSW is more suitable for joining
materials which have a lower melting point, high thermal conductivity and reflectivity.
Accordingly, most of the work reported were in welding of similar and dissimilar
aluminium alloys, as summarised in Table 3-1.

Table 3-1 Work reported in FSW of dissimilar materials.

Materials Sheet Joining Application Year Organisation /
thickness type Country
(mm)
Dissimilar Al Sungkyunkwan
4 butt Aerospace 2003
alloys University, Korea [115]
Tohoku University,
Al alloy/ Mg alloy 3 butt Automotive 2004
Japan [116]
DLR-German
Al alloy /
4 butt - 2005 Aerospace Centre,
Stainless steel
Germany [117]
Dissimilar Al National Institute of
5 butt Aerospace 2007
alloys Technology, India [118]
Osaka Prefecture
Al alloy/ Mg alloy 3 butt Automotive 2008
University, Japan [119]
Dissimilar Al Politecnico di Bari, Italy
0.8 butt - 2008
alloys [51]
Al alloy / Zn- 3 (Al alloy) Osaka University,
lap - 2008
coated steel 0.8 (steel) Japan [120]
Al alloy / Mg University of Ulsan,
2 butt - 2008
alloy Korea [121]
Korea Advanced
Low carbon 1 (Al alloy) Institute of Science and
lap Automotive 2009
steel/ Al alloy 0.6 (Steel) Technology, Korea
[122]
4 (Al alloy) Osaka University,
Al alloy / Ti alloy lap Aerospace 2009
2 (Ti alloy) Japan [123]
Dissimilar Al University of Salento,
4 butt - 2009
alloys Italy [124]

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 Chapter 3 – A literature review on welding of dissimilar materials

In comparison with the laser welding process, the main characteristic of FSW is its
higher welding efficiency. Aluminium alloys and magnesium alloys are commonly
welded by FSW due to their high thermal conductivity and reflectivity. Conversely,
acceptable welds of those two alloys cannot be easily obtained with laser welding. The
magnesium alloy to aluminium alloy welds produced with laser welding and FSW are
shown in Figure 3-1(a) and Figure 3-1(b), respectively. In comparison with FSW, a
smaller HAZ and a more severe degree of material mixing were found in laser welding,
where as a reduced formation of intermetallic brittle phases was obtained with FSW.

Figure 3-1 Welds of aluminium alloy to magnesium alloy obtained from: (a) laser
welding (Thickness of Mg alloy was 1.2 mm) [125]; (b) FSW (Thickness of both Al
alloy and Mg alloy was 2 mm) [121].

3.2.2 Laser welding of dissimilar materials

Laser welding is another option for welding dissimilar materials in industry due to its
advantageous characteristics, such as high energy density, low heat input, and rapid
heating and cooling rates [12]. It has been successfully applied for welding different
combinations of materials, as shown in Figure 3-2. Table 3-2 groups reported works in

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 Chapter 3 – A literature review on welding of dissimilar materials

laser welding of dissimilar materials. Their detailed data are sorted in Appendix 1 for
further reference. All of them show that the formation of intermetallic phases is the
most important factor influencing the weld quality in laser welding of dissimilar
materials.

Figure 3-2 Laser welding of dissimilar materials: (a) Cu to Al alloy [126]; (b) AISI 304
stainless steel to AISI 420 stainless steel [79]; (c) steel to Al alloy [127].

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 Chapter 3 – A literature review on welding of dissimilar materials

Table 3-2 Studies of laser welding of dissimilar materials.

Type of welding material Table in Appendix 1
Al alloy to other materials Table A1-1
Cu to other materials Table A1-2
Mg alloy to other materials Table A1-3
Ni alloy to other materials Table A1-4
Stainless steel to other materials Table A1-5
Ti alloy to other materials Table A1-6
Other material types Table A1-7

One clear difference between FSW and laser welding is that the weldable thickness of
the weld materials can range from micrometres to centimetres in scale. Furthermore,
advantages of laser welding (such as a smaller HAZ, high repeatability and high
economic efficiency) provide a range of opportunities in industrial applications.
Although several methods have been used for suppressing the formation of intermetallic
phases in previous works, this issue still is open for further development in laser
welding of dissimilar materials as well as in FSW.

3.2.3 Comparison between LBW and FSW

Figure 3-3 shows the cross-section of typically welds produced from laser welding and
FSW. Welds with smaller HAZ and higher ratio depth-to-width ratio are usually
obtained in laser welding, as shown in Figure 3-3(a). Figure 3-3(b) shows that welds
with wider HAZ are observed from FSW. In general, laser welding has the higher
potential in welding refractory materials within a short working duration. FSW is more
suitable in welding materials which have a low melting point, high thermal conductivity
and high reflectivity [114]. Characteristics of FSW and laser welding are compared in
Table 3-3.

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 Chapter 3 – A literature review on welding of dissimilar materials

Figure 3-3 Butt welding of AA6013-T6 aluminium alloy obtained from: (a) CO2 laser
welding; (b) FSW [128].

Table 3-3 Comparisons between LBW and FSW processesc.

LBW FSW
Joining efficiency + -
High aspect ratio + -
Small heat affected zone + -
High processing speed + -
Weld bead appearance + -
Weld low melting point materials - +
Weld high reflective materials - +
Weld high conductive materials - +
Weld thinner materials + -
Weld without porosity - +

3.3 Mechanisms in laser welding of dissimilar materials

In principle, the weld obtained in laser welding of dissimilar materials undergoes three
main process mechanisms within a very short duration: melting, mixing and
solidification [129]. The fusion and HAZ of the weld can undergo significant changes in
their chemical compositions, microstructures and mechanical properties during these
three phases [29].

c
‘+’ and ‘-’ mean advantages and disadvantages, respectively.

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 Chapter 3 – A literature review on welding of dissimilar materials

3.3.1 Melting

When the laser beam irradiates the top surface of the weld materials, both melting and
vaporisation may develop depending on the amount of heat input and the melting and
boiling points of the weld materials. In laser keyhole welding, the peak processing
temperature is normally higher than the boiling point of the weld materials, leading to
the formation of a vaporising keyhole and the loss of alloy elements [70]. In this process,
the weld materials transfer from the solid state to the liquid state. In laser welding of
dissimilar materials, the degree of material melting is strongly influenced by the thermal
conductivity, melting point and laser beam energy absorption of the weld materials. For
instance, at the same energy density condition, a lower degree of material melting can
be obtained from materials which have higher thermal conductivity, higher melting
point and lower energy absorption. Conversely, more melting can be produced when the
weld materials have lower thermal conductivity, lower melting point and higher energy
absorption. The thickness of the interfacial reaction layer between weld materials is
affected by the melting time [130].

3.3.2 Mixing

After the weld materials are molten, the temperature gradients in the weld may cause
significant surface tension forces (Marangoni force) at the top surface of the melt pool.
The Marangoni force induces a strong circulatory flow in the melt pool to force the
mixing of melting materials. In the melt pool, heat transfer, fluid flow and species
distributions occur during the welding process [129]. The degree of macro-segregation
in the weld is considered to be minimised when the melt pool is completely mixing by
the convective flow. Well-mixed melt pools can be obtained in laser welding and
electron beam welding because of their high energy density beams [29].

3.3.3 Solidification

When the laser beam moves away from the weld zone, the temperatures of the weld
materials drop rapidly to room temperature in a very short time. The molten materials
can quickly transfer from the liquid state to the solid state. The solidification rate in the
weld depends on its cooling rate and the degree of material mixing [129]. Different

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 Chapter 3 – A literature review on welding of dissimilar materials

microstructures can be found at each region of the weld due to the different
solidification rates. Furthermore, the solidification rate plays an important role in
determining the weld quality because it strongly affects the microstructure and the
formation of solidification cracks leading to variations of mechanical properties. For
example, in laser welding of A5083 Al alloy, Haboudou et al. [131] found that
columnar dendritic structures were formed at the edge and bottom areas of welds while
equiaxed dendrites were formed in the upper area of welds.

Solidification cracking and solute redistribution usually occur in the solidification

process [70]. The solute redistribution of alloy elements can induce segregation in the
weld which results in the formation of intermetallic phases [132]. The occurrence of
intermetallic phases relates to the solubility between the welding materials. Table 3-4
presents the weldability observed in laser welding of dissimilar pure metals according to
their solubility.

Table 3-4 Weldability of binary metal in laser welding processes [133].

W Ta Mo Cr Co Ti Be Fe Pt Ni Pd Cu Au Ag Mg Al Zn Cd Pb
Ta E
Mo E E
Cr E P E
Co F P F G
Ti F E E G F
Be P P P P F P
Fe F F G E E F F
Pt G F G G E F P G
Ni F G F G E F F G E
Pd F G G G E F F G E E
Cu P P P P F F F F E E E
Au * * P F P F F F E E E E
Ag P P P P P F P P F P E F E
Mg P * P P P P P P P P P F F F
Al P P P P F F P F P F P F F F F
Zn P * P P F P P F P F F G F G P F
Cd * * * P P P * P F F F P F G E P P
Pb P * P P P P * P P P P P P P P P P P
Sn P P P P P P P P F P F P F F P P P P F

E : Excellent G : Good F : Fair P : Poor * : No data available

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3.4 Technology for laser welding of dissimilar materials

Different methods have been tested to overcome common issues occurring in laser
welding of dissimilar materials. They include the use of pulsed wave lasers, offsetting
the laser beam on different positions and pre- or post-heating treatments on the welding
materials. Most of them were focused on suppressing the formation of intermetallic
brittle phases in the weld.

Compared with the continuous wave laser welding process, pulsed wave laser welding
can precisely control the heat input, the degree of material mixing and the solidification
rate [134]. During the pulsed wave laser welding process, the weld materials will not be
solidified individually because their heating and cooling processes occur at the same
time. It was believed that an appropriate selection of welding parameters, such as the
peak power, the average power, the pulse frequency and welding speed, helps to
improve the weld quality in laser welding of dissimilar materials.

In laser welding of aluminium alloy to steel, Borrisutthekul et al. [130] controlled the
heat flow by using backing blocks to minimise the influence of different thermal
conductivities between the welding materials, as shown in Figure 3-4. Their results
showed that the heat flow and the molten time could be precisely controlled by varying
both the thermal conductivity of back blocks and the welding speed. Sound welds, with
thinner interfacial reaction layer and higher joint strength, were obtained.

Figure 3-4 Illustration of laser welding of dissimilar materials with backing blocks
[130].

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 Chapter 3 – A literature review on welding of dissimilar materials

Berretta et al. [79] used a pulsed Nd:YAG laser to produce defect-free welds of AISI
304 austenitic stainless steels to AISI 420 martensitic stainless steels. Their results
showed that microstructures, hardness distributions and tensile strengths of the welds
varied with the offset distance of the laser beam. A harder weld with martensitic
structure was found when the laser beam was offset to the AISI 420 martensitic
stainless steel side. Better welds, however, were obtained when the laser beam was
either positioned on the interface of welding materials or offset to the AISI 304
austenitic stainless steel side.

3.5 Summary

Welding of dissimilar materials is increasingly used in industry with different welding

technologies owing to its benefits of saving the material cost and increasing the design
flexibility. Amongst these technologies, the laser welding process shows the potential in
this field due to its characteristics of precisely controlling the heat input and producing a
weld with a smaller HAZ and low distortion. Although several technologies have been
applied to improve the weld quality in laser welding of dissimilar materials, remaining
issues, such as the formation of intermetallic phases in the weld and welding the next
generation materials (i.e. aluminium alloy and magnesium alloy), are still open to be
developed.

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 Chapter 4 – Materials, equipments and experimental procedures

Chapter 4
Materials, equipments and experimental
procedures

4.1 Introduction

This chapter first focuses on the materials used in the present research work. The second
part details the equipments and the experimental procedures that were employed to
ensure accurate and repeatable results were obtained from all studies.

4.2 Materials

Four different materials were used in this work, these being Ti-6Al-4V titanium alloy,
Inconel 718 nickel alloy, DX54 Zn-coated steel and EN-AW-5754 aluminium alloy.
Their material properties and applications in the laser welding process are briefly
introduced in the following sections. Table 4-1 presents material properties of pure
metals which are key constituents of the above weld materials.

Table 4-1 Material properties of pure metals [24].

Properties Zn Fe Ti Ni Al
Melting point (°C) 419 1538 1668 1455 660
Boiling point (°C) 907 2861 3287 2913 2520
Density (g/cm3) 7.14 7.87 4.51 8.90 2.70
Specific heat (J/kg · °K ) 388 449 523 444 897
Thermal conductivity (W/m · °K ) 113 80 11 82 238
-6
Coefficient of expansion (10 /°K ) 30.2 12.1 8.6 13.4 23.5

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 Chapter 4 – Materials, equipments and experimental procedures

4.2.1 Titanium alloy

Titanium alloys (Ti alloy) have been widely used in the petrochemical, aerospace,
medical and leisure industries. Their characteristics are good corrosion resistance at
elevated temperature, high ratio of strength to weight, and good fatigue and creep
resistance. However, Ti alloys have an affinity with oxygen, hydrogen and nitrogen,
which readily generates embrittlements when welding them without inert gas protection
[109, 135-138]. Today, almost 50% of titanium alloy usage is in aerospace applications
[138] because they can save the total weight, operate at a higher temperature, have
better corrosion resistance and compatibility with composite materials [139].

Pure Ti has two allotropic crystal structures, as shown in Figure 4-1. When the
transformation temperature is above 883 °C, the b phase with the body centre cubic
(BCC) crystal structure is formed. The α phase with the hexagonal close-packed (HCP)
structure occurs when the transformation temperature is lower than 883 °C [137, 140].

Figure 4-1 Allotropic crystal structures of pure Ti.

Ti alloys are classified into commercial pure titanium (CP Ti), alpha titanium alloy (α Ti
alloy), beta titanium alloy (b Ti alloy) and alpha-beta titanium alloy (α-b Ti alloy). CP
Ti has good weldability [139] and it is commonly used in dental applications [141].

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 Chapter 4 – Materials, equipments and experimental procedures

Mechanical properties of CP Ti are strongly influenced by the minor interstitial

dissolved elements (i.e. oxygen, nitrogen and carbon) [141]. Four grades are included in
CP Ti according to their chemical composition, as shown in Table 4-2.

Table 4-2 Classification of commercial pure Ti [142].

Unalloy Tensile 0.2% Chemical composition (wt%)
grade strength yield Ti N C H O Fe
ASTM (MPa) strength (max.) (max.) (max.) (max.) (max.)
(MPa)
Grade 1 240 170 99.5 0.03 0.10 0.015 0.18 0.20
Grade 2 340 280 99.2 0.03 0.10 0.015 0.25 0.30
Grade 3 450 380 99.1 0.05 0.10 0.015 0.35 0.30
Grade 4 540 480 99.0 0.05 0.10 0.015 0.40 0.50

Properties of α, α-b and b Ti alloys are related to their microstructures, the volume
fraction of α and b phases [138]. α Ti alloys have a lower ductility than that of b Ti
alloys [143]. They are suitable for applications at elevated temperatures because their
characteristics of high resistance to plastic deformation, high creep resistance and lower
diffusion rate [138, 139]. b Ti alloys are particular suitable for applications requiring a
higher strength and corrosion resistance [138, 139]. b Ti alloys are heat treatable and
high formablilty. They have high strength, good toughness and high corrosion resistance
[140], but poor weldability. Characteristics of α-b Ti alloys are related to the b-stabiliser
content. Decreasing the b-stabiliser content of α-b Ti alloys can increase their ductility
and weldability.

Ti-6Al-4V α-b alloys contain about 6% Al for α stabilisation and 4% V for b

stabilisation. They have been widely used in industry due to their superior properties of
high tensile strength and good formability. Regarding applications, the usage of Ti-6Al-
4V is more than 50 % in both Europe and USA. Ti-6Al-4V alloys [138, 144]. They are
suitable for applications from 315 °C to 400 °C. The microstructure of Ti-6Al-4V alloys
can be modified from the b phase to the acicular α-b phase by controlling the cooling
rate. The b phase transforms and displaces to the martensite structure (α'-HCP or α''-
orthorhombic) at a high cooling rate, as shown in Figure 4-2. A slow cooling rate
induces the Widmanstätten α-b microstructure from the b phase. The acicular α-b

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 Chapter 4 – Materials, equipments and experimental procedures

structure is particularly attractive for aerospace applications requiring high specific

strength, high creep strength, fracture toughness and fatigue crack propagation
resistance [145].

Figure 4-2 Microstructure of fibre laser welding of Ti-6Al-4V Ti alloy: (a) the fusion
and heat affected zones, 100X; (b) the fusion zone, 200X [146].

In the laser welding of Ti alloys process, contaminations can cause the formation of
porosity, embrittlement, discoloration and poor penetration. Therefore, cleaning the
welding materials and providing a proper shielding gas arrangement are necessary in
welding of Ti alloys [107, 141, 147].

4.2.2 Nickel alloy

Nickel alloys (Ni alloys) are widely used in high temperature applications, such as jet
engines, rockets, nuclear power plants and the marine industry [25]. They have good
mechanical properties and good corrosion resistance at high temperatures. According to
chemical composition, Ni alloys can be grouped into binary Ni alloy (i.e. Ni-Cr alloy
and Ni-Cu alloy) and ternary Ni alloy systems. The Ni-Fe-base superalloy is widely
used in comparison with other ternary Ni alloys.

Both Ni (25 – 45%) and Fe (15 – 60%) are the main elements contained in Ni-Fe-base
superalloy. Cr and Mo usually are added to improve the oxidation resistance and
increase the strength, respectively [140]. Inconel 718 is an age-hardenable and high-

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 Chapter 4 – Materials, equipments and experimental procedures

strength alloy with good corrosion resistance. It is suitable for a wide range of
temperatures from -205 °C to 705 °C, as requested in the aerospace, power and nuclear
industries [148-150]. Its strength is increased by precipitating the Ni3Nb (γ'') and
Ni3(Al, Ti) (γ') phases, as shown in Figure 4-3(a) [151].

Figure 4-3 Microstructure of Inconel 718 Ni alloy: (a) the parent material; (b) the fusion
zone obtained from the electron beam welding process [152].

Issues developing in welding of Inconel 718 are the segregation of Nb, the formation of
brittle imtermetallic Laves phase, see in Figure 4-3(b), and the solidification cracking
and microfissuring in the heat-affected zone [153]. The size and distribution of Laves
phase relate to the degree of segregation and dendrite structure during the solidification
process [148, 149, 152-154]. Regarding to the Laves phase in the weld, it could be
restricted by applying a solution post-treatment at the temperature higher than 1040 °C
[152]. Using a pulsed wave laser welding process in controlling the cooling rate was
another method to minimise the segregation and the Laves phase in the weld [155].

4.2.3 Zn-coated steel

Zn-coated steels have properties of good corrosion resistance, formability, weldability

and paintability [156]. Electrogalvanised (EG) and galvannealed (GA) steels are
common types used in industry today. In comparison with galvannealed steels,
electrogalvanised steels are more expensive [157, 158]. Recently, Zn-coated steels are
widely used in manufacturing automotive components due to the increasing requirement
on corrosion resistance of vehicles. Initially, Zn-coated steels were used only on
unexposed automotive parts. They recently have been commonly applied for exposed

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 Chapter 4 – Materials, equipments and experimental procedures

automotive components by increasing the thickness of Zn coating and by applying

layers of paint on the Zn coating, as shown in Figure 4-4 [157].

Figure 4-4 Schematic illustration of a painting sample with the Zn-coated steel in the
automotive application [159].

Laser lap welding of Zn-coated steels is a common process used in the automotive
industry [160]. Issues that arise in laser lap welding of Zn-coated steels are the
formation of high Zn vaporisation, spatter and porosity, as shown in Figure 4-5 [161-
164]. Leaving an appropriate gap between the Zn coated steels [161, 165] and adding
copper powders between the Zn coated steels [49] have been used to minimise those
problems.

Figure 4-5 The formation of spatter on the weld surface in Nd:YAG PW laser lap
welding of DX56 Zn-coated steels [166].

4.2.4 Aluminium alloy

The usage of aluminium alloys (Al alloys) is at the second position compared with other
metals in the metal market. They are gradually and extensively used in industry due to

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 Chapter 4 – Materials, equipments and experimental procedures

their properties of high strength-to-weight ratio, good corrosion resistance, high thermal
and electrical conductivity, and are nontoxic. Industrial sectors for Al alloy use are
presented in Figure 4-6. Regarding manufacturing processes, Al alloys are classified as
either cast or wrought Al alloys [25]. Table 4-3 presents the Al alloy types, according to
their chemical composition. Al-Mg alloys (5xxx) and Ai-Mg-Si alloys (6xxx) are
commonly Al alloys used in the automotive applications [70]. Al-Cu alloys are suitable
for manufacturing electrical components. Al-Mn alloys have a moderate strength and
good workability, but they are not heat-treatable [140].

10%

Transport
11%
35%
Building & Construction

Packaging

Electrical
22%
Consumer Goods

22%

Figure 4-6 Industrial sectors of Al alloy in 2007 [167].

Table 4-3 Groups of Al alloys [140].

Wrought Al alloys Cast Al alloys
Pure aluminium 1xxx 1xx.x
Major alloyed elements
Cu 2xxx 2xx.x
Mn 3xxx -
Si with Cu and/or Mg - 3xx.x
Si 4xxx 4xx.x
Mg 5xxx 5xx.x
Mg and Si 6xxx -
Zn 7xxx 7xx.x
Sn - 8xx.x
Other element 8xxx 9xx.x

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 Chapter 4 – Materials, equipments and experimental procedures

Material properties of Al alloys (i.e. high thermal conductivity, high reflectivity and low
boiling point) strongly affect the results from the laser welding of Al alloy process [168,
169]. The wide range of solidification temperature easily causes cracks in the weld, as
shown in Figure 4-7 [170]. Nevertheless, welding of Al alloy is increasingly used in the
automotive industry, due to its benefit of reducing the vehicle weight.

Figure 4-7 The formation of cracks in Nd:YAG laser welding of AA2024 Al alloy [171].

4.3 Fibre laser welding system

4.3.1 Introduction to fibre lasers

The first fibre laser was demostrated in 1961 [172, 173]. Fibre lasers are pumped by
diode laser sources and posess a unique combination of laser power the beam quality
[174]. An Ytterbium-doped fibre laser with two diode laser pumps is illustrated in
Figure 4-8. From the inner to outer parts, a double-clad fibre laser is composed of an
inner quartz glass fibre (doped core), a double clad fibre (pump core) and an outer
cladding, as illustrated in Figure 4-9. The inner fibre is doped with a rare earth, such as
Erbium, Ytterbium, Neodymium or thallium atoms. The diameter of the doped core is
between 8 µm and 10 µm. The pump core has the diameter of more than 200 µm. The
outer cladding, made up of glasses or polymeric materials, has a low refractive index
coefficient to avoid signal loss [1, 4].

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Figure 4-8 Arrangement of an Ytterbium-doped large-core fibre laser with two pump
sources. (HR: high reflectivity; HT: high transmission) [175].

Figure 4-9 Schematic illustration of a double-clad fibre laser [176].

Fibre lasers include single mode and multi mode types. Currently, single mode fibre
lasers are available with power from several watts to 5 kW. The maximum output power
of a multi mode fibre lasers is more that 50 kW. The characteristics of fibre lasers are
good beam quality, high energy efficiency, low cost of maintenance and compact size.
They are suitable for most laser materials process applications. Furthermore, they have
been considered to replace other types of laser systems (i.e. Nd:YAG land CO2 lasers)
in industry [2]. Comparisons between commercial laser systems are presented in Table
4-4. Fibre lasers have a higher performance in economic efficiency, as shown in Figure
4-10. Because of their superior characteristics, fibre lasers are able to produce a deeper
penetration weld by using a low laser power [174]. Fibre lasers also have applications in
the sectors of medical devices, military and telecommunications.

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 Chapter 4 – Materials, equipments and experimental procedures

Table 4-4 Comparisons between conventional laser systems [174, 177, 178].
CO2 Nd:YAG HPDL Fibre Laser
1.07-1.08 (Ytterbium)
Wavelength (µm) 10.6 1.06 0.8-1.0
1.8-2.0 (Erbium)
1-3 (lamp pumped)
Efficiency (%) 5-15 20-50 60-70
10-20 (diode pumped)
50 (multi-mode)
Output power (kW) up to 45 up to 8 up to 8
5 (single-mode)
Average energy 6-8 5-9 3-5 8
2
10 10 10 10
intensity (J/mm )
Maintenance
1000 200 (Lamps) free 10000-100000
periods (hour)

Figure 4-10 Operating costs of commercial laser systems [179].

4.3.2 Set-up for fibre laser welding of dissimilar materials

Figure 4-11 illustrates the experimental setup used in this research work. An IPG YLR-
1000-SM ytterbium-doped single mode fibre laser machine was used. Its characteristics
are listed in Table 4-5. The HP1.5”(Z)/FL laser cutting head adapted to welding
manufacturing by Precitec KG International (see Figure 4-12) was used. A 190 mm
focal length lens was applied to focus the laser beam. The spot diameter of the focused
fibre laser beam is approximately 72 µm with a Gaussian intensity distribution. Argon
or nitrogen was used as the shielding gas, delivered by a 2 mm exit diameter coaxial

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nozzle. The shielding gas pressure and its flow rate were kept at 2 bar and 25 l/min,
respectively.

Figure 4-11 Experimental setup of fibre laser welding system.

Table 4-5 Characteristics of IPG YLR-1000-SM fibre laser.

Characteristics Value
Emission wavelength 1070 nm
Beam quality factor 1.1 M2
Output fibre core diameter 14 µm
Minimum output power 100 W
Maximum output power 1000 W

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Figure 4-12 Used laser head, jig and single axle stage.

In order to minimise the gap between the welding materials, a custom-made jig, as
shown in Figure 4-12 and Figure 4-13, was designed and used. Drawings of the custom-
made jig are presented in Appendix 2. Fixtures used in laser butt welding and laser lap
welding experiments are illustrated in Figure 4-14(a) and Figure 4-14(b), respectively.
A Baldor single axis high speed linear motor controlled stage was used in this work, as
shown in Figure 4-12.

Figure 4-13 Illustration of the custom-made jig.

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Figure 4-14 Illustrations of welding fixture for: (a) laser butt welding; (b) laser lap
welding.

4.4 Analytical facilities

4.4.1 Macrostructure and microstructure observation

After welding, welds were ultrasonic cleaned with acetone for 3 minutes. The weld
appearance was observed using an optical microscope and a scanning electron
microscope. The Polyvar-MET optical microscopy was manufactured by Reichert/Leica
Austria, with a personal computer running Solution DT software, as shown in Figure
4-15. The Hitachi-S3400N scanning electron microscope manufactured by Hitachi High
Technologies, Japan, as shown in Figure 4-16. Technical characteristics of Hitachi-
S3400N are listed in Table 4-6. The surface roughness of welds was measured using a
Veeco Wyko NT 1100 optical profiling system, as shown in Figure 4-17. Table 4-7
presents the specification of Wyko NT 1100 interferometer. For further evaluations,
welds were sectioned, mounted, ground, polished with diamond slurry to 3 µm surface
finish and etched. The Zn-coated steel samples were etched with Nital acid (3% HNO3
ethanol solution). Keller’s reagent (1 ml hydrofluoric acid, 2.5 ml hydrochloric acid, 2.5
ml nitric acid and 95 ml deionised water) was used to etch the Ti alloy and Al alloy
samples.

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Figure 4-15 PolyVar – MET optical microscope.

Figure 4-16 Hitachi S-3400N scanning electron microscope.

Table 4-6 Specification of Hitachi 3400N scanning electron microscope.

Item Value
Resolution SE 3.0 nm at 30 kV (High Vacuum Mode)
10 nm at 3 kV (High Vacuum Mode)
Resolution BSE 4.0 nm at 30 kV (Low Vacuum Mode)
Magnification 5X – 300,000X
Accelerating voltage 0.3 – 30 kV

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Figure 4-17 Veeco Wyko NT1000 optical profiling system.

Table 4-7 Technical specification of Veeco Wyko NT1000 optical profiling system.
Item Value
Vertical measurement range 0.1 nm to 1 mm

Vertical resolution < 1 Ǻ Ra

Vertical scan speed Up to 7.2 µm/sec

4.4.2 Microhardness test

Hardness measurement were made according to the British standard BS EN ISO 6507-
1:2005 [180] using a MICROMET 5114 microindentation hardness tester manufactured
by BUEHLER Instruments, as shown in Figure 4-18. The test samples are forced with a
diamond indenter. The loading force of this machine ranges from 1 gf to 2000 gf.
Diagonal lengths of indentations were measured under the magnification of 50X. The
test conditions (i.e. the testing locations, the loading force and the testing duration) are
described in each of the following experimental chapters.

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Figure 4-18 MICROMET 5114 microhardness test machine.

4.4.3 Shear force test

Shear force was measured according to the British standard of BS EN 10002-1: 2001
[181] at room temperature using an INSTRON 4507 universal testing machine, see
Figure 4-19, operating with a crosshead speed of 1 mm/s. The dimensions of shear force
testing samples used are described in each of the following experimental chapter. After
the shear force test, the fracture surface was observed using a Hitachi S-3400N SEM.

Figure 4-19 INSTRON 4507 universal testing machine.

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4.4.4 Chemical composition analysis

Chemical composition analysis, including raster area scanning, line intensity scanning
and point counting, was carried out using a Zeiss EVO 50 scanning electron microscope
equipped with an energy dispersive spectrometer (EDS) manufactured by OXFORD
Instruments, as shown in Figure 4-20. The specification of Zeiss EVO 50 SEM is listed
in Table 4-8.

Figure 4-20 Zeiss EVO 50 scanning electron microscope equipped with an Oxford
instruments EDS detector.

Table 4-8 Specification of the Zeiss EVO 50 scanning electron microscope.

Item Value
Resolution SE 3.0 nm at 30 kV
Resolution BSE 4.5 nm at 30 kV
Magnification 5X to 1,000,000X
Accelerating voltage 0.2 – 30 kV

4.4.5 Corrosion test

The corrosion test procedures were followed according to ASTM standards G1:1981,
G3:1989 and G61:1986 [182-184]. The corrosion resistance of welds was evaluated
using electrochemical polarisation analysis in 3.5% NaCl solution. The full setup of the
electrochemical test is presented in Figure 4-21. The welded samples were cleaned with

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acetone before the corrosion test. All samples had 1 cm2 of weld area exposed, which
the remaining surfaces were sealed with the 45-stopping off lacquer as illustrated in
Figure 4-22. In the electrochemical polarisation method, the reference electrode,
auxiliary electrode and working electrode were composed from a saturated calomel
electrode (SCE), a platinum electrode and the welded sample, respectively. Each sample
was started at -15 mVSCE of the open circuit potential with a scanning rate of 0.7
mV/sec until 1.0 VSCE. After testing, samples were ultrasonically cleaned with acetone
for 3 minutes.

Figure 4-21 Experimental setup of electrochemical polarisation corrosion test.

Figure 4-22 Schematic illustration of the corrosion test sample.

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Chapter 5
Fibre laser welding of Ti-6Al-4V titanium
alloy to Inconel 718 nickel alloy

5.1 Introduction

Titanium alloys like Ti-6Al-4V have been widely used in industry due to the excellent
characteristics of good corrosion resistance, high strength and creep resistance. The
largest use of Ti-6Al-4V alloy is in the aerospace industries, for example as static and
rotating components in the turbine engines [139]. Meanwhile, Inconel 718 nickel alloy,
a high temperature material, is also widely used in the aerospace industries. Because of
its superior mechanical properties and oxidation resistance at elevated temperatures,
Inconel 718 is particularly suitable for components in the high temperature regions of
aero engines and gas turbines [140, 151].

Today, the dissimilar materials welding process is increasingly attracting more attention
in industry because it can reduce the material costs and improve the design flexibility.
However, the formation of brittle phases, cracks and residual stresses still readily occur
in a weld between dissimilar materials due to the differences in physical and chemical
properties [12, 114, 134]. A limited amount of systematic research in this area [63, 130,
185-191] has been carried out until now. Schubert et al. [63] pointed out that controlling
the diffusion mechanism appropriately by applying a lower heat input can reduce the
formation of brittle intermetallic phases in dissimilar materials welds. They obtained a
better weld with a combination of a higher laser power and a higher welding speed in
welding aluminium-steel and aluminium-magnesium joints. Using a high energy density
laser beam to restrict the amount of energy input was another suggested method to

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control the heat diffusion and therefore minimise the thickness of the reactive interlayer
and avoid the formation of brittle intermetallic phases especially in cases of welding
steel-kovar, copper-steel and copper-aluminium joints [190]. Regarding the issue of
different conductivities between welding materials, applying a backing block below
welding samples has been pointed out as a method to control the heat flow and
effectively suppress the thickness of the intermetallic layer in welding of steel to
aluminium alloy [130].

According to Ti-Ni equilibrium diagram (Figure 5-1), three intermetallic phases, Ti2Ni,
TiNi and TiNi3, are readily formed in a Ti-Ni system at different chemical
compositions. Ti2Ni and TiNi3 phases have been recognised as the brittle phases which
can increase hardness and reduce strength. Conversely, the TiNi phase has properties of
high ductility, non magnetic, good corrosion resistance and good toughness at low
temperature. For these reasons, TiNi titanium alloy has been widely used in the medical
industry, such as dental applications [192], stents and surgical devices [193].

Figure 5-1 Ti-Ni equilibrium diagram [194].

Considering the welding of titanium and its alloys to nickel and its alloys: Seretsky and
Ryba [113] used a Nd:YAG laser to investigate spot welding Ti to Ni with and without

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TiNi filler. Cracks and incompletely mixed liquids were observed in the welds.
Chatterjee et al. [95, 195] butt welded Ti/Ni dissimilar materials using a CO2 laser to
investigate the solidification microstructure. They found that an asymmetric shape of
weld, macro-segregation and brittle intermetallic phases, Ti2Ni and TiNi3, were readily
generated within the weld with macroscopic cracks.

In spite of this, less work has been reported in this field related to fibre laser welding of
dissimilar materials. The purpose of this chapter is to investigate the influence of
processing parameters on the weld quality in fibre laser welding of Ti-6Al-4V to
Inconel 718. The melt pool shapes are analytically modelled using a two-dimensional
model.

5.2 Experimental investigation

Sheets of Ti-6Al-4V and Inconel 718 with 2 mm thickness were laser butt welded
together by a single mode fibre laser operating in the continuous wave mode. The
chemical compositions and physical properties of these two materials are shown in
Table 5-1, Table 5-2 and Figure 5-2.

Table 5-1 Chemical composition (wt.%) of Ti-6Al-4V and Inconel 718 [196, 197].
Ti-6Al-4V Inconel 718
Fe 0.40 d Balance
Al 5.5-6.75 0.20-0.80
V 3.5-4.5 -
Cr - 17.2-21.0
Nb+Ta - 4.75-5.50
Mo - 2.80-3.30
Ti Balance 0.65-1.12
Ni - 50.0-55.0

d
The maximum limit.

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Table 5-2 Physical properties of Ti-6Al-4V and Inconel 718 [24, 198-201].
Ti-6Al-4V Inconel 718
Melting point (°C) 1655 1260-1336
Boiling point (°C) 3315 2917
Density (g/cm) 4.42 8.91
Specific heat (J/kg K) 610 435
Coefficient of Expansion (10-6 /K ) 8.0 13.0
Latent heat (kJ/kg) 290 272
Solidus temperature (°C) 1605 1260
Liquidus temperature (°C) 1655 1336
Thermal conductivity (W/m K)
at 20 °C 5.8 11.4
at ~Tm/2 17.5 21.3
Thermal diffusivity (m2/s)
at 20 °C 2.15ä10-6 2.94ä10-6
at ~Tm/2 6.49ä10-6 5.50ä10-6

Figure 5-2 Thermal properties of Ti-6Al-4V and Inconel 718 [202-204].

(Cp and k mean specific heat and thermal conductivity, respectively.)

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The full setup of this experiment is described in chapter 4. During welding, argon was
supplied co-axially and laterally as the shielding gas. The shielding gas prsessure and
flow rate were kept at 2 bar and 25 l/min, respectively. Both Ti-6Al-4V and Inconel 718
samples for welding were 25 mm × 50 mm × 2 mm. The welding fixture in fibre laser
butt welding is shown in Figure 4-14(a). The surface roughness of Ti-6Al-4V and
Inconel 718 samples were, approximately, 4.66 Ra and 4.36 Ra, respectively. Before
welding, each sample was cleaned with acetone. The focal position of the laser beam
was set at the top surface of the plates in this experiment. The spot diameter of the
focused laser beam was approximately 72 µm.

A series of experiments was carried out to investigate the correlation of laser power,
welding speed and the offset of the laser beam from the interface to achieve the weld
quality of full penetration welds. In Table 5-3, the one variable at a time (OVAT)
method was used. Firstly, to investigate the influence of the laser power on the weld
quality, four levels of the laser power, 700, 800, 900 and 1000 W, were tested at a fixed
welding speed when the laser beam was positioned on the interface of samples. Next, a
study of the effect of welding speed was carried out by using different welding speeds,
60, 80 and 100 mm/s at a fixed laser power of 1000 W. In the final experiment, three
different offset positions of the laser beam - on the interface of samples, offset from the
interface of the welding materials to the Ti-6Al-4V side 35 µm and offset from the
interface of the welding materials to the Inconel 718 side 35 µm - were tested while
other parameters were kept constant to find the relationship between the laser beam
offset position and the weld quality.

Table 5-3 Experimental matrix.

Welding speed (mm/s) Laser power (W)
60 700, 1000e
80 700, 800, 900e, 1000e
100 700, 800e, 900, 1000e

e
means that three different laser beam offset positions - on the interface, offset from the
interface of the welding materials to the Ti-6Al-4V side 35 µm and offset from the
interface of the welding materials to the Inconel 718 side 35 µm - were carried out
individually with these combinations of welding parameters.

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After welding, all samples were sectioned across the weld, mounted in conductive resin,
polished with diamond abrasives to 1 µm surface finish and etched with Keller’s reagent
for further examination. The weld bead geometry including the weld width and depth
was measured using optical microscopy with a personal computer running Solution DT
software as shown in Figure 4-15. The mean diameter of porosity and length of crack in
welds were also observed and calculated from the cross section of welds by optical
microscopy. Microstucture and phenomena of micro-segregation within welds were
observed by means of optical microscopy and scanning electron microscopy equipped
with backscattered electron imaging (BEI) and energy dispersive spectrometry (EDS).
Profiles of microhardness including the base metals and the weld were tested using a
Vickers microhardness machine, with a 100 g load applied for 10 seconds. Figure 5-3
schematically illustrates evaluations of the weld dimension and the hardness distribution.

Figure 5-3 Schematic diagram of the weld dimension and hardness tests.

5.3 Experimental results

5.3.1 The weld geometry

As shown in Figure 5-4, each full penetration weld has near parallel sides under all the
values of the laser power, welding speed and the laser beam offset position used in the
experiments. In each case, the weld profile on the Ti-6Al-4V side is straighter than on
the Inconel 718 side. The weld geometries were clearly different between three different
laser beam offset positions when laser power and welding speed were 800 W and 100
mm/s, respectively, as shown in Figure 5-4.

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Figure 5-4 Macrostructure of the cross sectional welds when laser power and welding
speed were kept at 800 W and 100 mm/s, respectively, for the different laser beam
offset positions : (a) on the bond; (b) offset 35 μm on the Ti-6Al-4V side; (c) offset 35
μm on the Inconel 718 side.

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For the case of the laser beam positioned at the joint interface, the relationship found
between the weld geometry, welding speed and laser power is shown in Figure 5-5.
When welding speed was kept at 80 mm/s and 100 mm/s, results show that, in both
cases, the weld width randomly varied when laser power increased from 700 W to 1000
W as shown in Figure 5-5(a). Meanwhile, the influence of welding speed on the weld
width with the constant laser power of 1000 W was shown in Figure 5-5(b). The weld
width decreased from 564 µm to 351 µm when welding speed was increased from 60
mm/s to 100 mm/s.

(a)

(b)
Figure 5-5 The weld width of full penetration welds with: (a) different laser powers; (b)
different welding speeds at a constant laser power of 1000 W. (The laser beam was
positioned on the interface of welding plates.)

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The relationship between the laser beam offset position and the weld width is shown in
Figure 5-6. When the laser power and welding speed were 900 W and 80 mm/s,
respectively, a weld width of 603 µm was obtained when the laser beam was positioned
on the interface. Meanwhile, the weld width with the laser beam offset to the Ti-6Al-4V
side and the Inconel 718 side was around 495 µm and 458 µm, respectively. With the
laser power increased to 1000 W, a narrower weld width was obtained when the laser
beam was positioned on the interface. A slightly wider weld was obtained when the
laser beam was offset to the Ti-6Al-4V side.

Figure 5-6 The weld widths with three different laser beam offset positions at the
constant welding speed of 80 mm/s.

5.3.2 The weld defects

The relationship between the formation of porosity, laser power, welding speed and the
laser beam offset position is shown in Figure 5-7. Porosity was produced at a wide
range of parameter combinations. The offset position of the laser beam probably was
not the main factor to determine the formation of porosity. The diameter of micro-
porosities observed from this work ranged from 15 µm to 172 µm. Figure 5-8 shows the
relationship between the formation of crack, laser power, welding speed and the laser
beam offset position. Cracks were produced at a wide range of parameter combinations.
However, crack-free welds were more readily obtained at a higher laser power and a
higher welding speed. As previously, the offset position of the laser beam was not a key

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factor to influence the formation of crack in the weld. The mean crack length in each
weld was between 63 µm and 663 µm.

Figure 5-7 Relationship between the formation of porosity, laser power, welding speed
and the laser beam offset position. (“Centre”, “Ti side” and “Ni side” mean the laser
beam was positioned on the interface, offset to the Ti-6Al-4V side and offset to the
Inconel 718 side, respectively.)

Figure 5-8 Relationship between the formation of crack, laser power, welding speed and
the laser beam offset position. (“Centre”, “Ti side” and “Ni side” mean the laser beam
was positioned on the interface, offset to the Ti-6Al-4V side and offset to the Inconel
718 side, respectively.)

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5.3.3 Hardness distribution of the weld

Figure 5-9 shows the relationship between hardness distributions and laser power when
the laser beam was positioned on the interface. It indicates that higher hardnesses
occurred near the FZ in comparison with the parent materials. No clear trend was found
between laser power and hardness variations. When laser power and the laser beam
offset position were 700 W and on the interface, respectively, the influence of welding
speed on hardness variations is shown in Figure 5-10. Results show that less hardness
variations between the FZ and parent materials was obtained with the welding speed of
60 mm/s while more significant hardness variations were found at a higher welding
speed of 80 mm/s or 100 mm/s.

Figure 5-9 Hardness distributions of welds in fibre laser welding of Ti-6Al-4V to

Inconel 718 at the constant welding speed of 80 mm/s.

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Figure 5-10 Hardness distributions of welds in fibre laser welding of Ti-6Al-4V to

Inconel 718 at the constant laser power of 700 W.

Figure 5-11 displays the relationship between the laser beam offset position and
hardness variations when laser power and welding speed were 900 W and 80 mm/s,
respectively. Obviously, hardness variations near the FZ was found when the laser beam
were positioned on the interface or offset to the Ti-6Al-4V side as shown in Figure
5-11(a) and Figure 5-11(b), respectively. On the other hand, in Figure 5-11(c), the
hardness variation near the FZ was minimal when the laser beam was offset to the
Inconel 718 side.

(a)
Figure 5-11 Hardness distributions in the fibre laser welding of Ti-6Al-4V to Inconel
718 welds with different laser beam offset positions when laser power and welding
speed were 900 W and 80 mm/s, respectively: (a) on the interface; (b) offset to the Ti-
6Al-4V side; (c) offset to the Inconel 718 side.

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(b)

(c)

Figure 5-11 (Continued).

5.3.4 Micro-segregation of the weld

Welds produced at different welding conditions are shown in Figure 5-12 and Figure
5-13. Figure 5-12 is a partial cross-sectioned area observed under an optical microscope
when welding conditions were 800 W, 60 mm/s and the laser beam was offset on the
Inconel 718 side. It is likely that vortices occurred in the weld producing different
microstructures. Hardness and chemical compositions of Points A-G are tabulated in
Table 5-4 and Table 5-5, respectively. Higher hardnesses were obtained at Points B, F
and G. According to the Ni-Ti phase diagram [194] and their chemical compositions in
Table 5-5, Points B and G could be identified as the TiNi3 phase. Meanwhile, Point D

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with 814.1 Hv was classified as the Ti2Ni phase and Point A (263.5 Hv) and Point E
(375.4 Hv) are the hardness of un-welded Ti-6Al-4V and Inconel 718, respectively.

When welding conditions were 700 W, 80 mm/s and the laser beam was offset on the
Ti-6Al-4V side, a backscattered electron image near the top area of weld was taken and
is shown in Figure 5-13. Because the atomic number of Ni is higher than Ti, the Ti-6Al-
4V and Inconel 718 can be easily identified as the black and grey colour, respectively.
Several material mixes were found within the weld. The molten material near the Ti-
6Al-4V side was understandably richer in this alloy than that near the Inconel 718 side
and clear vortices were found in the weld. Hardness at Points H-M are listed in Table
5-4. A high hardness of 389.2 Hv occurred at Point I.

Figure 5-12 Micro-segregation in Ti-6Al-4V/Inconel 718 weld at 800 W, 60 mm/s and

with the laser beam offset to the Inconel 718 side (Optical microscope image).

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Figure 5-13 Micro-segregation in Ti-6Al-4V/Inconel 718 weld at 700 W, 80 mm/s and

the laser beam was offset to the Ti-6Al-4V side (SEM-Backscattered electron image).

Table 5-4 Hardness of Points A-G in Figure 5-12 and Points H-M Figure 5-13.
Hardness (Hv)
A 263.5
B 889.0
C 319.1
D 814.1
E 375.4
F 906.2
G 909.1
H 251.2
I 389.2
J 344.7
K 353.0
L 367.0
M 368.6

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Table 5-5 Chemical composition (wt.%) of Points A-G in Figure 5-12.

A B C D E F G
Ti - 19.12 7.78 55.42 90.29 33.40 20.85
Al - - - - 5.50 2.31 -
V - - - 3.50 4.21 - -
Ni 55.27 44.39 50.62 25.05 - 35.06 45.59
Cr 21.36 17.23 18.72 6.00 - 12.72 18.26
Nb - 4.84 4.72 - - - -
Fe 17.42 14.42 16.18 7.28 - 11.23 15.30
Hg 5.95 - - - - 5.28 -
S - - 1.98 - - - -
Tm - - - 3.74 - - -
Phase Inconel TiNi3 Inconel Ti2Ni Ti-6Al-4V Unknown TiNi3
718 718

5.4 Discussion

Usually, the formation of cracking can be discussed in terms of metallurgical and

mechanical factors and previous research has highlighted two factors that could
influence the formation of cracks within a Ti/Ni or Ti alloy/Ni alloy weld. Firstly, two
intermetallic brittle phases, Ti2Ni and TiNi3, which are readily produced within the weld
during welding of a titanium alloy and a nickel alloy, can increase the susceptibility to
failure at relatively low stresses. Secondly, the large differences of thermo-physical
properties between Ti-6Al-4V and Inconel 718 can generate the stresses that actually
cause the formation of cracks within the weld.

In order to clearly realise the relevance between processing parameters and the weld
quality particularly in the formation of intermetallic brittle phases and cracks in welds, a
simple two-dimensional analytical model was developed and compared with the
experimental results. The model focuses on the relationships between laser power,
welding speed and the laser beam offset position and the melt pool properties and
behaviours, including the melt area, melt ratio and cooling rate.

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5.4.1 Development of analytical model for welding dissimilar materials

The thermal distribution in both Ti-6Al-4V and Inconel 718 welding plates were
modelled individually according to Rosenthal’s equation for two dimensional flow of
heat [205] as shown in Eq.(5-1):

q ′ −λvx
T ( x, y ) − T0 = e K 0 (λvr ) (5-1)
2πk

where T(x,y) is the temperature at point (x,y) (±C) , T0 is the original sample
temperature (20 ±C), q′ is the rate of heat per unit length (W/m), k is the thermal
conductivity (W/m ±C), l is the thermal diffusivity (m2/s), v is welding speed (m/s), K0
is the modified Bessel function of the second kind and zero order, and r = ( x2 + y2 )1/2 is
the distance from the heat source (m). In order to increase the precision of modelling
results, values of thermal conductivity and thermal diffusivity at around half of the
melting point of each material are used, as shown in Table 5-2 and Figure 5-2. Heat
transfer across the interface of the two welding materials is ignored at this stage. From
results of the thermal distribution, the melt pool size is defined according to the melt
points of Ti-6Al-4V (1655 ºC) and Inconel 718 (1260 ºC) as shown in Figure 5-14(a)
and Figure 5-14(b). In Figure 5-14(a), on the Ti-6Al-4V side, the melt pool width, the
melt pool length in the forward direction and the melt pool length in the rear direction
are presented as WTi, L1Ti and L2Ti, respectively. Meanwhile, in Figure 5-14(b), WNi,
L1Ni and L2Ni are the melt pool width, the melt pool length in the forward direction and
the melt pool length in the rear direction on the Inconel 718 side, respectively.

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Figure 5-14 Schematic diagram of the melt pool calculated according to Rosenthal’s
equation with characteristic dimensions shown: (a) the Ti-6Al-4V side; (b) the Inconel
718 side; (c) the melt pool curve in laser dissimilar materials welding.

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 Chapter 5 – Fibre laser welding of Ti-6Al-4V titanium alloy to Inconel 718 nickel alloy

The Ti-6Al-4V and Inconel 718 melt pools are then taken as ellipses; one ellipse
represents the area in front of the beam axis and one is the area behind the beam axis.
The mean length of the separate melt pools in the forward and rear directions are
( L1Ti + L1 Ni ) ( L + L2 Ni )
L1 = and L2 = 2 Ti , respectively. This is taken as the length of the
2 2
combined pool. The melt pool is described in Eq.(5-2) to Eq.(5-5) and Figure 5-14(c).

On the Ti-6Al-4V side (y > 0):

x2 y2
For x > 0: 2
+ =1 (5-2)
L1 WTi2

x2 y2
For x < 0: 2
+ =1 (5-3)
L2 WTi2

On the Inconel 718 side (y < 0):

x2 y2
For x > 0: 2
+ =1 (5-4)
L1 W Ni2

x2 y2
For x < 0: 2
+ =1 (5-5)
L2 W Ni2

After that, the melt pool area (mm2), the melt ratio and the cooling rate of the melt pool
(°C/sec) [206] are calculated as shown in Eq.(5-6) to Eq.(5-9), respectively.

The melt pool area:

π × WTi ( L1 + L2 ) π × W Ni ( L1 + L2 )
Melt pool area = + (5-6)
4 4

The melt ratio:

π × W Ni ( L1 + L2 ) (5-7)
× t Ni
V Ni 4
Melt ratio = =
VTi π × WTi ( L1 + L2 )
× t Ti
4

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 Chapter 5 – Fibre laser welding of Ti-6Al-4V titanium alloy to Inconel 718 nickel alloy

where VNi, VTi, tNi and tTi are the melt volume in the Inconel 718 side and Ti-6Al-4V
side, and the thickness of the Inconel 718 plates and Ti-6Al-4V, respectively. Due to the
same thickness of the Ti-6Al-4V and Inconel 718 plates, Eq.(5-7) can be rewritten to
Eq.(5-8).

WNi
Melt ratio = (5-8)
WTi

•
According to Hofmeister et al. [206], the cooling rate ( T ) of a melt pool for conductive
cooling is related to the length of the pool by an expression of the form as shown in
Eq.(5-9).

•
Log ( T ) = − 2 × log( L1 + L2 ) + 3 (5-9)

From the models it can be seen that when laser power and welding speed are kept at
1000 W and 80 mm/s, a longer and bigger melt pool is obtained when the laser beam is
offset to the Inconel 718 side as shown in Figure 5-15(a). Similar lengths of melt pool
are obtained whether the laser beam is positioned on the interface or offset to the Ti-
6Al-4V side. The melt area in the Inconel 718 side is slightly wider than in the Ti-6Al-
4V side when the laser beam is positioned on the interface. A bigger melt area is found
in the Ti-6Al-4V side when the laser beam is offset to the Ti-6Al-4V side. Similar
trends are observed when welding speed increases from 80 mm/s to 100 mm/s, as
shown in Figure 5-15(b).

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 Chapter 5 – Fibre laser welding of Ti-6Al-4V titanium alloy to Inconel 718 nickel alloy

(a)

(b)

Figure 5-15 The melt pool curves in fibre laser welding of Ti-6Al-4V to Inconel 718
obtained at three different laser beam offset positions with: (a) 1000 W and 80 mm/s; (b)
1000 W and 100 mm/s.

Relationships between the formation of cracks and the melt pool behaviours including
the melt pool area, the melt ratio and the cooling rate are shown in Table 5-6 and Figure
5-16. The cooling rate increases as the melt pool area decreases. There is a higher
possibility to produce crack-free welds when the melt pool area and the cooling rate are
in the range of 0.45 ~ 0.95 mm2 and 1142 ~ 3423 °C/sec, respectively. Welds with
cracks readily occur when the heat input is higher than 16 J/mm or lower than 9 J/mm.
Crack-free welds with the smaller melt ratio (symbol “○”) are observed within a very

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 Chapter 5 – Fibre laser welding of Ti-6Al-4V titanium alloy to Inconel 718 nickel alloy

small processing window (the melt pool area and the cooling rate are around 0.44 ~ 0.59
mm2 and 1946 ~ 2930 °C/sec, respectively.).

Table 5-6 Detailed values from the analytical modeling with different welding
parameters.
Speed Power Laser Heat Melt Melt Cooling The formation
(mm/s) (W) beam input pool area ratio rate of crack
positionf (J/mm) (mm2) (°C/sec)

60 700 Centre 11.67 0.39 1.27 4625.16 Crack

60 1000 Ti side 16.67 0.95 0.70 1021.69 Crack
60 1000 Centre 16.67 1.05 1.60 1061.33 Crack
60 1000 Ni side 16.67 1.35 3.60 689.49 Crack
80 700 Centre 8.75 0.21 1.26 9268.46 Crack
80 800 Centre 10.00 0.30 1.46 5405.81 Crack-free
80 900 Ti side 11.25 0.44 0.55 2930.67 Crack-free
80 900 Centre 11.25 0.46 1.25 3139.48 Crack-free
80 900 Ni side 11.25 0.59 2.98 1860.01 Crack-free
80 1000 Ti side 12.50 0.59 0.56 1946.53 Crack-free
80 1000 Centre 12.50 0.64 1.24 1983.89 Crack-free
80 1000 Ni side 12.50 0.83 2.76 1142.71 Crack-free
100 700 Centre 7.00 0.16 1.24 10711.01 Crack-free
100 800 Centre 8.00 0.21 1.21 7800.86 Crack-free
100 800 Ti side 8.00 0.22 0.50 6055.21 Crack
100 800 Ni side 8.00 0.24 4.16 5267.95 Crack
100 900 Centre 9.00 0.29 1.26 5077.40 Crack-free
100 1000 Centre 10.00 0.39 1.30 3423.88 Crack-free
100 1000 Ni side 10.00 0.51 2.85 1933.88 Crack-free

f
Centre, Ti side and Ni side mean the laser beam was positioned on the interface, offset
to the Ti-6Al-4V side and offset to the Inconel 718 side, respectively.

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 Chapter 5 – Fibre laser welding of Ti-6Al-4V titanium alloy to Inconel 718 nickel alloy

Figure 5-16 Relationships between the formation of crack, the melt pool area, melt ratio
and cooling rate in analytical modelling.

5.4.2 Discussion on experimental results

The amount of heat input can determine chemical composition of the weld [207]. It also
determines the cooling rate, which is inversely proportional to the square of the melt
pool length [206]. Thermal strains caused by high cooling rates can increase the crack
initiation rate, but a higher thermal gradient resulting in a rapid cooling rate in the weld
can also reduce the grain size to increase solidification crack resistance [70].
Additionally, a rapid cooling rate may induce nonequilibrium solidification in the weld
and thus amount of segregation in the solidified pool [132, 190]. Modelled results
(Figure 5-16) indicate that crack-free welds were produced at a wide range of cooling
rates so together these effects do not seem to be dominant in determining if cracking
will occur.

The melting ratio of fused materials is another important factor that determines the
formation of defects in dissimilar materials welds [190]. Producing a bond similar to a
brazed joint, by melting one material to induce another one to melt, has been suggested
as a method to avoid the formation of intermetallic phases within the weld [12]. Perhaps
because of this mechanism, the majority of the crack-free welds were produced at a
higher melt ratio in Figure 5-16. It is possible that when the beam was positioned on the

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Inconel 718 side, the lower melting point and higher thermal conductivity of Inconel
718 meant the heat could dissipate more quickly and less material melted than when on
the Ti-6Al-4V side. The smaller hardness variation that occurred (Figure 5-11(c))
indicates less formation of Ti-Ni intermetallics, which can increase strength and
hardness but decrease ductility. Due to Ti-6Al-4V having a lower thermal conductivity
than Inconel 718, when the laser beam was offset to the Ti-6Al-4V side, more heat
could accumulate in the Ti-6Al-4V plate. This could have caused a larger thermal
gradient and hence a strong Marangoni fluid flow, assisting the formation of the brittle
intermetallic phases and increasing hardness variations, as indicated in Figure 5-11 (b).

For optimum properties it is important to avoid the formation of these intermetallic

phases in the welds [101, 208]. In this case it is possible to achieve this by appropriately
restricting the size and extent of the melt pool and the solidification time. When the
laser beam is offset to the Inconel 718 side, the significant reduction of the melt area in
the Ti-6Al-4V side and the wider melt area in the Inconel 718 side (Figure 5-15) may
cause less vigorous convective flow in the molten zone around the keyhole, avoiding
the formation of intermetallic phases in the weld because most of heat input can be lost
quickly on the Inconel 718 side before enough heat is transferred into the Ti-6Al-4V
side to induce severe microsegregation [29].

Crack-free welds were also readily observed at higher welding speed, as shown in
Figure 5-8. Although higher speed is normally related to higher cooling rate, any direct
relation between cracking and cooling rate has already been considered. It is therefore
likely that other factors apart from cooling rate and intermetallic formation played a
secondary role in determining the final state of a weld. Melt pool geometry (elongation
at higher speeds), keyhole geometry and stability effects and slight difference beam
absorption at different traverse velocities may have contributed.

5.5 Conclusions

The effects of three processing parameters, laser power, welding speed and offset
distance of the laser beam from the interface, were investigated individually during fibre
laser welding of Ti-6Al-4V to Inconel 718. Experimental results indicated that when
welding 2 mm thick sheets of Ti-6Al-4V to Inconel 718 with an IPG 1 kW fibre laser,
crack-free welds could be readily obtained at higher laser powers and welding speeds. A

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better quality weld with less hardness variations and less chance of cracks can be
generated by offsetting the laser beam approximately 35 µm from the interface to the
Inconel 718 side and using a combination of a higher laser power and a higher welding
speed. This is attributed to this method suppressing the formation of Ti-Ni intermetallic
brittle phases. According to results from analytical modelling, properly controlling the
heat input, the melt pool area and the cooling rate at the same time can readily result in
a crack-free weld.

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 Chapter 6 – Pulsed wave fibre laser welding of Zn-coated steel to Al alloy

Chapter 6
Pulsed wave fibre laser welding of Zn-
coated steel to Al alloy

6.1 Introduction

Reducing vehicle weight has become an important task in the automotive industry due
to the latest policies related to environmental protections and avoid unnecessary energy
waste, issued by the European Union [209]. Reducing the weight of a vehicle by 10%
can increase its fuel economy by 6% [210]. Figure 6-1 presents the relationship between
the vehicle weight and the fuel efficiency. For these reasons, as shown in Figure 6-2,
replacing steels with lightweight materials, such as Al alloys, Mg alloys or composite
materials, on car bodies is considered as a potential approach to meet latest
environmental laws [70]. Al alloys is another common material used on car bodies in
order to reduce vehicle weight.

In certain areas of car manufacture, organic glues are often used for temporally joining
the Zn-coated steels and Al alloys before permanent welding. The stability of such
temporary joining by glues needs improving. Laser “stitching” or low strength welding
could be considered as an alternative. Some challenges exist in joining Zn-coated steel
to Al alloy by laser welding, due to the material properties of the weld materials being
significantly different. Porosity, spatter and intermetallic brittle phases are readily
produced in laser lap welding of Zn-coated steel to Al alloy [211].

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 Chapter 6 – Pulsed wave fibre laser welding of Zn-coated steel to Al alloy

Figure 6-1 Relationship between the vehicle weight and fuel efficiency [212].

Figure 6-2 The use of lightweight materials on car bodies in: (a) 1977; (b) 2006. (Glass,
rubber and ceramic materials are included in the “Other” category.) [210].

Currently, laser lap welding of Zn-coated steels is widely applied in the automotive
industry [160]. Due to the low boiling point of the Zn (approximately 907 °C), the
challenge in this process is that Zn vaporisation from the coating occurs before the steel
has reached its melting point (around 1536 °C). According to Figure 6-3, Zn vapour is
able to cause relatively high pressure within the keyhole during the welding process and
induce violent fluid flow in the melt pool. Accordingly, the formation of defects, such
as blisters, craters and open cavities are easily obtained in this process, which can affect
the weld quality [161-164]. Tzeng [213] pointed out that a good weld appearance could
be obtained under appropriate pulsed welding conditions. However, the formation of
porosity in the weld was not easily avoided. In some previous work, the weld quality
has been continuously developed and improved using several technologies, such as

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using a bifocal hybrid laser system to stabilise the melt pool [166], pre-drilling vent
holes along the weld line [214, 215] or keeping a gap between two welding materials
[161, 165, 216] to allow Zn vapour to escape through the melting zone. A filler material
was also used to improve the weld quality in laser welding of steel to Al alloy [217,
218].

Figure 6-3 The relationship between vapour pressure of pure metals and temperature
[219].

The challenge in laser welding of Al alloy to steel is to account for the differences in
physical and chemical properties between two materials and their poor miscibility,
which result in the formation of intermetallic brittle phases and an interfacial reaction
layer in the weld [127, 218]. According to the Al-Fe phase diagram as shown in Figure
6-4, five intermetallic phases, Fe3Al, FeAl, FeAl2, Fe2Al5 and FeAl3, can be formed at
different chemical compositions. Amongst these, Fe-rich intermetallic phases, FeAl and
Fe3Al, have higher ductility and toughness while Al-rich intermetallic phases, FeAl3,
FeAl2 and Fe2Al5, are harder and more brittle [134, 220]. The properties of these
intermetallic phases are tabulated in Table 6-1. During the rapid heating and cooling
process, FeAl3 and Fe2Al5 are the principle phases that occur in the weld [221]. Fe2Al5
is considered as extremely brittle and its hardness can be above 1000 Hv [220].

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Figure 6-4 Fe-Al equilibrium diagram [194]

Table 6-1 Classifications of intermetallic compounds in the Fe-Al binary system [220].
Type wt.% of Fe Hardness (Hv) Compressive strength (MPa)
Fe3Al 80.06 330 560
FeAl 67.31 470 670
FeAl2 50.72 Unknown Unknown
Fe2Al5 45.16 1013 240
FeAl3 40.70 892 200

In general, the thickness of the interfacial reaction layer is a key factor in determining
the weld quality in laser welding of un-coated steel to Al alloy [220]. Katayama [134]
found that a crack-free weld was produced when the melt area of Al alloy was
controlled appropriately. To consider the weld geometry and material properties of
welding materials, Lee et al. [222-224] applied a defocused laser beam in laser lap
welding of dissimilar materials. The laser beam only irradiated and melted the top
material; the bottom one was melted via the heat conducted from the top material. Sierra
et al. [127] predicted that the weld quality was determined by the weld penetration
depth in the bottom material. The deeper the penetration depth was, the thicker the
interfacial reaction layer produced, resulting in a slight hardness variation and a lower

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shear force. Meanwhile, Borrisutthekul et al. [130] also found that the weld strength
increased when the interfacial reaction layer was reduced in laser lap welding of Al
alloy to steel. Their results showed that increasing welding speed and thermal
conductivity of the backing block, which was put below the bottom welding material,
could effectively reduce the molten time and control the heat flow; therefore, the
diffusion in the weld could be restricted.

Compared with continuous wave laser welding, pulsed wave laser welding generally
can precisely control the heat input by means of properly selecting laser power, welding
speed, spot size of the laser beam, the wave shape, pulse frequency, pulse duration and
pulse repetition rate [134]. For this reason, pulsed wave laser welding was applied in
this preliminary study, in which the process temperature has to be paid more attention
due to the lower melting point of Zn. The aim of this study was to investigate the
relationship between the weld quality and processing parameters, such as laser power,
welding speed and the pulse frequency, in the lap welding of Zn-coated steel on Al
alloy using a single mode fibre laser. A series of analyses, which included mechanical
tests and metallurgical evaluations, were carried out.

6.2 Experimental procedure

A sheet of 0.7 mm thick DX54 Zn-coated steel (the top material) and 2.5 mm thick EN-
AW-5754 Al alloy (the bottom material) were lap welded in this study. Chemical
composition of these two materials is listed in Table 6-2. Physical properties of Fe, Al
and Zn elements are listed in Table 6-3. The experimental setup of the fibre laser
welding system is shown in Chapter 4. Before welding, both materials were cleaned
with acetone and clamped with a custom-made welding fixture, as shown in Figure
4-14(b), to minimise the gap between two weld materials. During welding, argon was
supplied co-axially as the shielding gas. The shielding gas pressure and flow rate were
kept at 2 bar and 25 l/min, respectively.

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Table 6-2 Chemical composition (wt.%) of DX54 Zn-coated steel and EN-AW-5754 Al
alloy.
DX54 Zn-coated steel 5754 Al alloy
C 0.0028 -
Si 0.0030 0.4
Mn 0.1300 0.5
P 0.0080 -
S 0.0180 -
Mg - 2.6-3.2
Cr 0.0180 -
Mo 0.0010 -
Ni 0.0180 -
Ti 0.0520 -
N 0.0029 -
Al 0.0430 Balance
Fe Balance 0.4

Table 6-3 Physical properties of Fe, Al and Zn elements [24].

Fe Al Zn
Melting point (±C) 1538.0 660.1 419.5
Boiling point (±C) 2861 2520 907
3
Density (g/cm ) 7.87 2.70 7.14
Specific heat (J/kg K) 449 897 388
Thermal conductivity (W/m K ) 80 238 113
Coefficient of expansion (10-6/K) 12.1 23.5 30.2

Three parameters chosen were as variables in this long-pulsed welding study. They
were laser power, welding speed and pulse frequency. The control of pulse frequency
was achieved by setting a constant “off” time (t2 in Eq.(2-4) at Page 55) of 10 ms. In
effect this means that greater effective power is decreased at lower frequency. The
detailed experimental matrix is shown in Table 6-4. Firstly, the influence of laser power
was investigated when welding speed and the pulse frequency were fixed. Secondly,
welding speed was increased from 50 mm/s to 100 mm/s when laser power was kept at

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300 W with different pulse frequencies. Finally, the effect of pulse frequency with
constant laser power and welding speed was studied. Because the aim of this work is
investigating the keyhole mode welding in fibre laser welding of Zn-coated steel on Al
alloy, the focal position of laser beam was set on the top surface of the Zn-coated steel
through all the work.

Table 6-4 Experimental parameter matrix.

Power (W) 250 275 300
Speed (mm/s) 50 70 100 50 70 100
1.64 1.64
1.76 1.76
1.96 1.96
2.17 2.17
2.44 2.44 2.44 2.44
Frequency (Hz)
2.78 2.78
3.23 3.23 3.23 3.23
3.85 3.85 3.85 3.85
4.76 4.76 4.76 4.76
6.25 6.25 6.25 6.25

After welding, the weld appearance was observed with optical microscopy. Later, welds
were sectioned, mounted, ground, polished with diamond slurry to 3 µm surface finish
and etched. Nital acid (3% HNO3 ethanol solution) and Keller’s reagent were used to
etch Zn-coated steel and Al alloy, respectively, for further evaluations.

The weld shape and defects were measured and observed using optical microscopy.
Hardness profiles in three different penetration depth levels, 250, 500 and 750 µm, were
carried out using a microhardness testing machine with a 50 g loading force and the
testing duration of 10 seconds. Figure 6-5 shows the hardness testing method applied
here. According to the standard of BS EN 10002-1: 2001 [181], the shear force test was
carried out at room temperature using an universal testing machine INSTRON 4507
operating with a crosshead speed of 1 mm/s. The dimensions of shear force testing
sample are schematically illustrated in Figure 6-6. Three welding lines were carried
with specific distances in each sample. After the shear force test, the fracture surface

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was observed using scanning electron microscopy. Microsegregation within welds was
investigated using a scanning electron microscopy equipped with energy dispersive
spectrometer (EDS).

Figure 6-5 Schematic diagram of hardness test.

Figure 6-6 Schematic diagram of shear force testing sample.

6.3 Results
6.3.1 The weld morphology and beam geometry

The weld appearances produced at different welding conditions are shown in Figure
6-7. In Figure 6-7(a), clear blowholes were observed when the laser power, welding
speed and pulse frequency were 250 W, 70 mm/s and 6.25 Hz, respectively. On the
other hand, a smooth weld appearance without any blowhole was produced at 300 W
power, 100 mm/s speed and 1.64 Hz pulse frequency.

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 Chapter 6 – Pulsed wave fibre laser welding of Zn-coated steel to Al alloy

Figure 6-7 Surface weld appearance of long pulsed-wave fibre laser lap welding of Zn-
coated steel on Al alloy: (a) 250 W, 70 mm/s and 6.25 Hz; (b) 300 W, 100 mm/s and
1.64 Hz.

Weld sections of Zn-coated steel on Al alloy produced by a fibre laser with different
welding conditions are shown in Figure 6-8. All welds were with the parallel shape on
the steel side, while the weld widths in the Al alloy were significantly wider. Noticeable
undercut welds were found in the samples shown in Figure 6-8(a) and Figure 6-8(f).
Porosity also appeared at the adjacent area of steel and Al alloy, as shown in Figure
6-8(a).

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Figure 6-8 Macrostructure of fibre laser welding of Zn-coated steel on Al alloy: (a) 250
W, 50 mm/s and 6.25 Hz; (b) 250 W, 50 mm/s and 4.76 Hz; (c) 250 W, 50 mm/s and
3.85 Hz; (d) 250 W, 50 mm/s and 3.23 Hz; (e) 250 W, 70 mm/s and 3.85 Hz; (f) 300 W,
50 mm/s and 3.85 Hz

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The relationship between the weld depth, laser power and pulse frequency is shown in
Figure 6-9(a). When the welding speed was kept at 50 mm/s, the weld depth increased
as the laser power increased at a specific pulse frequency. If the welding speed and laser
power were kept at 250 W and 50 mm/s, respectively, the weld depth decreased with the
increasing pulse frequency. A similar trend also could be obtained at welding conditions
of 300 W and 50 mm/s: the weld depth decreased with increasing pulse frequency but
by a smaller amount. Figure 6-9(b) indicates the influence of welding speed and pulse
frequency on weld depth at a constant laser power of 300 W. At fixed laser power and
pulse frequency, the weld depth could be increased when the welds speed was decreased
from 70 mm/s to 50 mm/s. In addition, the weld depth was slightly increased with an
increasing pulse frequency.

(a)

(b)
Figure 6-9 The weld depth at: (a) welding speed of 50 mm/s; (b) laser power of 300 W.

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 Chapter 6 – Pulsed wave fibre laser welding of Zn-coated steel to Al alloy

In general, increasing the pulse frequency will reduce the power dekivery by the laser
(Page 129). This would suggest that a greater weld depth will be achieved at lower
frequency. This is generally true for a single material radiate by laser. However, for the
experiments described here which involve two materials in contact, the depth of
penetration is also gonverned by the thermalconducity of both materials. Sicnce the
thermalconductility of Al alloy is higher than that of Zn-coated steel. Sufficient heat is
conductive to the bottom of the weld to allow it to melt even the relative lower power.

6.3.2 Hardness distribution

Figure 6-10 compares hardness distributions at different laser powers, 250 W and 300
W, when welding speed and pulse frequency were fixed at 50 mm/s and 3.85 Hz. In
Figure 6-10(a), hardness profiles at 250 µm and 500 µm penetration levels are similar.
The highest hardness of both them occurred at the centre of FZ, while the mean
hardness of Zn-coated steel was approximately 95.1 Hv. The hardness was increased
sharply within the FZ at the 750 µm penetration level. A similar distribution is found in
Figure 6-10(b). Hardness near the FZ was increased with the penetration depth
increased at the welding conditions of 300 W, 50 mm/s and 3.85 Hz. The higher
hardness was also obtained near the FZ at the 750 µm penetration level.

(a)

Figure 6-10 Microhardness distribution at three penetration levels in the weld at 50

mm/s and 3.85 Hz with: (a) 250W; (b) 300 W.

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 Chapter 6 – Pulsed wave fibre laser welding of Zn-coated steel to Al alloy

(b)

Figure 6-10 (Continued).

A comparison of hardness distributions in samples produced using two different pulse

frequencies with the same laser power and welding speed are shown in Figure 6-11. In
Figure 6-11(a), the hardness at three penetration levels of 250 µm, 500 µm and 750 µm
were slightly increased in the FZ in comparison to the mean hardness of the base
materials. The higher hardness was obtained in the centre of FZ at the penetration depth
level of 250 µm. Figure 6-11(b) shows hardness profiles at a lower pulse frequency,
3.23 Hz, with the same laser power and welding speed as in Figure 6-11(a). At the 250
µm penetration level, hardness increased in the FZ. At the 500 µm penetration level,
hardness in the centre of FZ was higher in comparison with that of the 250 µm
penetration level. Meanwhile, the hardness was abruptly increased in the FZ at the 750
µm penetration level.

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 Chapter 6 – Pulsed wave fibre laser welding of Zn-coated steel to Al alloy

(a)

(b)

Figure 6-11 Microhardness distribution at three penetration levels in the weld at 300 W
and 70 mm/s with: (a) 3.85 Hz; (b) 3.23 Hz.

6.3.3 Micro-segregation

When the welding conditions were 250 W, 50 mm/s and 3.23 Hz, the cross-section of
weld observed by an optical microscope and a scanning electron microscope are shown

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in Figure 6-8(d) and Figure 6-12, respectively. In Figure 6-12(a), Fe, Al and Zn
elements can be readily identified according to their colour in a backscattered electron
image. Therefore, it was easy to distinguish the location of Fe, Al and Zn elements after
the melting, mixing and solidification processes. Following the colour rule in the
backscattered image, severe mixing in the interfacial reaction layer occurred on the Al
alloy side. A magnified image of Figure 6-12(a) is illustrated in Figure 6-12(b). Not
only cracking, but also multi-phases, appeared clearly. The hardness and chemical
compositions of Points A to F are tabulated in Table 6-5. Amongst these, a higher
hardness was obtained in Point E and its phase could be identified as intermetallic
brittle phase FeAl3, according to the Al-Fe phase diagram [194] and EDS results.

Figure 6-12 Backscattered electron images and EDS results of the weld produced at 250
W, 50 mm/s and 3.23 Hz: (a) microstructure (X400); (b) the magnified image of Section
1 in Figure 9(a) (X1000); (c) the distribution of Fe element; (b) the distribution of Al
element.

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Table 6-5 Hardness and chemical compositions of Points A-F in Figure 6-12(b).
Point Hardness Chemical composition (wt%) Phase
(Hv) Al Fe Zn
A 204.9 65.44 34.56 - FeAl3
B 173.9 7.72 92.28 - Rich-Fe
C 153.3 8.12 91.88 - Rich-Fe
D 164.7 91.35 8.65 - Rich-Al
E 630.4 63.69 35.74 0.57 FeAl3
F 132.6 100.00 - - Rich-Al

Figure 6-12(c) and Figure 6-12(d) indicate the distribution of Fe and Al elements,
respectively. They show that steel could penetrate into the Al alloy side, while a part of
Al alloy disappeared and was replaced by steel. Also, the thickness of interfacial
reaction layer was increased with an increasing penetration depth as shown in Figure
6-12(c) and Figure 6-12(d).

6.3.4 Shear force

Results of shear force tests are shown in Figure 6-13. When laser power and welding
speed were kept at 250 W and 50 mm/s, respectively, decreasing pulse frequency could
increase the shear force. The shear force was increased when laser power was increased
from 250 W to 300 W at 50 mm/s and 3.23 Hz. However, the shear force could be
reduced by increasing welding speed from 50 mm/s to 70 mm/s when laser power and
welding speed were kept at 300 W and 3.23 Hz. Figure 6-14 shows fracture surfaces on
the Zn-coated steel and Al alloy sides when the welding conditions were 300 W, 70
mm/s and 3.23 Hz. It indicates that the failure was caused by brittle fracture.

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Figure 6-13 Shear force of fibre laser lap welding of Zn-coated steel to Al alloy.

Figure 6-14 Fracture surfaces of the shear force testing sample at 300 W, 70 mm/s and
3.23 Hz: (a) on the Al alloy side; (b) on the Zn-coated steel side.

6.4 Discussion

One of challenges in laser welding of Zn-coated steel on Al alloy is that the melt pool is
unstable due to the pressure from Zn vaporisation, resulting in porosity and blowholes
as shown in Figure 6-7(a). Most porosity was found near the mating surface between
Zn-coated steel and Al alloy (Figure 6-8(a)) because the higher cooling rate of fibre
laser welding could not provide enough time for porosity to exit from the keyhole
before the melting zone solidified. Generally, heat input, which is a function of welding
speed, laser power and pulse frequency, plays a key role in determining the weld

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geometry, including the weld width and penetration depth. Hence, however, the effect
of high thermal conductivity of Al alloy has to be especially considered in laser keyhole
mode welding. In this project, therefore, the weld width on the steel side was narrower
than the weld width on the Al alloy side; the weld penetration depth on the Al alloy side
was more sensitive to the pulse frequency at lower power and lower welding speed,
such as 250 W and 50 mm/s (Figure 6-9(a)). Conversely, welding speed and laser power
become more important in determining the weld geometry (Figure 6-9(b)) when the
total heat input was higher.

The formation of intermetallic phases is another key issue in welding of dissimilar

materials because this can increase hardness and reduce the toughness of the weld.
Generally, according to the Al-Fe phase diagram, the formation of intermetallic phases
can occur when the amount of Al element exceeds approximately 12 wt% [194, 225].
Depending on chemical composition, intermetallic phases in the Al-Fe binary system
can be classified into five types, FeAl, Fe3Al, FeAl3, FeAl2 and Fe2Al5. Amongst these,
FeAl3 and Fe2Al5 with higher hardness are recognised as being easily obtained in the
weld [220] (Figure 6-12 and Table 6-5). Accordingly, the hardness in the FZ
significantly increases on the Al alloy side (Figure 6-10 and Figure 6-11) because of the
occurrence of these intermetallic brittle phases within the weld. The phases mean the
weld ductility can be reduced which causes the brittle fracture (Figure 6-14).

As well as the formation of intermetallic phases, the thickness of interfacial reaction

layer is important in laser welding of dissimilar materials. In general, the thicker the
interfacial reaction layer is, the higher the possibility that intermetallic phases occur
[130]. According to Figure 6-12(c) and Figure 6-12(d), the thickness of interfacial
reaction layer has a positive correlation with the penetration depth. The thickness of the
interfacial reaction layer will increase as the penetration depth increases. For this
reason, the formation of cracks was easily observed near the bottom area of the FZ on
the Al alloy side (Figure 6-8(d) and Figure 6-12(b)).

The mechanical properties of the weld are also related to the weld geometry as well as
micro-segregation within the weld. It can be said that appropriate weld dimensions,
including sufficient penetration depth in the Al alloy, are factors that determine the weld
shear force especially in the laser micro welding process. Hence, precisely controlling
heat input to increase the weld penetration depth (or the ratio of weld width to depth)

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can be an effective method to improve the weld shear force in fibre laser welding Zn-
coated steel to Al alloy (Figure 6-13).

6.5 Conclusions

Laser welding of Zn-coated steel to Al alloy was carried out using an IPG 1 kW single
mode fibre laser. Narrower and more parallel-shaped keyhole welds can be produced
with the laser in the pulsed wave welding mode. Results show that the weld hardness in
the FZ on the Al alloy side can be increased significantly due to the formation of
intermetallic phases. Pulse frequency is a key factor, especially at the lower heat input
welding condition, but laser power and welding speed are important in determining the
weld geometry when the heat input is higher than a critical value. Brittle intermetallic
phases were readily found in the FZ on the Al alloy side. All welds broke with a brittle
fracture in the shear force test.

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Chapter 7
Continuous wave fibre laser welding of
Zn-coated steel to Al alloy

7.1 Introduction

As aforementioned, laser welding is used in the automotive industry because of its

characteristics of high beam quality, small HAZ and high repeatability. Until now, most
research work associated with laser welding of automotive materials has addressed the
lap welding of either Zn-coated steels [160, 166, 216, 226] or un-coated steel to Al alloy
[82, 127, 217, 220, 222-225, 227]. Hardly any work has been found on laser welding of
Zn-coated steel to Al alloy without using a filler material. In laser lap welding of Zn-
coated steel to Al alloy, common issues include the formation of porosity, spatter and
intermetallic phases and the unstable melt pool resulting from the lower boiling point of
Zn (907 °C) compared with the melting point of Fe (1538 °C) and the boiling point of
Al (2520 °C). These were also described in the previous chapter in the context of gap-
free pulsed wave fibre laser welding of Zn-coated steel to Al alloy [120, 211].

In laser welding, the shielding gas is another important factor that influences the weld
quality. One of its roles is to protect the melt pool from oxidation [45]. Theoretically, an
inert shielding gas with a higher density can provide better protection over the melt pool.
In laser welding, the important properties of the shielding gas include its physical
properties, chemical composition, the flow rate and distribution [65]. Regarding the
physical properties of a shielding gas, its thermal conductivity and density have been
found to be more important than the ionisation potential in the high energy density laser
welding process [103].

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The purpose of the study described in this chapter was to investigate the feasibility and
characteristics of using laser welding of Zn-coated steel and Al alloys as a “stitching”
method to replace the use of organic glues in the automotive industry. The relationship
between the welding quality and processing parameters, such as welding speed, laser
power, the number of welding passes and type of the shielding gas, have been
investigated in laser lap welding of Zn-coated steel on Al alloy using a single mode
continuous wave fibre laser. A series of analyses including metallurgical evaluations,
mechanical tests were carried out. A particular focus of this chapter was to achieve a
sound weld that can meet the industrial requirement for “stitching” applications, in
terms of weld strength, surface finish, micro-segregation and hardness variations.

7.2 Materials and methods

The experimental materials consisted of 1.0 mm thick DX54 Zn-coated steel and 1.0
mm thick EN-AW-5754 Al alloy with the Al alloy being placed at the bottom part of
the lap welding to avoid the high reflection to the laser beam by the Al alloy. The Zn
layer, approximately 10 µm thick, was coated on both sides of the steel. Chemical
compositions of DX54 Zn-coated steel and EN-AW-5754 Al alloy are tabulated in
Table 6-2. Physical properties of Fe, Al and Zn are listed in Table 6-3. IPG 1kW single
mode continuous wave fibre laser with a wavelength of 1070 nm and a single axis high
speed liner motor controlled stage were applied in this work. The spot diameter of the
focused fibre laser beam was approximately 72 µm. The full setup is shown in Figure
4-11. Before welding, both materials were cleaned with acetone and clamped with a
custom-made welding fixture, see Figure 4-14(b), to minimise the gap between the two
weld materials. During welding, Ar or N2 gas was supplied co-axially as the shielding
gas with a constant flow rate of 25 l/min and pressure of 2 bar.

Four parameters were chosen as variables in this study. They were laser power, welding
speed, the number of welding passes and type of shielding gas. Properties of the Ar and
N2 shielding gases are listed in Table 2-4. Detailed experimental matrices of the single
pass and double pass welding results are shown in Table 7-1 and Table 7-2, respectively.
Comparative work with the two different shielding gases was carried out in both the
single pass and double pass welding cases. Firstly, the influence of laser power was
investigated at a constant welding speed of 100 mm/s in single pass welding. Secondly,

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three combinations of the first pass welding conditions with a focused laser beam were
chosen as listed in Table 7-2 and a defocused laser beam was applied in a second pass
welding. The laser power of the second pass weld was varied from 150 W to 250 W,
while the welding speed and the focal point position of the laser beam (f.p.p.) were kept
at 75 mm/s and 2 mm above the top surface of Zn-coated steel, respectively. The spot
size of the defocused laser beam was approximately 81 µm.

Table 7-1 Experimental matrix of the single pass welding.

Shielding Welding speed Power Maximum porosity
gas (mm/s) (W) diameterg (µm)
400 W -
450 W -
500 W -
Ar 100
550 W -
600 W 112.3
650 W -
400 W X
450 W X
500 W 91.6
N2 100
550 W -
600 W -
650 W 29.0

g
‘-’ and ‘X’ mean porosity was not observed in the weld and materials were not welded
successfully, respectively.

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Table 7-2 Experimental design of the double pass welding.

Welding conditions Maximum porosity
Shielding gas
1st pass 2nd pass diameterh (µm)
150 W 90.6
600 W &
200 W 65.0
75 mm/s
250 W 40.5
150 W -
600 W &
Ar 200 W -
100 mm/s
250 W -
150 W 48.2
650 W &
200 W 44.5
100 mm/s
250 W 26.9
150 W -
600 W &
200 W -
75 mm/s
250 W -
150 W 32.9
600 W &
N2 200 W -
100 mm/s
250 W 39.4
150 W -
650 W &
200 W -
100 mm/s
250 W -

For further evaluations, welds were sectioned, mounted, ground, polished with diamond
slurry to 3 µm surface finish and etched with Nital acid (3% HNO3 ethanol solution)
and Keller’s reagent for Zn-coated steel and Al alloy, respectively. The weld geometry
shape, including the weld width and depth, was measured using optical microscopy with
a personal computer running Solution DT. The diameters of pores in welds were
observed and calculated from the cross section of welds, also using optical microscopy.
Hardness profiles at three different penetration depth levels, 250, 750 and 1100 µm,
were measured using a microhardness testing machine with a 50 g loading force and a
testing duration of 10 seconds. Figure 7-1 shows the hardness testing method applied in

h
‘-’ means porosity was not observed in the weld and materials were not welded
successfully, respectively.

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this work. According to the BS EN 10002-1: 2001 standard [181], the shear force test
was carried out at room temperature using an INSTRON 4507 universal tensile testing
machine operating with a crosshead speed of 1 mm/s. The dimensions of a shear force
test sample are schematically illustrated in Figure 7-2. Three welding lines were carried
with specific distances in each sample. After the shear force test, the fracture surfaces
were examined using a scanning electron microscope. Chemical analysis was
undertaken using an energy dispersive spectrometer (EDS).

Figure 7-1 Schematic illustration of the hardness test points and the weld dimension.

Figure 7-2 Schematic illustration of shear force testing sample.

7.3 Results
7.3.1 The weld beam geometry

Figure 7-3 shows the relationship between the weld bead width, penetration depth and
laser power in the single pass welding with different shielding gases at a constant
welding speed of 100 mm/s. In Figure 7-3(a), when Ar gas was used as the shielding
gas, the weld width (WTop-steel indicated in Figure 7-1), abruptly varied when laser power

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was increased from 400 W to 650 W. On the other hand, when N2 gas was used, the
weld width slightly increased when laser power was increased from 400 W to 600 W
before decreasing again at 650 W. With respect to the weld penetration depth on the Al
alloy side (DAl), a clear relationship between the weld depth and laser power were found.
In Figure 7-3(b), the weld depths increased with increasing laser power using both Ar
and N2 as the shielding gas. Furthermore, at the same welding condition, weld depths
with Ar gas were always deeper than those with N2 gas.

(a)

(b)
Figure 7-3 The weld dimension obtained from the single pass welding at 100 mm/s with
different shielding gases: (a) the weld width on the Zn-coated steel side (WTop-Steel); (b)
the weld depth on the Al alloy side (DAl).

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The relationships between welding parameters and the weld dimension with double pass
welding are shown in Figure 7-4 and Figure 7-5. It can be seen that with Ar gas the
weld width (WTop-steel) increased when laser power of the second defocused beam weld
was increased as shown in Figure 7-4(a). However, in Figure 7-4(b) there was no clear
relationship between the weld width and laser power of the second pass weld with N2
gas. Figure 7-5 illustrates the weld macrostructures produced by double pass welding
with different shielding gases. All of them have a near-parallel shape on the steel side
and the widths of the FZ on the Al alloy side are significantly wider than those on the
steel side.

(a)

(b)
Figure 7-4 The weld width of the double pass welding with different shielding gases: (a)
Ar gas; (b) N2 gas.

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Figure 7-5 Macrostructure of double pass welding when the first pass welding was 650
W,100 mm/s, f.p.p. of 0 mm, and accompanied with the second pass welding and
shielding gases were: (a) 150 W, 75 mm/s, f.p.p. of +2 mm and Ar gas; (b) 150 W, 75
mm/s, f.p.p. of +2 mm and N2 gas; (c) 200 W, 75 mm/s, f.p.p. of +2 mm and Ar gas; (d)
200 W, 75 mm/s, f.p.p. of +2 mm and N2 gas; (e) 250 W, 75 mm/s, f.p.p. of +2 mm and
Ar gas; (f) 250 W, 75 mm/s, f.p.p. of +2 mm and N2 gas.

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7.3.2 Weld defects

Table 7-1 and Table 7-2 list the formation of porosity in welds produced by single pass
and double pass welding, respectively. It was observed that porosity occurred randomly
in both the single pass and double pass welding. The pore diameters ranged from 26.9
µm to 112.3 µm. Most of the porosity occurred near the interface between Zn-coated
steel and Al alloy as indicated in Figure 7-5(a), Figure 7-5 (c) and Figure 7-5 (e). Figure
7-6 shows the top surface appearance of the single and double pass welds with Ar and
N2 shielding gases. No matter what type of shielding gas was used, an unstable melt
pool and the formation of spatter could be easily observed in the single pass welds, as
shown in Figure 7-6(a) and Figure 7-6(b). In Figure 7-6(c) and Figure 7-6(d), clearly
improved surfaces were obtained after applying a second pass. With respect to micro-
cracks, they could be easily observed near the interfacial area of the FZ, as shown in
Figure 7-7(a) and Figure 7-7(b).

Figure 7-6 Top appearance of welds produced from: (a) single pass welding with Ar
gas; (b) single pass welding with N2 gas; (c) double pass welding with Ar gas; (d)
double pass welding with N2 gas. (the first pass welding parameters: 650 W, 100 mm/s,
f.p.p. of 0 mm; the second pass welding parameters: 200 W, 75 mm/s, f.p.p. of +2 mm.)

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Figure 7-7 Backscattered electron image of weld cross sections obtained for the double
pass welding with: (a) Ar gas; (b) N2 gas. (the first pass welding parameters: 650 W,
100 mm/s, f.p.p. of 0 mm; the second pass welding parameters: 200 W, 75 mm/s, f.p.p.
of +2 mm.)

7.3.3 Hardness distribution

Hardness distributions at three different penetration levels in single and double pass
welds (the first pass welding parameters: 600 W, 100 mm/s, f.p.p. of 0 mm; the second
pass welding parameters: 250 W, 75 mm/s, f.p.p. of +2) with different shielding gases
are shown in Figure 7-8 and Figure 7-9. In single pass welding, with Ar gas shrouding,
a higher hardness was found near the FZ at the 250 µm penetration level while the mean
hardness of steel was approximately 101.3 Hv as found in Figure 7-8(a). A similar

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hardness distribution was obtained at the 750 µm penetration level. Hardness near the
FZ slightly is increased. Meanwhile, hardness in the centre of the FZ at the 1100 µm
penetration level was higher than that at the 250 µm and 750 µm penetration levels. The
mean hardness in the unaffected zone of Al alloy was around 67.9 Hv. With N2 gas,
there was no significant increase in hardness in the centre of the FZ compared with Al
alloy parent material at the 1100 µm penetration level, as shown in Figure 7-9(a),
indicating the possibility of fewer brittle phases.

(a)

(b)

Figure 7-8 Hardness distribution obtained from single and double passes welding with
Ar gas (the first pass welding parameters: 600 W, 100 mm/s, f.p.p. of 0 mm; the second
pass welding parameters: 250 W, 75 mm/s, f.p.p. of +2 mm): (a) single pass weld; (b)
double pass weld.

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(a)

(b)

Figure 7-9 Hardness distribution obtained from single and double passes welding with
N2 gas (the first pass welding parameters: 600 W, 100 mm/s, f.p.p. of 0 mm; the
second pass welding parameters: 250 W, 75 mm/s, f.p.p. of +2 mm): (a) single pass
weld; (b) double pass weld.

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In double pass welding (Figure 7-8(b), the hardness gradient increased with increasing
penetration depth with Ar gas. In comparison with hardness near the FZ, the hardness at
the 1100 µm penetration level was higher than that measured at the 250 µm and 750 µm
penetration levels. Meanwhile, hardness distributions obtained at the same welding
conditions using N2 as the shielding gas are shown in Figure 7-9(b). Hardness
distributions at the 250 µm and 750 µm penetration levels have similar trends to those
obtained with Ar gas. However, the hardness gradient at the 1100 µm penetration depth
was lower using N2 gas.

7.3.4 Micro-segregation and Microstructure

In double pass welding (the first pass welding parameters: 650 W, 100 mm/s, f.p.p. of 0
mm, and the second pass welding parameters: 200 W, 75 mm/s and f.p.p. of +2 mm),
the weld microstructure produced with Ar and N2 gases are shown in Figure 7-5(c) and
Figure 7-5(d), respectively. The backscattered electron images of Figure 7-5(c) and
Figure 7-5(d) are displayed in Figure 7-7(a) and Figure 7-7(b), respectively. In Figure
7-7(a), a significant amount of steel penetrated into the Al alloy when Ar was applied as
the shielding gas. Less steel was found in the Al alloy side and, more Al diffused into
the steel side when N2 gas was used, as shown in Figure 7-7(b).

SEM backscattered electron images and EDS mapping analysis of Figure 7-5(a) and
Figure 7-5(b) are displayed in Figure 7-10(a) and Figure 7-10(b), respectively. In Figure
7-10(a), the top area of the Al alloy side has been replaced by the steel in the double
pass weld case (the first pass welding parameters: 650 W, 100 mm/s, f.p.p. of 0 mm,
and the second pass welding parameters: 150 W, 75 mm/s, f.p.p. of +2 mm) with Ar gas.
Less steel penetrated into the Al alloy side and more Al was kept in the Al alloy side
when the N2 was used as the shielding gas as shown in Figure 7-10(b).

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Figure 7-10 Backscatter electron images and EDS mapping analysis of double pass
welding (the first pass welding parameters: 650 W, 100 mm/s, f.p.p. of 0 mm; the
second pass welding parameters: 150 W, 75 mm/s, f.p.p. of +2 mm) with different
shielding gases: (a) Ar gas; (b) N2 gas.

Chemical compositions at the 1100 µm penetration level in Figure 7-8(b) and Figure 7-9
(b) are shown in Figure 7-11(a) and Figure 7-11(b), respectively. More Fe was found in
the FZ with Ar gas (Figure 7-11(a)) than with N2 gas (Figure 7-11(b)). At the same time,
more Al element was observed near the FZ in Figure 7-11(b) than in Figure 7-11(a).
The relative amount of Zn element was similar when shrouding with Ar or N2 gas.

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(a)

(b)

Figure 7-11 Relative chemical compositions at the 1100 μm penetration level obtained
from double pass welding (the first pass welding parameters: 600 W, 100 mm/s, f.p.p.
of 0 mm; the second pass welding parameters: 250 W, 75 mm/s, f.p.p. of +2 mm) with
different shielding gases: (a) Ar gas; (b) N2 gas.

Figure 7-12 shows welds produced by double pass welding (the first pass welding
parameters: 600 W, 100 mm/s, f.p.p. of 0 mm, and the second pass welding parameters:
200 W, 75 mm/s, f.p.p. of +2 mm) with different shielding gases. Compared to the weld
produced with N2 gas in Figure 7-12(b), the degree of materials mixing is more
significant when Ar was used as the shielding gas, as shown in Figure 7-12(a). The
hardness and chemical compositions of Points A to I in Figure 7-12 are tabulated in

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Table 7-3. Significant and random variations of chemical composition between Points A
to E were obtained when the weld was shrouded with Ar gas. The higher hardness was
obtained in Point B. Meanwhile, likely Fe3Al and FeAl phases were observed at points
C and D, respectively. When N2 was used as the shielding gas, the amount of Fe
decreased significantly from Points F and G to H and I.

Figure 7-12 Welds produced at double pass welding (the first pass welding parameters:
600 W, 100 mm/s, f.p.p. of 0 mm, and the second pass welding parameters: 200 W, 75
mm/s, f.p.p. of +2 mm) with different shielding gases: (a) with Ar gas; (b) with N2 gas.

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Table 7-3 Hardness and chemical compositions of Points A-I in Figure 7-12.
Point Hardness Chemical compositions (at%)
(Hv) Al Fe Zn
A 199.0 0.92 98.75 0.33
B 281.5 37.85 43.37 18.78
C 147.2 28.11 65.98 5.91
D 170.0 51.83 37.98 10.19
E 242.7 51.45 47.61 0.94
F 179.4 0.74 99.05 0.21
G 157.2 5.85 93.68 0.47
H 148.9 92.50 1.46 6.04
I 207.7 92.75 2.63 4.61

7.3.5 Shear force

The relationship between the shear force and type of shielding gas used is shown in
Figure 7-13. For single pass welding at 600 W and 100 mm/s, the shear force was
higher when using N2 gas than using Ar gas. Similar results appeared in double pass
welding (the first pass welding parameters: 600 W, 80 mm/s, f.p.p. of 0 mm, the second
pass welding parameters: 200 W, 75 mm/s, f.p.p. of +2 mm), a higher shear force was
obtained when N2 gas was used as the shielding gas. Under N2 gas shrouding condition,
both single pass and double pass welds have shear forces satisfying the company
standard (4 kN) with single pass welding marginally passing the minimum industrial
requirement.

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Figure 7-13 Shear force of single pass welds (600 W, 100 mm/s, f.p.p. of 0 mm) and
double pass welds (the first pass welding parameters: 600 W, 80 mm/s, f.p.p. of 0 mm;
the second pass welding parameters: 200 W, 75 mm/s, f.p.p. of +2 mm) with different
shielding gases.

7.4 Discussion

The main problems in laser welding of Zn-coated steel to Al alloy are the formation of
spatter, porosity and intermetallic brittle phases in the weld. Since the boiling point of
Zn (907 ±C) being significant lower than the boiling point of Al (2520 ±C) and the
melting point of Fe (1538 ±C) [24], Zn vaporisation can occur and induce an unstable
melt pool and spatter, as shown in Figure 7-6(a) and Figure 7-6(b). Because the unstable
melt pool is difficult to avoid, using a second pass welding with a lower power
defocused laser beam to modify the top surface of the weld has been shown to be an
effective method in improving the weld appearance (Figure 7-6(c) and Figure 7-6(d))
without singificantly changing mechanical properties (Figure 7-8 and Figure 7-9).

As described in chapter 6, the mechanical properties of a weld can vary depending on

the presence of intermetallic brittle phases in the weld. For instance, the weld could
become harder and more brittle when the phase of FeAl3, FeAl2 or Fe2Al5 is present in
the weld. Conversely, the formation of Fe-rich intermetallic phases, such as FeAl and
Fe3Al, contributes to the weld’s ductility [134, 220, 228]. Due to the rapid heating and
cooling rates of laser welding, material reactions, mixing and interdiffusion and the
occurrence of intermetallic phases, especially FeAl3 and Fe2Al5 can be readily induced

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 Chapter 7 – Continuous wave fibre laser welding of Zn-coated steel to Al alloy

in the weld [229]. Therefore, a higher hardness can be obtained deep in the FZ where
intermetallic phases are easily produced (Figure 7-8, Figure 7-9, Figure 7-12, and Table
7-3).

The type of shielding gas using is another key factor affecting the weld quality,
particularly in welding Zn-coated steel with Al alloys. In laser keyhole welding, the
shielding gas has more opportunities to interact with the keyhole especially when the
melt pool is relatively unstable [230]. For this reason, physical properties of the
shielding gas, such as the ionisation potential, density and thermal conductivity, have to
be considered in the laser keyhole welding process. Amongst these, the thermal
conductivity and density of shielding gas are more important than its ionisation
potential in high energy density laser welding [103]. When laser keyhole welding is
carried out with a higher thermal conductivity shielding gas, the heating and cooling
rates can increase resulting in decreased the weld penetration depth (Figure 7-3). Also,
reducing the molten phase time in laser dissimilar materials welding has been found to
limit the formation of intermetallic phases in the weld due to the restricted heat flow and
the diffusion activity [130]. Accordingly, the extent of mixing materials can be
restricted, and welds with low hardness variations and a better shear force can be
obtained using the higher thermal conductivity N2 shielding gas (Figure 7-9 and Figure
7-13) in the process.

The density of shielding gas can also influence the formation of porosity and the
efficiency of melt pool protection [103]. Shielding gas with a high density has a higher
possibility of being trapped in the keyhole and in the weld after solidification. Hence,
less porosity can be expected in fibre laser double pass welding of Zn-coated steel on Al
alloy with N2 gas than with Ar gas (Table 7-1, Table 7-2 and Figure 7-5).

7.5 Conclusions

Laser lap welding of Zn-coated steel to Al alloy was carried out using a single mode
fibre laser. Results show that double pass welding could effectively produce welds with
better appearance than single pass welding. Furthermore, the thermal conductivity and
density of shielding gas play an important role in determining the weld quality. Welds
with less hardness variation and greater shear force were obtained when N2 gas was

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used as the shielding gas. The welds produced by double pass welding with N2 gas
exceed the industrial requirements in terms of shear forces for “stitching” applications.

It is interesting to compare the results here with those from a less comprehensive
investigation into the fibre laser welding of un-coated steels. This is described in detail
in Appendix 3. In essence, these investigations show that thermal conductivity of Al
alloy and Zn vaporisation occurring in the porcess play crucil roles in determining the
weld quality. With a result, fibre laser welding of un-coated steels is easier than that of
Zn-coated steel to Al alloy [231].

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Chapter 8
Corrosion performances in fibre laser
welding of dissimilar materials

8.1 Introduction

The reduction of vehicle weight and the improvement of corrosion resistance are
increasingly important for the automotive industry [162]. Zn-coated steel is commonly
used in car bodies due to its good corrosion resistance (in a wide range of chemical
environments), weldability, formability and paintability [158, 232]. For similar reasons,
Al alloy is another common material used for automotive applications. In a corrosive
environment, the Zn coating is dissolved before the steel because Zn has a lower
electrochemical potential than iron [16]. A Zn-coated steel corrosion process can be
divided into four stages, as illustrated in Figure 8-1 [233]. Firstly, the Zn coating is
corroded but the remaining coating still protects the entire substrate from the corrosion
environments. Secondly, a partial Zn coating has been lost. However, Zn corrosion
products and the remaining Zn coating can protect the substrate from the corrosion
environments. Thirdly, the substrate is slightly corroded. Meanwhile, Zn corrosion
products can functionally restrict the corrosion rate. Finally, the substrate steel is
exposed to the corrosion environment and both the Zn coating and corrosion products
are absent. The corrosion rate of the substrate is equal to that of the un-coated steel in a
similar corrosion environment.

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Figure 8-1 Schematic corrosion processes of Zn-coated steel [234].

In Chapter 7, results show that the shielding gas type could significantly affect
mechanical properties of welds in the fibre laser lap welding of Zn-coated steel to Al
alloy process. Previous work mainly focused on the corrosion resistance of Zn-coated
steel particularly in automotive applications [157, 234-240]. Few papers have been
reported on the corrosion behaviour in the laser welded Zn-coated steel process [16,
241, 242] or the laser welding of dissimilar materials process [243, 244]. Hardly any
publication is cited on the corrosion performance in fibre laser welding of similar or
dissimilar materials.

In the laser welding process, the rapid heating and cooling rates induce micro-
segregation, microstructure transformation and loss of alloy elements in the weld. They
can affect both the mechanical properties and corrosion resistance of the weld [30, 245].
Previous work reported that intergranular corrosion was obtained from the laser welding
of Zn-coated steel process, as shown in Figure 8-2(a) [16]; micro-segregation,
unfavourable phases and solidification cracking caused pitting and galvanic corrosions
in the laser welding of stainless steel process [30]; the localised corrosion, see Figure
8-2(b), was observed along the boundary of the FZ in the laser welding of AZ31 Mg
alloy process resulting from the solute concentration in the weld [246]; the grain size of
the weld influenced its corrosion resistance in the electron beam welding of Ti-6Al-4V
[247]. According to these findings, the corrosion resistance of a weld could be improved
by decreasing the surface roughness, reducing the formation of defects [248] or post-
heating the welded materials (i.e. the laser surface melting process) [249-252].

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Figure 8-2 Corrosion in laser welding of:(a) Zn-coated steel [16]; (b) AZ31 Mg alloy
[246].

This section aims to characterise and evaluate the corrosion performance in fibre laser
lap welding of Zn-coated steel to Al alloy for automotive applications. The influences
of the shielding gas type and the number of welding passes were investigated using the
electrochemical polarisation test method. Evaluations also included surface appearance
observation, surface roughness and chemical composition analyses before and after the
corrosion test.

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8.2 Experimental work

8.2.1 Electrochemical polarisation technique

Today, accelerated corrosion tests are commonly used in the laboratory to investigate
the corrosion performance within a short duration. Results obtained from accelerated
corrosion tests can only be used for indicating the corrosion behaviour compared with
that in the natural corrosion environment [253]. Regarding accelerated corrosion tests,
the electrochemical polarisation method has been widely used to evaluate the corrosion
resistance especially for the localised corrosion. It has the advantages of being sensitive
to the corrosion rate, having a short experimental duration and well-established
theoretical understanding. In an electrochemical polarisation test, the relative corrosion
rate and corrosion tendency can be compared according to the value of critical current
density and corrosion potential, respectively [254]. Basic knowledge of corrosion
science is briefly introduced in Appendix 4.

In an electrochemical polarisation test, the reaction rate (anode or cathodic reaction) can
be obtained by controlling the anodic or cathodic potential with a potentiostat
measurement. An example of a hypothetical anodic polarization curve is illustrated in
Figure 8-3.

• Point A: The anode reaction rate is equal to the cathodic reaction rate. Its current
density is close to zero.
• From Point A to Point B (The active region): Oxidation is the main mechanism
occurring in this period
• Point B: The value of corrosion potential and current density are named “The
passivation potential” and “The critical current density”, respectively.
• From Point B to Point C: The current density decreases with the increasing
corrosion potential.
• Point C: Its current density is named “The passivation current density”.
• From Point C to Point D (The passive region): The potential increases with a
constant current density before reaching Point D.
• After Point D: The amount of increasing corrosion potential depends on the
occurrence of pitting corrosion, crevice corrosion, the transpassive dissolution or
the oxide evolution [254].

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Figure 8-3 Illustration of an anodic polarisation plot [183].

8.2.2 Corrosion experiment

The specifications of weld materials, DX54 Zn-coated steel and EN-AW-5754 Al alloy,
and the experimental set-up are identical to those presented in Chapter 7. The shielding
gas type and the number of welding passes were chosen as variables in this work. The
experimental matrix is listed in Table 8-1. In double pass welding, the second pass was
carried out with a lower power defocused laser beam. In all cases, the second pass
welding parameters were 200 W, 75 mm/s with the focal point position (f.p.p.) of the
laser beam 2 mm above the top surface of Zn-coated steel. The spot diameter of the
defocused fibre laser beam at the surface was approximately 81 µm. Two different
gases, Ar and N2, were used as the shroud. Their physical properties are listed in Table
2-4. The corrosion tests were carried out using the electrochemical polarisation test
method with 3.5% NaCl solution at room temperature. The testing equipment,
parameters and procedures are described in Chapter 4. After corrosion tests, the surface
appearance and chemical composition profiles of uncorroded and corroded welds were
compared using a scanning electron microscope (SEM) equipped with energy dispersive
spectrometer (EDS) and a Wyko NT1100 optical profiling system.

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Table 8-1 Experimental matrix and electrochemical parameters.

Shielding Welding parameters i Electrochemical
Gas parameters j
P1 S1 F1 P2 S2 F2 I E
2
(W) (mm/s) (mm) (W) (mm/s) (mm) (A /cm ) (mV)
600 75 0 37.3 -631.3
600 100 0 39.8 -664.4
650 100 0 39.3 -585.3
Ar
600 75 0 200 75 +2 39.9 -518.6
600 100 0 200 75 +2 43.5 -583.3
650 100 0 200 75 +2 35.5 -589.8
600 75 0 24.7 -688.3
600 100 0 30.1 -684.6
650 100 0 38.9 -626.1
N2
600 75 0 200 75 +2 30.6 -631.4
600 100 0 200 75 +2 33.8 -645.0
650 100 0 200 75 +2 29.4 -608.3

8.2.3 Results

Compared with the uncorroded Zn-coated steel (Figure 8-4(a)), clear corrosion and
corrosion productions could be observed on the appearance of parent Zn-coated steel
after testing, as shown in Figure 8-4(b). The corrosion potential and critical current
density of the parent Zn-coated steel were -648.9 mV and 41.1 A/cm2, respectively.
Detailed electrochemical parameters of all tested samples are listed in Table 8-1.

i
P1, S1 and F1 are laser power, welding speed and the focal point position of the laser
beam in the first pass welding, respectively. P2, S2 and F2 are laser power, welding
speed and the focal position of the laser beam in the second pass welding.
j
I and E are the critical current density and its corrosion potential, respectively.

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Figure 8-4 Top surface appearance of the un-welded Zn-coated steel: (a) uncorroded
area; (b) corroded area.

8.2.3.1 Effect of the shielding gas type

Top surface appearances of corroded single pass welds produced with Ar and N2 gases
are shown in Figure 8-5(a) and Figure 8-5(b), respectively. Relatively severe corrosion
was observed when N2 gas was used. At each specific combination of single pass
welding parameters (600 W at 75 mm/s, 600 W at 100 mm/s or 650 W at 100 mm/s),
higher critical current densities and less negative corrosion potentials were obtained
from welds with Ar gas than those with N2 gas shrouding, as shown in Table 8-1.

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Figure 8-5 Top surface appearances between corroded single and double pass
welds(Single pass welding parameters were 650 W, 100 mm/s, f.p.p. of 0 mm. Double
pass: the second pass welding parameters: 200 W, 75 mm/s, f.p.p. of +2 mm.): (a) the
corroded single pass weld with Ar gas; (b) the corroded single pass weld with N2 gas;
(c) the corroded double pass weld with Ar gas; (d) the corroded double pass weld with
N2 gas.

Polarisation curves of single and double pass welds with different shielding gases are
shown in Figure 8-6. In Figure 8-6(a) of single pass welds, the weld produced with a
higher current density and lower negative corrosion potential was obtained when Ar
was used as the shielding gas compared with the use of N2 gas. In Figure 8-6(b), similar
results were obtained in double pass welds. The double pass weld with Ar shielding gas
showed better corrosion resistance than with N2 gas.

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 Chapter 8 – Corrosion performances in fibre laser welding of dissimilar materials

(a)

(b)
Figure 8-6 Polarisation curves with different shielding gases(Single pass welding
parameters were 600 W, 100 mm/s and f.p.p. of 0 mm. Double pass welding: the second
pass welding parameters: 200 W, 75 mm/s, f.p.p. of +2 mm).: (a) single pass welding;
(b) double pass welding.

For single pass welding at 600 W, 100 m/s and f.p.p. of 0 mm, the chemical
composition profiles from the top surface of uncorroded welds with Ar and N2 gases are
shown in Figure 8-7(a) and Figure 8-8(a), respectively. In both cases, the FZ was
composed of Fe and Zn with little Al present. Zn then sharply decreased at the boundary
of the FZ and became the main element on the un-affected surface. A weld with a
slightly wider affected zone was obtained when N2 gas was used. After the corrosion
test, Fe became the major element on the corroded weld surface as shown in Figure

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8-7(b). Figure 8-7(c) and Figure 8-8(b) are chemical composition profiles of double
pass welds with Ar and N2 gases, respectively. Results show that a smaller affected
zone with a higher average concentration of Zn was obtained with Ar gas in Figure
8-7(c) than the use of N2 gas in Figure 8-8(b).

(a)

Figure 8-7 Chemical composition profiles on different pass welds with Ar gas(Single
pass welding parameters were 600 W, 100 mm/s and f.p.p. of 0 mm. Double pass
welding:the second pass welding parameters: 200 W, 75 mm/s, f.p.p. of +2 mm).: (a)
the uncorroded single pass weld; (b) the corroded single pass weld; (c) the uncorroded
double pass weld.

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Chapter 8 – Corrosion performances in fibre laser welding of dissimilar materials

(b)

(c)

Figure 8-7 (Continued).

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 Chapter 8 – Corrosion performances in fibre laser welding of dissimilar materials

(a)

(b)

Figure 8-8 Chemical composition profiles on different pass welds with N2 gas (Single
pass welding parameters were 600 W, 100 mm/s and f.p.p. of 0 mm. Double pass
welding: the second pass welding parameters: 200 W, 75 mm/s, f.p.p. of +2 mm).: (a)
the uncorroded single pass weld; (b) the uncorroded double pass weld.

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8.2.3.2 Effect of number of welding pass

When Ar gas was used, more clear corrosion was found on the corroded single pass
weld (Figure 8-5(a)) than on the corroded double pass weld (Figure 8-5(c)). The
polarisation curves of single and double pass welds are shown in Figure 8-9. It can be
seen that the less negative corrosion potential was obtained from the double pass weld
than from the single pass weld. The chemical composition profiles of single and double
uncorroded welds are shown in Figure 8-7(a) and Figure 8-7(c), respectively. In the
double pass welding, the average concentration of Zn on the FZ was higher and a wider
affected zone was observed. In Table 8-1, a higher critical current density and a less
negative corrosion potential (-518.6 mV) were obtained in the double pass weld with
the first pass welding parameters of 600 W, 75 mm/s and f.p.p. of 0 mm than with the
single pass weld produced at 600 W, 75 mm/s and f.p.p. of 0 mm (-631.3 mV). The
surface roughness of corroded single and double pass welds are shown in Figure 8-10(a)
and Figure 8-10(b), respectively. In both single pass and double pass welding, a higher
surface roughness was obtained on the welded surface than on the surface of the un-
welded Zn-coated steel.

Figure 8-9 Polarisation curves between single and double pass welding with Ar gas.
(The first pass welding parameters: 600 W, 100 mm/s, f.p.p. of 0 mm; the second pass
welding parameters: 200 W, 75 mm/s, f.p.p. of +2 mm).

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(a)

(b)

Figure 8-10 The surface roughness of corroded welds obtained from different number of
welding passes with Ar shielding gas (The first pass welding parameters: 600 W, 75
mm/s, f.p.p. of 0 mm; the second pass welding parameters: 200 W, 75 mm/s, f.p.p. of
+2 mm.): (a) single pass weld; (b) double pass weld.

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When N2 gas was used, less corrosion was found on the corroded double pass weld
(Figure 8-5(d)) than on the corroded single pass weld in Figure 8-5(b). The polarisation
curves of corroded single (600 W, 100 mm/s and f.p.p. of 0 mm) and double (the first
pass welding parameters: 600 W, 100 mm/s, f.p.p. of 0 mm; the second pass welding
parameters: 200 W, 75 mm/s, f.p.p. of +2 mm) pass welds are shown in Figure 8-11.
Circles І and П highlight the peak of the maximum (or critical) and minimum current
density, respectively. In Circle І, a higher current density was obtained from the parent
Zn-coated steel while the single pass weld gave a smaller value. The double pass weld
had a less negative corrosion potential and a higher critical current density when the
first pass welding parameters were 600 W and 75 mm/s or 600 W and 100 mm/s,
compared with a single pass weld of the same welding parameters. The chemical
composition profiles of single and double uncorroded welds are shown in Figure 8-8(a)
and Figure 8-8(b), respectively. A wider HAZ with a higher average concentration of
Zn was obtained in double pass weld, as shown in Figure 8-8(b).

Figure 8-11 The polarisation curves of single and double pass welds with N2 gas.
(Single pass welding parameters were 600 W, 100 mm/s and f.p.p. of 0 mm. Double
pass welding: the second welding pass parameters: 200 W, 75 mm/s, f.p.p. of +2 mm.
Circles І and П highlight the peak of maximum and minimum current density,
respectively.)

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 Chapter 8 – Corrosion performances in fibre laser welding of dissimilar materials

8.2.4 Discussion

In laser welding, the corrosion resistance of a weld is related to its morphology,

chemical composition and electrochemical properties. Since Zn has the lower
electrochemical properties, it is dissolved before iron [16]. According to Figure 8-11,
the Zn coating starts to dissolve at the peak of maximum (or critical) current density as
highlighted at Circle І. The Zn coating was dissolved completely until the current
density reached the peak of minimum current density, as highlighted at Circle П. After
passing the peak at Circle П, the current density sharply increased as to the steel begins
to dissolve [16]. Finally, iron became the main element on the corroded weld (Figure
8-7 (b)).

The higher the surface roughness of a weld, the lower is its corrosion resistance (Figure
8-10). In fibre laser lap welding of Zn-coated steel to Al alloy, the high Zn vaporisation
readily produced resulting in an unstable melt pool. The surface finish could be
improved by applying the double pass welding process and therefore a lower corrosion
rate was obtained with Ar or N2 gas shrouding (Figure 8-5, Figure 8-9 and Figure 8-11).

Since the peak temperature of the melt pool in laser keyhole welding can be higher than
the boiling point of the weld materials, properly protecting the melt pool from oxidation
is important. The shielding gas type and the design of shrouding system can influence
the degree of oxidation. Previous work has shown that Ar gas could more effectively
protect the melt pool than the use of N2 gas because of its higher density and the
neutrally metallurgical effect [103]. Conversely, N2 gas is active and easily reacts with
iron, creating nitrides that can affect the corrosion resistance [255, 256]. Similar results
were found in this work. For both single and double pass welds, the lower corrosion rate
was obtained when Ar gas was used (Table 8-1 and Figure 8-5).

In addition to the melt pool oxidation, variations of chemical composition and losses of
alloying elements are other factors that determine the corrosion resistance of welds. The
FZ was composed of Zn, Fe and less Al, as shown in Figure 8-7(a), Figure 8-7(c) and
Figure 8-8. The residual Zn on the FZ was believed to be induced from the
concentration, spatter and the formation of Fe-Zn intermetallic phases during the laser
welding process [16, 257]. At the fusion zone boundaries, the concentration of Zn
sharply decreased as shown in Figure 8-7(a) and Figure 8-8(a). It was because that the

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Zn vaporisation reacted with the melting iron and therefore formed Fe-Zn intermetallic
phases during laser welding. However, the temperature in the HAZ was not high enough
to form the Fe-Zn intermetallic phases. During the second pass welding, more Zn from
the surface of unaffected zone was melted and mixed within the melt pool. Thus, more
Zn was observed on the top surface of double pass welds (Figure 8-7(c) and Figure
8-8(b)). On another point of view, the Fe-Zn intermetallic phases (i.e. FeZn13, FeZn7
and Fe3Zn10) and corrosion products (i.e. Zn(OH)2, ZnCl2·6Zn(OH)2 and
ZnCl2·4Zn(OH)2 [258-260]) on the weld could increase the Zn concentration. Both of
them could help to decrease the corrosion rate [240, 242, 259] and protect the weld
[240, 259]. For these reasons, in some cases, the corrosion resistance of welds were
slightly higher than the parent Zn-coated steel, particularly in double pass welding with
Ar gas (Figure 8-9).

8.3 Conclusions

In laser welding, the corrosion resistance of welds is related to variations of chemical

composition and microstructure in the weld and are controlled by the heating and
cooling rates in the weld. The proper selection of processing parameters is important for
improving the corrosion resistance of welds. In fibre laser welding of Zn-coated steel to
Al alloy, the shielding gas type and the number of welding passes could significantly
influence the corrosion performance of welds as well as their mechanical properties.
Results showed that the corrosion resistance of welds was improved by using the double
pass welding process. The Ar gas had more potential to produce a weld with less
corrosion.

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 Chapter 9 – General discussion on the basic characteristics of fibre laser welding of dissimilar materials

Chapter 9
General discussion on the basic
characteristics of fibre laser welding of
dissimilar materials

9.1 Introduction

Although laser welding has shown as one of the suitable methods for joining dissimilar
materials, some challenges still remain to be solved. Such problems include significant
variations of mechanical properties, the formation of intermetallic brittle phases and
thermal distortions. Amongst these, the presence of intermetallic brittle phases plays the
most important role in determining the weld quality. It can be said that any factor which
can induce the intermetallic brittle phases must be studied and analysed. This chapter
presents the general features of fibre laser welding of dissimilar materials. It aims to
understand the role of influential factors controlling the weld quality. Furthermore, the
potential of fibre lasers for welding different materials are outlined.

The interacting effects observed in fibre laser welding of dissimilar materials are shown
in Figure 9-1. In laser welding of dissimilar materials, the joining materials basically
undergo melting, mixing and solidification processes. As soon as materials melt, they
mix in the weld which induces the occurrence of intermetallic phases before
solidification. The extent of intermetallic phases depends on the degree of material
mixing which is related to the temperature gradient in the melt pool. In theory, the fewer
intermetallic phases forming in a weld, the better mechanical properties the weld has.

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 Chapter 9 – General discussion on the basic characteristics of fibre laser welding of dissimilar materials

Figure 9-1 The interacting effects in fibre laser welding of dissimilar materials.

9.2 The main factors in fibre laser welding of dissimilar

materials

Slightly different mechanisms were observed between laser butt and lap welding of
dissimilar materials. In fibre laser butt welding of dissimilar materials, the amount of
heat input, the cooling rate and, especially, the laser beam spot position on the weld are
important in determining the weld quality. This is demonstrated by the results obtained
from fibre laser butt welding of Ti-6Al-4V titanium alloy to Incoenal 718 nickel alloy in
Chapter 5. In laser lap welding of dissimilar materials, intermetallic phases normally
occurred at the bottom of the welded materials. The weld penetration depth and the
extent of the intermetallic phases are mainly related not only to the amount of heat input
but to the thermal conductivity of the weld materials. The lower melting and boiling
points of the weld materials produced deeper welds as observed in results presented in
Chapter 6 and in Appendix 3. From the viewpoint of metallurgy, the shielding gas type
is another factor influencing the weld quality in fibre laser lap welding of dissimilar
materials as found in Chapter 7.

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 Chapter 9 – General discussion on the basic characteristics of fibre laser welding of dissimilar materials

Generally speaking, as shown in Figure 9-2, three factors have to be considered when
using a fibre laser to weld dissimilar materials. Firstly, the amount of heat input, which
is actually absorbed by the weld materials, has to be precisely controlled. In laser
welding of dissimilar materials, the laser beam absorption is a function of the
reflectivity of the welding materials, laser power, welding speed and so on. Furthermore,
the thermal conductivity and the melting and boiling points of the weld materials play
an important role in determining the weld quality. A clear understanding of the
relationship between the weld quality and these three aspects needs to be achieved by
means of a series of experimental investigations. It is an effective approchwhen
launching a new project on fibre laser welding of dissimilar materials.

Figure 9-2 Influenced factors in fibre laser welding of dissimilar materials.

It is interesting to compare these results using a fibre laser with published results which
are based on a Nd:YAG laser. On the face of its laser beam wavelength, power and
frequency, it would be expected that they would achieve fibre laser welding
performances.

Some results from fibre laser and CO2 laser welding of Ti alloy to Ni alloy are shown in
Figure 9-3(a) and Figure 9-3(b), respectively. A smaller and parallel-shaped weld was
obtained in fibre laser welding compared with CO2 laser welding. The degree of

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 Chapter 9 – General discussion on the basic characteristics of fibre laser welding of dissimilar materials

material mixing phenomenon in the weld was relatively severe in CO2 laser welding as
compared with fibre laser welding. The results shows that the affected zone and the
formation of intemetallic phases can be minimised in fibre laser welding of dissimilar
materials owing to its better beam quality and higher energy density. It is believed that
fibre lasers have the potential to overcome common problems occurring in laser welding
of dissimilar materials. They include the joining of materials with different properties,
the reduction of intermetallic phases and thermal distortions (particularly for
applications requiring laser keyhole mode welding). Compared with other conventional
laser systems, there is a high possibility to produce welds with better quality and less
variation of mechanical properties by using fibre lasers with a proper control of the
processing parameters.

Figure 9-3 Laser welding of Ti alloy to Ni alloy using: (a) CW single mode fibre laser;
(b) CW CO2 laser (White arrows highlighting the fusion interfaces.) [95].

In conclusion, the features of fibre laser welding of dissimilar materials can be briefly
summarised.

• The weld geometry is mainly determined by the thermal conductivity and

melting point of the welding materials.
• A higher hardness in the weld can be attributed to the formation of intermetallic
phases and the refinement of the microstructure.
• The thermal conductivity and the melting and boiling points of the welding
materials can affect the rate of heat dissipation in the weld which modifies the
weld quality.

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 Chapter 10 – Conclusions and recommendation for future work

Chapter 10
Conclusions and recommendation for
future work

10.1 Conclusions

Conclusions are included at the end of each chapter in this thesis. This thesis has
covered a wide range of aspects of fibre laser welding of dissimilar materials, some
important results and conclusions are collected and presented below.

Fibre laser butt welding of Ti-6Al-4V titanium alloy to Inconel 718 nickel alloy

1. A better quality weld with less hardness variations and less formation of cracks
could be generated by offsetting the laser beam approximately 35 µm from the
interface to the Inconel 718 side and using a combination of a higher laser power
and a higher welding speed.

2. Hardness variations in welds were related to the occurrence of Ti-Ni

intermetallic brittle phases (Ti2Ni and TiNi3).

3. The formation of cracks in the weld is closely related to the amount of heat input,
the melt pool area and the cooling rate in the weld.

4. There was no clear evidence to show that the formation of crack and porosity
was strongly controlled by the offset position of the laser beam alone.

Fibre laser lap welding of Zn-coated steel to Al alloy

1. Narrow and parallel-shaped keyhole welds were produced with an IPG 1 kW

single mode fibre laser (Chapters 6 and 7).

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 Chapter 10 – Conclusions and recommendation for future work

2. The hardness in the FZ of the Al alloy side sharply increased resulting from the
occurrence of intermetallic brittle phases. All welds broke with a brittle fracture
under the shear force test (Chapters 6 and 7).

3. The pulse frequency was an influencing factor, especially at the lower heat input
welding condition, but laser power and welding speed were important in
determining the weld geometry when the heat input was higher than a critical
value (Chapter 6).

4. Double pass welding could effectively produce welds with better appearance and
corrosion resistance than the use of single pass welding (Chapters 7 and 8).

5. The shielding gas type plays an important role in determining the weld quality.
Welds with less hardness variation and greater shear force were obtained when
N2 gas was used as the shielding gas. A weld with the better corrosion resistance
was produced in Ar gas shrouding (Chapters 7 and 8).

10.2 Recommendation for future work

Melt pool behaviour observation

Results presented in this thesis show that the shielding gas type could influence
mechanical properties and corrosion resistance of welds. Therefore, monitoring the melt
pool behaviour in the fibre laser keyhole welding process with different shielding gases
by using a high speed camera is recommended. Elucidating the relationship between the
shielding gas type, residual stresses and the corrosion resistance of welds in fibre laser
welding of dissimilar materials is also needed.

Process window optimisation

The formation of porosity and cracks in the weld is a common issue in the laser welding
of dissimilar material process. A deep and systematic investigation of the main
parameters producing porosity and cracks in fibre laser welding of dissimilar materials,
through a non-destructive testing (NDT) method, is needed. An optimal process window
for producing defect-free welds would be expected to be obtained.

185
 Chapter 10 – Conclusions and recommendation for future work

Process modelling

Another issue in the laser welding of dissimilar materials process is the occurrence of
intermetallic brittle phases. Results from this thesis have shown that it plays an
important role in determining the weld quality. It was believed that the amount of
intermetallic phases is closely related to the melting velocity, the convective flow and
the degree of micro-segregation in the weld. For these reasons, a study on the
convective flow developing in fibre laser welding of dissimilar materials by means of
computational fluid dynamics (CFD) simulation software is desirable.

Laser power delivery mode

Currently, light alloys, such as aluminium alloys and magnesium alloys, are
increasingly used in industry. Challenges in laser welding of these materials are the
lower energy efficiency resulting from their high thermal conductivity and high
reflectivity. Precisely controlling the energy density could be an effective method to
overcome this issue. Therefore, fibre lasers with good beam quality and high energy
density have potential in this field. Investigating the effect of laser power delivery mode
(pulsed wave and continuous wave) in fibre laser welding of high thermal conductivity
materials is important for the joining of light alloys.

186
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 Appendix 1

Appendix 1

Table A1-1 Laser welding of Al alloys to other materials.

Materials Sheet Joining Laser Application Year Organisation /
thickness type type Country
(mm)
Be:0.1 PW Nd: Nuclear BIAS, Germany
Al / Be lap 2005
Al:1 YAG Aeronautical [174]
University of
Science and
Al / Be - butt Nd:YAG Aeronautical 2006 Technology
Beijing, China
[261]
Al / Ni- or
1.2 PW BLZ, Germany
Ag-based spot/lap Electronics 2004
Nd:YAG [262]
filler / Cu
PW BLZ, Germany
Al / Cu 1.2 spot/lap Electronics 2006
Nd:YAG [263]
Nagaoka
CW Nd- University of
Al / Mg 1 lap Automotive 2005
YAG Technology,
Japan [191]
Dalian University of
Al / Ce PW laser
1.7 lap Transportation 2006 Technology ,
filler / Mg –TIG
China [264]
Dalian University of
Mg: 1.2
Al / Mg lap Laser Transportation 2007 Technology ,
Al: 1.7
China [125]
Dalian University of
Al: 1.7
Al / Mg lap Nd:YAG - 2009 Technology ,
Mg: 1.2
China [19]
Al / Zn-
based filler
CW LTm, France
/ Zn Steel: 0.77 lap Automotive 2007
Nd:YAG [218]
coated-
steel
Al / steel
Al / 1
BIAS, Germany
Coated 1.3 lap Nd:YAG Automotive 1997
[15]
steel 1.5
Al / Ti
Al: 1 Automotive Osaka University,
Al / PET lap Diode 2008
PET: 2 Electronics Japan [265]
SPCC: 1 butt CW Osaka University,
Al / SPCC Automotive 2005
Al: 1 or 2 lap Nd:YAG Japan [227]
Tokyo Institute of
Steel: 1
Al / steel lap YAG Automotive 2005 Technology,
Al: 1.2
Japan [224]
Al alloy
Steel: 0.77 CW LTm, France
/Zn filler / Tee Automotive 2006
Al:1.2 Nd:YAG [266]
steel

203
 Appendix 1

Table A1-1 (Continued.)

Materials Sheet Joining Laser Application Year Organisation /
thickness type type Country
(mm)

Tokyo Institute of
Steel: 1
Al / steel lap YAG Automotive 2006 Technology,
Al: 1.2
Japan [222]
Steel: 1.2 CW GERAILP,
Al / steel lap Automotive 2007
Al: 1.3 Nd:YAG France [127]
Nagaoka
Steel: 1.2 University of
Al / steel lap Laser Automotive 2007
Al: 1.6 Technology ,
Japan [130]
Nagoya
Al / steel 0.5 lap CW fibre Automotive 2009 University ,
Japan [267]
PW Iranian National
Al / steel 2 lap Automotive 2010
Nd:YAG Centre, Iran [21]
Al / Zn- University of
coated - lap Nd:YAG Automotive 2005 Bayreuth ,
steel Germany [211]
Al / filler /
Osaka University
Zn-coated 1 lap Nd:YAG Automotive 2009
, Japan [217]
steel
Al/ Zn-
Al: 1.15 Nd:YAG- BIAS, Germany
coated butt - 2009
steel:1 MAG [268]
steel
Al / 316
National Laser
Stainless CW
0.9 lap Electronics 2007 Centre , South-
steel Nd:YAG
Africa [269]
Al / Cu
University of
Science and
Ti: 1.5
Al / Ti lap CO2 Transportation 2008 Technology
Al: 1
Wuhan, China
[188]
Harbin Institute of
Al /filler/ Ti 3 butt CO2 - 2010 Technology ,
China [270]

204
 Appendix 1

Table A1-2 Laser welding of Cu to other materials.

Materials Sheet Joining Laser Application Year Organisation /
thickness type type Country
(mm)

Beijing University
Cu: 1.5
Cu / Brass lap Nd:YAG Electronics 2009 of Technology ,
Brass: 1.1
China [271]
Fraunhofer
Spot
PW Institute for Laser
Cu / steel 250 µm lap Automotive 2003
Nd:YAG Technology,
Tee
Germany [55]
Indian Institute of
Cu / Fe 60 butt CW CO2 Electronics 2005 Science , India
[272]
Shanghai
Power- Jiaotong
Cu / Fe 7 butt CO2 2009
generation University, China
[273]
Kovar /
tool steel
Welding research
Cu / tool PW Chemical
1 butt 2004 centre, Korea
steel Nd:YAG Electrical
[190]
Cu / Al
alloy
PW BIAS, Germany
Cu / steel 0.1 lap Electronics 2005
Nd:YAG [174]
CuFe2P / Laser Zentrum
Cu PW
35, 70 µm lap Electronics 2002 Hannover e.V.,
CuFe2P / Nd:YAG
Cu / Pl Germany[274]
Indian Institute of
Cu / Ni 50 butt CW CO2 - 2005 Science , India
[275]

Table A1-3 Laser welding of Mg alloy to other materials.

Materials Sheet Joining Laser Application Year Organisation /
thickness type type Country
(mm)
Mg / 304 Dalian University
Mg: 1.7 YAG-
Stainless lap Automotive 2008 of Technology,
steel:1.2 GTA
steel China [9]
Dalian University
Mg / Ni Mg: 1.7 Laser –
lap - 2010 of Technology,
layer/steel Steel: 1.2 TIG
China [276]
AZ31/AM6
Hunan University,
0/ZK60 2 butt CO2 - 2008
China [277]
Mg alloy
Warsaw
AZ91 / AM
University of
50 Mg 4.5 butt CO2 Automotive 2009
Technology,
alloy
Poland [278]

205
 Appendix 1

Table A1-4 Laser welding of Ni alloy to other materials.

Materials Sheet Joining Laser Application Year Organisation /
thickness type type Country
(mm)
Ti-Ni
based
shape Diameter:
Spot PW Osaka University
memory 0.15~ 0.35 Medical 2003
lap Nd:YAG , Japan [279]
alloy(NT- mm
E4) wire /
3O4 SS
Superalloy
Chinese
K418
CW Academy of
alloy / 3.5 butt Engine 2007
Nd:YAG Sciences, China
steel
[280]
42CrMo
Indian Institute of
Ni / Ti 20 butt CW CO2 Aerospace 2006 Science , India
[95]
The University of
Fibre
Ni / Ti 2 butt Aerospace 2008 Manchester, UK
laser
[281]

Table A1-5 Laser welding of stainless steels to other materials.

Materials Sheet Joining Laser Application Year Organisation /
thickness type type Country
(mm)
304
stainless Cidade
PW
steel / 420 0.8 butt - 2007 Universita´ria ,
Nd :YAG
stainless Brazil [79]
steel
Computational
Au /
Materials
Stainless - butt Nd:YAG Jewelry 2008
Laboratory ,
steel
Switzerland [282]
316
stainless Dublin City
steel / 2 butt CO2 - 2008 University ,
carbon Ireland [283]
steel

206
 Appendix 1

Table A1-6 Laser welding of Ti alloy to other materials.

Materials Sheet Joining Laser Application Year Organisation /
thickness type type Country
(mm)
Single
Welding research
Ti: 0.1 mode
Ti / Steel lap Transportation 2008 centre, Korea
Steel: 0.7 CW fibre
[189]
laser
Ti64 / 316 National Laser
CW
Stainless 0.9 lap Medical 2007 Centre , South-
Nd:YAG
steel Africa [269]
Korean Atomic
Energy Research
Ti / Ta 0.2 tube Nd:YAG Nuclear 2008
Institute, Korea
[284]

Table A1-7 Laser welding of hard metals to other materials.

Materials Sheet Joining Laser Applications Year Organisation /
thickness type type Country
(mm)
Hard
CO2
metals Technical
CW
K10 / steel University of
2.5 butt Nd:YAG Cutting tools 2003
Hard Lisbon , Portugal
PW
metal K40 [285]
Nd:YAG
/ steel

207
 Appendix 2

Appendix 2

Figure A2-1 Drawing of the top part of custom-mode jig.

208
 Appendix 2

Figure A2-2 Drawing of the bottom part of custom-mode jig.

209
 Appendix 3

Appendix 3
Fibre laser welding of un-coated steels

A3.1 Introduction

In laser welding of dissimilar materials, weld quality is strongly affected by the

differences of chemical and physical properties between weld materials. This section
summarises the results of fibre laser lap welding of un-coated steels and Zn-coated steel
to Al alloy. Experiments were undertaken using an IPG 1 kW continuous wave fibre
laser. The weld geometry, microstructure and hardness distributions were evaluated.
The purpose of this comparative work is to point out the most influential material
properties and the generic characteristics of the fibre laser welding of dissimilar
materials.

A3.2 Experiments

In fibre laser welding of Zn-coated steel to Al alloy, the specifications of DX54 Zn-
coated steel and EN-AW-5754 Al alloy are identical to those presented in Chapter 7.
The Zn coatings on both sides of the DX54 Zn-coated steel were removed using #400
SiC papers for the fibre laser welding of un-coated steels process. The fixture and the
experimental set-up were similar with those used in Chapter 7. The laser power and the
number of welding passes were chosen as variables, while the welding speed was kept
constant at 100 mm/s. The welding process was shrouded with Ar gas. The
experimental matrix is tabulated in Table A3-1. After welding, the samples were
sectioned, mounted, polished and etched for further examinations. Nital acid (3% HNO3
ethanol solution) was used to etch un-coated steels and Zn-coated steels while Al alloy
was etched by Keller’s reagent. The weld geometry was observed and measured using
an optical microscope, as illustrated in Figure 7-1. Furthermore, the microstructure and
microsegregation were examined using a scanning electron microscope. Hardness

210
 Appendix 3

distribution was measured at three different penetration depth levels (i.e. 250 µm, 750
µm and 1100 µm) of the welds using a Vickers hardness testing machine, as shown in
Figure 7-1and Chapter 4.

Table A3-1 Parameters used in fibre laser lap welding of un-coated steels and Zn-coated
steel to Al alloy.
Welding parametersk
The first pass The second pass
P1 S1 F1 P2 S2 F2
(W) (mm/s) (mm) (W) (mm/s) (mm)
400 100 0
450 100 0
500 100 0
550 100 0
600 100 0
650 100 0
600 100 0 200 75 +2
650 100 0 200 75 +2

A3.2.1 Geometry and microstructure

The relationship between laser power and the weld penetration depth at the bottom of
the welded materials is shown in Figure A3-1. Similar trends of increasing the weld
penetration depth with increasing laser power were found in both cases of fibre laser
welding Zn-coated steel to Al alloy and un-coated steels. Deeper welds were produced
in fibre laser lap welding of Zn-coated steel to Al alloy than those obtained in fibre laser
welding of un-coated steels at each specific combination of welding parameters. Cross-
sectional welds are compared in Figure A3-2. At the top of the welded materials,
parallel-shaped FZs were obtained in both cases. In fibre laser welding of un-coated

k
P1, S1 and F1 are power, welding speed and the focal point position of the laser beam in
the first pass welding, respectively; P2, S2 and F2 are power, speed, and the focal point
position of laser beam in the second pass welding, respectively.

211
 Appendix 3

steels, the weld width linearly decreased as the weld depth increased from the top to the
bottom of welding materials, as shown in Figure A3-2(b). A finer grain size in the FZ
was obtained in fibre laser welding of Zn-coated steels, see Figure A3-2(b), as
compared with that in fibre laser welding of un-coated steel to Al alloy. At the bottom
of the welded materials, a wider FZ with a clear mixture of materials was found in fibre
laser welding of Zn-coated steel to Al alloy, as shown in Figure A3-2(a).

Figure A3-1 The relationship between laser power and the weld penetration depth in
single pass welding with a constant welding speed of 100 mm/s.

212
 Appendix 3

Figure A3-2 Cross sections of single pass welds at 450 W and 100 mm/s in single pass
welding of: (a) Zn-coated steel to Al alloy; (b) un-coated steels.

A3.2.2 Hardness

Hardness distribution of welds obtained in fibre laser double pass welding of Zn-coated
steel to Al alloy and those from un-coated steels are shown in Figure A3-3(a) and
Figure A3-3(b), respectively. At the penetration depths of 250 µm and 750 µm, slightly
higher hardness near the FZ were obtained from fibre laser welding of un-coated steels
than those obtained from fibre laser welding of Zn-coated steel to Al alloy. Meanwhile,
at the 1100 µm penetration level, the hardness near the FZ claerly increased in fibre

213
 Appendix 3

laser welding of Zn-coated steel to Al alloy. Similar hardness distributions between the
three penetration depth levels were found in fibre laser welding of un-coated steels, as
show in Figure A3-3(b).

(a) (b)
Figure A3-3 Hardness distributions in fibre laser double pass welding of: (a) Zn-coated
steel to Al alloy; (b) un-coated steels. (The first pass welding parameters: 600 W, 100
mm/s, f.p.p. of 0 mm; the second pass welding parameters: 200 W, 75 mm/s, f.p.p. of
+2 mm).

214
 Appendix 3

A3.2.3 Micro-segregation

A cross-section of the weld produced by fibre laser single pass welding of Zn-coated
steel to Al alloy is shown in Figure A3-4. A material mixing phenomenon was observed
in the Al alloy side according to the colour rule of SEM backscattered electron image.
In fibre laser double pass welding of un-coated steels, the weld with refined
microstructure and without any occurrence of material mixing phenomena was obtained,
as shown in Figure A3-5.

Figure A3-4 Backscattered electron image of fibre laser single pass welding of Zn-
coated steel to Al alloy with Ar gas shrouding. (650 W, 100 mm/s and f.p.p. of 0 mm.)

Figure A3-5 Backscattered electron image of fibre laser double pass welding of un-
coated steels with Ar gas shrouding. (The first pass welding parameters: 600 W, 100
mm/s, f.p.p. of 0 mm; the second pass welding parameters: 200 W, 75 mm/s, f.p.p. of
+2 mm).

215
 Appendix 3

A3.3 Discussion

Differences between fibre laser welding of Zn-coated steel to Al alloy and un-coated
steels have been discovered. The weld geometry is strongly determined by the thermal
conductivity of the welding materials. The FZ could not be consistent between welding
materials in fibre laser welding of dissimilar materials. The fusion area was sharply
increased in the Al alloy side from fibre laser welding of Zn-coated steel to Al alloy
(Figure A3-2). Higher thermal conductivity weld materials produced deeper and wider
welds. The thermal conductivity of pure aluminium is approximately three times higher
than pure iron, see Table 6-3. Therefore, deeper and wider fusion areas at the bottom
material side were obtained in fibre laser welding of Zn-coated steel to Al alloy than
those in fibre laser welding of un-coated steels (Figure A3-1 and Figure A3-2).

Another difference is the formation of intermetallic phases. Most of the intermetallic

phases in welds of Zn-coated steel to Al alloy are brittle and hard, which increases the
hardness of the weld but reduces its strength. The hardness was abruptly increased near
the FZ in fibre laser welding of Zn-coated steel to Al alloy (Figure A3-3) because of the
presence of the material mixing phenomenon (Figure A3-4). Excluding the intermetallic
phases, hardness is also related to the grain size. A smaller grain size produced at a high
cooling rate can increase hardness in the FZ. In comparison with fibre laser welding of
Zn-coated steels to Al alloy, higher hardness were obtained near the FZ of the top
welded material side in fibre laser welding of un-coated steels (Figure A3-3(b)). This is
because heat was quickly dissipated by Al alloy or by Zn vaporisation in fibre laser
welding of Zn-coated steel to Al alloy. Whereas, heat was transferred slowly and easily
retained in the fusion area during the fibre laser welding of un-coated steels. The
localised heat may cause a higher cooling rate which may induce a finer microstructure
in the weld leading to higher hardness.

216
 Appendix 4

Appendix 4
Introduction of corrosion mechanisms

Metal corrosion is an electrochemical process. Four essential elements (i.e. anode,

cathode, electrolyte and electrical circuit) are necessary in Figure A4-1. The main
reaction at the anode and the cathode are oxidation reaction and reduction reaction,
respectively. In a corrosion environment, the anode and electrolyte present as the
corroded material and the corrosion media, respectively. The corrosion reaction
equations on the anode and cathode are presented in Eq.(A4-1) and Eq.(A4-2),
respectively.

The anodic reaction:

M → Mn- + ne- (A4-1)

The cathodic reaction:

2H- + 2e- → H2 (A4-2)

where M, n and e are the involved metal, valence of the corroding metal species and
electrons. The number of electrons lost at the anode is equal to the number of electrons
gained at the cathode. The corrosion rate is higher for the more anodic metals. In the
other words, the more noble material has the best corrosion resistance, while the
reactive metal with more negative electrode potential can be corroded more easily.
Table A4-1 presents electrochemical properties of pure metals.

Figure A4-1 Illustration of an electrochemical process [254].

217
 Appendix 4

Table A4-1 Electrochemical properties of pure metals [208].

Corrosion can be classified as the uniform corrosion, pitting corrosion, galvanic

corrosion, intergranular corrosion, crevice corrosion, dealloying, erosion corrosion,
environmentally induced cracking (i.e. stress corrosion, corrosion fatigue and hydrogen-
induced cracking). [286]. Figure A4-2 illustrates corrosion types commonly observed in
a weld. According to the British standard of BS EN ISO 8044:2000, their definitions
and mechanism are briefly introduced in the following sections [287]:

218
 Appendix 4

Figure A4-2 Schematic of corrosion types: (a) uniform corrosion; (b) pitting corrosion;
(c) galvanic corrosion; (d) intergranular corrosion; (e) crevice corrosion; (d) stress
corrosion [286].

Uniform corrosion: The corrosion rate and the metallurgical effect are almost identical
at every location of the metallic surface, as shown in Figure A4-2(a).

Pitting corrosion: It is formed at a localised area of the passive metal surface as shown
in Figure A4-2(b). The degree of pitting corrosion relates to its chemical
composition and microstructure [288].

Galvanic corrosion: This occurs when the joining materials have different corrosion
potentials (Figure A4-2(c)). The less noble material has a lower corrosion rate. In
the weld, it normally is observed at the adjacent area between the affected and un-
affected zones [30, 286, 289].

Intergranular corrosion: This forms at the grain boundaries (Figure A4-2(d)) and
relates to the formation of anodic precipitated phases at the grain boundaries.
Anodic precipitated phases can induce intergranular corrosion due to a reduction of

219
 Appendix 4

the corrosion resistance occurring at the grain boundaries compared with that of the
remainder grains. An improper heat treatment and welding process readily cause
intergranular corrosion.

Crevice corrosion: It shows at the mating surface between one metal to another metal,
as shown in Figure A4-2(e).

Stress corrosion: Three factors (i.e. surface tensile stress, susceptible alloy and specific
environment) are simultaneously required to cause stress corrosion, see Figure
A4-2(f). Stress corrosion can be minimised by increasing the compression stress
on the exposed surface of materials [245].

220