Computational Fluid Dynamic Modelling of Laser Additive Manufacturing Process and Effect of Gravity

COMPUTATIONAL FLUID DYNAMIC MODELLING OF LASER ADDITIVE MANUFACTURING PROCESS AND EFFECT OF GRAVITY

A thesis submitted to

The University of Manchester

For the degree of

Doctor of Philosophy (PhD)

In the Faculty of Science and Engineering

2017

Heng Gu

School of Mechanical, Aerospace and Civil Engineering

List of contents

List of contents ...... 1

List of figures ...... 6

List of tables ...... 12

Nomenclatures ...... 13

Acronyms ...... 16

Abstract ...... 17

Declaration ...... 18

Acknowledgement ...... 20

Chapter 1 Introduction ...... 21

1.1 Research rationale ...... 21

1.2 Aim and objectives of the project ...... 23

1.3 Thesis structure ...... 24

Chapter 2 Literature review of laser additive manufacturing ...... 27

2.1 Introduction ...... 27

2.2 Laser basics ...... 27

2.2.1 Spontaneous emission and stimulated emission ...... 28

2.2.2 Population Inversion ...... 30

2.2.3 Laser Cavity ...... 30

2.3 Laser beam characteristics ...... 30

2.3.1 Monochromaticity ...... 31

2.3.2 Coherence ...... 31

2.3.3 Directionality...... 32

2.3.4 Brightness ...... 32

2.4 Types of lasers...... 32

2.4.1 Diode Lasers...... 33

2.4.2 Fibre lasers ...... 34

List of contents

2.4.3 Disk lasers ...... 35

2.5 Laser additive manufacturing ...... 36

2.5.1 Pre-placed wire and powder ...... 37

2.5.2 Powder injection...... 38

2.5.3 Wire feeding ...... 40

2.5.4 Combined wire and powder ...... 41

2.6 Process Parameters ...... 42

2.6.1 Laser power and specific energy ...... 42

2.6.2 Process velocity ...... 43

2.6.3 Gas-powder flow rate ...... 44

2.6.4 Overlap ...... 45

2.6.5 Deposition pattern ...... 47

2.7 Mechanical and thermal behaviour ...... 49

2.7.1 Stress ...... 49

2.7.2 Porosity ...... 51

2.7.3 Cracking and distortion ...... 53

2.7.4 Microstructure ...... 54

2.7.5 Surface unevenness ...... 56

2.7.6 Re-melting ...... 60

2.8 Melt pool behaviour ...... 62

2.8.1 Marangoni flow ...... 62

2.8.2 Gravitational force ...... 64

2.8.3 Melting and solidification ...... 68

2.8.4 Overhang ...... 72

2.9 Modelling ...... 75

2.9.1 Thermal-mechanical modelling ...... 75

2.9.2 Multi-physical modelling ...... 80

2.10 Summary and discussion ...... 90 Chapter 3 Formulation of a CFD model for laser additive manufacturing process ...... 92

List of contents

3.1 Introduction ...... 92

3.2 Previous work ...... 92

3.3 Governing equations ...... 96

3.4 VOF model ...... 98

3.5 Surface tension ...... 99

3.6 Melting and solidification ...... 101

3.7 Buoyancy force ...... 101

3.8 Numerical implementations ...... 102

3.9 User-defined energy, mass and momentum sources ...... 103

3.9.1 Energy source ...... 103

3.9.2 Mass source ...... 107

3.9.3 Momentum source ...... 108

3.10 Boundary conditions ...... 109

3.10.1 Internal boundary condition ...... 110

3.10.2 External boundary condition ...... 111

3.11 Temperature gradient and solidification rate ...... 111

3.12 Summary ...... 113 Chapter 4 Study and prediction of laser metal deposition with various structures ...... 114

4.1 Introduction ...... 114

4.2 Material ...... 114

4.3 Experimental equipment ...... 115

4.3.1 Laser applied ...... 115

4.3.2 Powder delivery system ...... 117

4.4 Modelling strategy ...... 118

4.5 Experimental results and discussion ...... 122

4.5.1 Influence of laser power ...... 122

4.5.2 Influence of powder feed rate ...... 123

4.5.3 Influence of process velocity ...... 124

4.6 Modelling of single track ...... 126

List of contents

4.7 Comparison of influence factors ...... 128

4.8 Multiple layer deposition ...... 134

4.9 Deposition pattern ...... 136

4.10 Re-melting ...... 140

4.11 Modelling of deposition with multiple adjacent passes ...... 141

4.12 Modelling of re-melting process ...... 143

4.13 Modelling of deposition with complex shapes...... 145

4.14 Modelling of geometries with overhang structures ...... 146

4.15 Conclusion ...... 148 Chapter 5 CFD modelling of laser metal deposition process and effect of gravity ...... 150

5.1 Introduction ...... 150

5.2 Modelling strategy ...... 151

5.3 Experimental procedures ...... 155

5.4 Results and discussion ...... 158

5.4.1 Deposition morphology ...... 158

5.4.2 Melt pool development ...... 162

5.4.3 Temperature development ...... 169

5.5 Surface unevenness at both ends ...... 171

5.5.1 Bulge formation ...... 171

5.5.2 Effect of process conditions on bulge formation ...... 173

5.5.3 Effect of system instability on bulge formation ...... 175

5.6 Effect of surface tension coefficient ...... 180

5.6.1 Effect of surface tension coefficient on non-flat deposition ...... 180

5.6.2 Effect of surface tension coefficient on bulge formation ...... 184

5.7 Effect of gravity levels and zero gravity ...... 188

5.8 Conclusion ...... 193 Chapter 6 CFD modelling of narrow gap welding based additive manufacturing process ...... 195

6.1 Introduction ...... 195

List of contents

6.2 Previous work ...... 196

6.2.1 Effect of gravity on welding process ...... 196

6.2.2 Narrow gap welding ...... 198

6.2.3 Dissimilar welding ...... 200

6.3 Numerical modelling of narrow gap welding ...... 202

6.3.1 Modelling strategy ...... 202

6.3.2 Lack of fusion ...... 206

6.3.3 Effect of gravity direction on narrow gap welding process ...... 209

6.3.4 Effect of gravitational level on narrow gap welding process ...... 212

6.3.5 Effect of surface tension coefficient on narrow gap welding ...... 215

6.3.6 Narrow gap welding of thick sections ...... 218

6.4 Numerical modelling of narrow gap dissimilar welding...... 221

6.4.1 Modelling strategy ...... 221

6.4.2 Modelling of bead asymmetry ...... 222

6.4.3 Modelling of re-melting process ...... 224

6.5 Conclusion ...... 224

Chapter 7 Conclusion and Future work recommendations ...... 227

7.1 Conclusions ...... 227

7.2 Future work recommendations ...... 232

Reference ...... 233

Word count: 69,124

List of figures

Figure 1.1 Diagram for thesis structure...... 26 Figure 2.1 Basic construction of a laser system [8] ...... 28 Figure 2.2 Three major processes of electromagnetic radiation: (a) absorption, (b) spontaneous emission, (c) stimulated emission ...... 29 Figure 2.3 Schematic of a diode laser system [12] ...... 34 Figure 2.4 Principle of a cladding pumped fibre laser [13] ...... 35 Figure 2.5 Concept of a thin disk laser [14] ...... 36 Figure 2.6 Pre-placed powder direct metal deposition [16] ...... 38 Figure 2.7 Lateral powder injection [12] ...... 39 Figure 2.8 Coaxial powder injection [20] ...... 40 Figure 2.9 Laser additive manufacturing using wire, (a) front feeding, (b) rear feeding, wire placed at (c) leading edge, (d) centre and (e) trailing edge of melt pool [23, 24]... 40 Figure 2.10 Combined wire and powder deposition [27] ...... 42 Figure 2.11 Powder stream structure (a) pre-waist stage, (b) waist stage, (c) post-waist stage [43] ...... 45 Figure 2.12 Average of grain size vs. Overlap [44] ...... 46 Figure 2.13 A schematic view of overlap between adjacent bead [29] ...... 47 Figure 2.14 Deposition patterns (a) raster, (b) offsetout, (c) offsetin, (d) fracture [49] . 48 Figure 2.15 Stress in x, y, z-direction (a) along the centre line in deposition direction (b) at the middle height of wall in width (Y) direction [57] ...... 50 Figure 2.16 Lack of fusion and residual gas porosity [61] ...... 52 Figure 2.17 Microstructure change from bottom to the top layer of deposition [75] ..... 54 Figure 2.18 Optical photos of longitudinal section for different deposition patterns [78] ...... 55 Figure 2.19 Bulge effect at the beginning of deposition [80] ...... 56 Figure 2.20 Bead geometry at the beginning of the deposition with different surface tension gradient (a) Negative, (b) positive, (c) Mixed [85] ...... 57 Figure 2.21 Melt pool on the edge with different deposition sequence [48] ...... 58 Figure 2.22 Simulation of humping effect in high-speed gas tungsten arc welding [91]59 Figure 2.23 Cross-section of bead when applying different process velocity (a) 200 mm/min, (b) 400 mm/min, (c) 700 mm/min, (d) 1000 mm/min [118] ...... 64 Figure 2.24 Buoyancy force and convection flow driven by buoyancy force [120]...... 66 Figure 2.25 (a) Melt pool shape formed by Marangoni stress only, (b) Melt pool shape formed by gravity force only, (c) Melt shape formed by the combination of those two forces together [122] ...... 67 Figure 2.26 Impact of temperature gradient and solidification rate on microstructure formation [129] ...... 69 Figure 2.27 Growth rate and temperature gradient on solidification boundary with different melt pool shape [120] ...... 70

List of figures

Figure 2.28 Illustration of overhang, L is layer thickness, St is layer top surface and Sb is layer bottom surface [134] ...... 72 Figure 2.29 Two different methods to produce overhang structures[136]...... 73 Figure 2.30 Contact angle of a water droplet adhering on a glass window [142] ...... 74 Figure 2.31 Stress components of a single track laser deposition (a) x-direction, (b) y- direction, (c) z-direction, (d) von Mises equivalent stress [151] ...... 76 Figure 2.32 Phase fraction of martensite during laser metal deposition [160] ...... 78 Figure 2.33 Temperature distribution and bead geometry [172] ...... 83 Figure 2.34 Bead deposited with various process velocities...... 84 Figure 2.35 Temperature distribution of third layer after 6s [182] ...... 86 Figure 2.36 Coupled Level set and VOF method [190] ...... 87 Figure 2.37 Melt pool surface with and without free deformation [197]...... 90 Figure 3.1 Conservation in a discrete element [213] ...... 96 Figure 3.2 Mass balance in a small volume element dV [214] ...... 96 Figure 3.3 Molecular transport of the momentum in the x-direction in an arbitrary ...... 97 Figure 3.4 (a) Interface between two phases. (b) VOF function of the two liquids [217] ...... 98 Figure 3.5 Normal stress and tangential stress on the free interface ...... 99 Figure 3.6 Comparison between cells on the curved interface and flat surface ...... 105 Figure 3.7 Schematic of the laser metal deposition process ...... 109 Figure 3.8 Boundary conditions of the model ...... 110 Figure 4.1 Stainless steel 316L powder ...... 115 Figure 4.2 Laserline LDL 160-1500 Diode laser ...... 116 Figure 4.3 Measured power output and nominal power ...... 116 Figure 4.4 Variation between spot size and the distance from nozzle tip ...... 117 Figure 4.5 FST powder feeder ...... 117 Figure 4.6 Coaxial nozzle used for the experiments [98] ...... 118 Figure 4.7 Schematic of laser metal deposition with coaxial powder feeding ...... 118 Figure 4.8 Mesh for the calculation domain ...... 119 Figure 4.9 Boundary conditions and initial phase patch of the calculation domain ..... 120 Figure 4.10 Illustration of boundary conditions on cross-section plan ...... 120 Figure 4.11 Single layer deposition using different laser power ...... 122 Figure 4.12 Single layer deposition using different powder feed rate ...... 124 Figure 4.13 Cross-section of single deposition track (A) 8 mm/s, (B) 10 mm/s, (C) 12 mm/s, (D) 14 mm/s, (E) 16 mm/s, (F) 18 mm/s, (G) 20 mm/s ...... 125 Figure 4.14 Simulated deposition and temperature distribution using 1 kW laser power, 1.7 mm beam diameter, 0.28 g/s powder feed rate and 8 mm/s process velocity ...... 126 Figure 4.15 Development of melt pool and velocity field 0.588 s, 1.2 s, 1.896 s, 2.4 s ...... 127 Figure 4.16 Effect of process velocity on deposition cross section (comparison A-G with Figure 4.13) ...... 127 Figure 4.17 Effect of process velocity on deposition top view (comparison A-G) ...... 128 Figure 4.18 Effect of process velocity on deposition side view (comparison A-G) ..... 128 Figure 4.19 Effect of laser power on layer height ...... 129

List of figures

Figure 4.20 Effect of powder feed rate on layer height ...... 129 Figure 4.21 Effect of process velocity on layer height ...... 129 Figure 4.22 Effect of laser power on layer width...... 130 Figure 4.23 Effect of process velocity on layer width ...... 130 Figure 4.24 Effect of process velocity on layer width ...... 130 Figure 4.25 Effect of laser power on dilution ...... 131 Figure 4.26 Effect of powder feed rate on dilution ...... 131 Figure 4.27 Effect of process velocity dilution ...... 131 Figure 4.28 Effect of factors on deposition height ...... 133 Figure 4.29 Effect of factors on deposition width ...... 133 Figure 4.30 Cross-sections of 5 samples ...... 135 Figure 4.31 Four methods to print L (A) Two segments, (B) Raster, (C) offset in, (D) offset out...... 138 Figure 4.32 Four deposited letter L by using different deposition methods ...... 138 Figure 4.33 Two methods to print C, (A) raster (B) offset out ...... 139 Figure 4.34 Path design for letter P and deposited sample ...... 139 Figure 4.35 (A) Deposition of re-melted sample L (B) Comparison of the side wall between re-melted sample and Figure 4.31 (D) ...... 140 Figure 4.36 Laser deposition process of five adjacent passes ...... 141 Figure 4.37 Melt pool development history during depositing five adjacent passes .... 142 Figure 4.38 Side view of the deposition formation ...... 143 Figure 4.39 Re-melting process with excessive heat input ...... 143 Figure 4.40 Surface finish comparison before and after re-melting process ...... 144 Figure 4.41 Deposition of letters ‘LPRC’ and temperature distribution ...... 145 Figure 4.42 Deposition of letters ‘LPRC’ with cooling between features ...... 145 Figure 4.43 Deposition of second and third overhanging layers ...... 146 Figure 4.44 Multilayer deposition with overhanging structures ...... 147 Figure 4.45 Multilayer deposition with complex overhanging structures ...... 147 Figure 5.1 Mesh for the calculation domain ...... 151 Figure 5.2 Boundary conditions and initial phase patch of the calculation domain ..... 152 Figure 5.3 Illustration of boundary conditions on cross-section plan ...... 153 Figure 5.4(a) Cavitar laser illumination system (b) High-speed camera in horizontal position ...... 156 Figure 5.5 Schematic diagrams of wire laser deposition process (a) flat (b) vertical ... 156 Figure 5.6 Experimental set-up equipped with high-speed camera system ...... 157 Figure 5.7 2-layer deposition result and cross-section (a) top view, (b) experimental cross section, (c) cross-section of modelling result ...... 157 Figure 5.8 Side views of depositions in the vertical up position with a laser spot size of (a) Ø3 mm [case 3], (b) Ø2 mm [case8] ...... 158 Figure 5.9 Deposition in flat, vertical down, vertical up, overhead and horizontal positions with Ø3 mm, P=3 kW, v=0.01 mm/s, f=0.02 m/s (2 layers) ...... 159 Figure 5.10 Deposition in flat, vertical down, vertical up, overhead and horizontal positions with Ø2 mm, P=1.3 kW, v=0.01 mm/s, f=0.02 m/s (2 layers) ...... 160

List of figures

Figure 5.11 Deposition in vertical up position (3 layers), v=0.01 mm/s, f=0.009 m/s, Case 11: Ø2 mm, P=1.3 kW Case 12: Ø3 mm, P=3 kW ...... 161 Figure 5.12 Deposition in vertical up position, Case13: Ø2 mm, P=3 kW, v=0.01 mm/s, f=0.02 m/s (2 layers) Case 14: Ø3 mm, P=4 kW, v=0.01 mm/s, f=0.045 m/s (1 layer) ...... 161 Figure 5.13 Temperature and melt pool-velocity field history for case 8, (a&f:0.36 s, b&g:1.44 s, c&h:1.80 s, d&i:1.908 s, e&j:2.196 s)...... 162 Figure 5.14 melt pool evolution for cases with big spot size ...... 164 Figure 5.15 melt pool evolution for cases with small spot size ...... 165 Figure 5.16 Comparison of melt pool evolution for cases with big and small spot size ...... 165 Figure 5.17 Five allocated points on the substrate surface ...... 169 Figure 5.18 Temperature history of 4 selected points for Case 6 ...... 169 Figure 5.19 Temperature history of 4 selected points for Case 1 ...... 170 Figure 5.20 Temperature history of middle point P1 for different cases ...... 171 Figure 5.21 Bulge at the beginning of case 4 (second layer, 0.18 s) ...... 172 Figure 5.22 Bulge morphology at 0.42 s and projection of the bulge morphology at 2.2s ...... 174 Figure 5.23 Acceleration and deceleration of the robot arm at starting and ending of the deposition process ...... 175 Figure 5.24 Deposition with (a) and without (b) consideration of system instability... 175 Figure 5.25 Melt pool and velocity field with (a) and without (b) consideration of system instability in the middle of second layer deposition (1.2 s and 1.1 s) ...... 176 Figure 5.26 Comparison of melt pool development between with and without consideration of system instability ...... 177 Figure 5.27 Top view (a) and side view (b) of 7-layer deposition ...... 178 Figure 5.28 Image from the high-speed camera (a) Starting part (b) Ending Part...... 179 Figure 5.29 Modelling of deposition without synchronized heat and mass feed ...... 179 Figure 5.30 Deposition in vertical up position with negative, positive and mixed ∂γ/∂T, Ø2 mm, P=1.3 kW, v=0.015 mm/s, f=0.02 m/s ...... 182 Figure 5.31 Comparison of melt pool development for different surface tension coefficient ...... 182 Figure 5.32 Comparison of temperature history for different surface tension coefficient ...... 183 Figure 5.33 Comparison of melt pool development for different surface tension coefficient with lower process velocities ...... 183 Figure 5.34 Surface tension on bulge formation (a) negative (b) positive (c) mixed ... 184 Figure 5.35 Temperature history of middle point P1 with different surface tension coefficient ...... 185 Figure 5.36 Melt pool and velocity field for different surface tension coefficient (a) negative 4.3 s, (b) negative 5.4 s, (c) positive 4.3 s, (d) positive 5.4 s, (e) mixed 4.3 s, (f) mixed 5.4 s ...... 186 Figure 5.37 Melt pool development of cases with different surface tension coefficient ...... 187

List of figures

Figure 5.38 Comparison of the first deposited layer under different gravity levels (Ø=2 mm, P=1.3 kW, v=0.01 mm/s, f=0.02 m/s) ...... 188 Figure 5.39 Comparison of the second deposited layer under different gravity levels (Ø=2 mm, P=1.3 kW, v=0.01 mm/s, f=0.02 m/s) ...... 189 Figure 5.40 Melt pool evolution for cases with different gravity values ...... 189 Figure 5.41Comparison of deposition cross-section with different gravity values ...... 190 Figure 5.42 Contact angle and aspect ratio with different gravity values ...... 191 Figure 5.43 Comparison of deposition with different process parameters ...... 192 Figure 5.44 Comparison of deposition cross-section and melt pool with different process parameters (a,g): P=1.3 kW, v=0.01 mm/s, f=0.02 m/s, (b,h): increased laser power P=1.6 kW (c,i): increased velocity v=0.015 mm/s, (d,j): further increased velocity v=0.02 mm/s, (e,k): reduced wire federate f=0.01 m/s, (f,l): reduced wire federate f=0.01 m/s and power P=1 kW ...... 192 Figure 6.1 Schematic of narrow gap welding process ...... 202 Figure 6.2 Mesh for the calculation domain ...... 203 Figure 6.3 Boundary conditions and initial phase patch of the calculation domain ..... 203 Figure 6.4 Illustration of boundary conditions on cross-section plan ...... 204 Figure 6.5 First layer narrow gap welding, low laser power (a,b), higher laser power(c,d), after re-melting (e)(f) ...... 206 Figure 6.6 Middle plane cross-section and melt pool velocity field: low laser power (a)(b), higher laser power (c) (d), after re-melting (e)(f) ...... 207 Figure 6.7 second and third layer for different directions : flat (a)(b), horizontal (c)(d), overhead (e)(f), vertical up (g)(h), vertical down (i)(j) ...... 209 Figure 6.8 Second layer melt pool and third layer temperature: flat (a)(b), horizontal (c)(d), overhead (e)(f), vertical up (g)(h), vertical down (i)(j) ...... 210 Figure 6.9 Melt pool development history of first two layers...... 211 Figure 6.10 Melt pool development history of three layers ...... 212 Figure 6.11 Second and third pass for vertical up with: 0 g (a)(b), 1 g (c)(d), first and second pass for vertical up 5 g (e)(f) ...... 213 Figure 6.12 Melt pool-velocity field and temperature distribution at end of second pass: 0 g (a)(b), 1 g (c)(d), 5 g (e)(f) ...... 214 Figure 6.13 Melt pool development history of first two passes ...... 214 Figure 6.14 Melt pool development history of three passes ...... 215 Figure 6.15 Second and third pass of the weld with different surface tension coefficient: negative ∂γ/∂T (a)(b), positive ∂γ/∂T (c)(d), mixed ∂γ/∂T (e)(f) ...... 216 Figure 6.16 Second pass melt pool-velocity field and third pass temperature distribution: negative ∂γ/∂T (a)(b), positive ∂γ/∂T (c)(d), mixed ∂γ/∂T (e)(f) ...... 216 Figure 6.17 Melt pool development history of first two passes ...... 217 Figure 6.18 Melt pool development history of three passes ...... 218 Figure 6.19 Weld bead development ...... 218 Figure 6.20 Cross-section development (middle y plane) ...... 219 Figure 6.21 Cross-section development (middle x plane) ...... 219 Figure 6.22 Melt pool and temperature distribution of first four passes ...... 220 Figure 6.23 Melt pool development history ...... 220

List of figures

Figure 6.24 Schematic sketch of the calculation domain ...... 221 Figure 6.25 (a) 3 kW flat position (b) 3 kW horizontal position (c) 4 kW flat position ...... 223 Figure 6.26 (a,b)3 kW flat position (c,d) 3 kW horizontal position (e,f) 4 kW flat position ...... 223 Figure 6.27 (a,b,c) before re-melting, (d,e,f) after re-melting ...... 224 Figure 6.28 (a) before re-melting, (b)after re-melting ...... 224

List of tables

Table 4.1 Chemical composition of 316L powder [227] ...... 115 Table 4.2 Relationship between laser power and deposition results ...... 123 Table 4.3 Relationship between powder feed rate and deposition result ...... 124 Table 4.4 Relationship between process velocity and deposition result ...... 125 Table 4.5 Experimental results of design expert ...... 132 Table 4.6 Analysis of variance table (ANOVA) ...... 133 Table 4.7 Deposited dimension by applying different process parameters ...... 135 Table 5.1 Thermos-physical properties of 316L [114] ...... 154 Table 5.2 Process parameters in different cases for the simulation and experiments ... 155 Table 5.3 Cases with different surface tension-temperature coefficient ...... 180 Table 6.1 Chemical composition of SA 508 Gr.3 Cl.2 steel and Alloy 52 (wt.%) [263] ...... 221

Nomenclatures

ℎ∗ Planck constant (J s)

휈 Electromagnetic radiation frequency (Hz)

E1 Lower energy state

E2 Higher energy state

퐿 Cavity length (m)

휆 Wavelength of the laser radiation (m)

휏 Coherence time (s)

o 휃푚 Diffraction limit angle ( ) a Beam waist radius (m)

-2 -1 퐵휆 Monochromatic brightness of laser (W m sterad )

푃 Laser power (W)

o 휃0 Divergence angle ( )

퐸 Specific energy (J kg-1)

D Spot diameter (m)

V Process velocity (m/s)

푄푚 Powder mass rate (kg/s)

H Overall layer heigh (m) h Single layer height (m) k Overlap ratio (%)

0 -1 γ푚 Surface tension of pure metal (N m )

푇0 Reference melting temperature (K)

푅 Gas constant (J K−1 mol−1)

Г Surface excess at saturation 푠

푘 Entropy segregation constant

Nomenclatures

훥퐻0 Enthalpy of segregation (J)

푎푖 Activity of species 𝑖 in solution (%)

G Temperature gradient (K m-1)

-1 Vs Solidification velocity (m s )

푇̇ Cooling rate (K s-1)

∆푇 Temperature difference between solidus and liquidus temperature (K)

2 -1 퐷퐿 Diffusion coefficient (m s )

F Liquid fraction (%)

ρ Material density (kg m3) t Time (s) u, v, w Component velocity in x, y and z direction (m s-1)

-3 -1 푆푚 Volumetric mass source addition rate (kg m s )

푝 Static pressure (Pa)

휏̿ Stress tensor (N m-2)

퐹⃗ External body forces (N)

퐻 Enthalpy (J)

푘 Thermal conductivity (W m−1 K−1)

-3 푆ℎ Volumetric heat source (W m )

푝2, 푝1 Pressures on two sides of the fluid (Pa)

훾 Surface tension coefficient (N m-1 K-1)

푅1, 푅2 Two radius defining the surface curvature (m)

κ Surface curvature (m-1) n Surface normal

-3 휌푎푣푔 Average density of the interfacial cell (kg m )

-3 휌푖, 휌푗 Density of phase i and j (kg m )

ℎ푟푒푓 Reference enthalpy (J)

Nomenclatures

푇푟푒푓 Reference temperature (K)

-1 퐶푝 Specific heat at constant pressure (J K )

퐿 Latent heat of the material (J kg-1)

퐴푚푢푠ℎ Mushy zone constant

훽 Thermal expansion coefficient (K-1)

푀푆 Self-adaptive mass addition rate (kg m-3 s-1)

-2 -1 ℎ푐 Heat convection coefficient (W m K )

σ Boltzmann constant (J K−1)

ε Radiation emissivity

T푒푛푣 Environment temperature (K)

-1 Gx Gy Gz Temperature gradient on x, y and z direction (K m )

C Laser concentration coefficient

푟푛 Radial distance from the heat source centre on the circular disk (m)

-2 푞0 Maximum heat flux at the centre of heat source model (W m )

-3 퐻푆푛 Heat source term on a certain cell (W m )

푇푡푖푚푒푠푡푒푝 Actual time of each calculation interval (s)

3 푉푐푒푙푙 Cell volume (m )

휂 Laser power efficiency (%)

-3 퐻푆푡표푡푎푙 Total heat source input (W m )

푀푆 Mass source term for each selected cell (kg m-3 s-1)

-3 -1 퐹푤푖푟푒 Mass addition rate (kg m s )

푟푤푖푟푒 Wire radius (m)

α Angle between wire feeder and sample surface (o)

-1 ΔH0 Standard heat of adsorption (J kg mol )

μ Dynamic viscosity (kg m-1 K-1)

푓훼 Sulphur concentration in substrate (ppm)

Acronyms

CW Continuous wave DC Direct current RF Radio frequency LAM Laser additive manufacturing CAD Computer-aided design SLS Selective laser sintering SLM Selective laser melting LOM Laminated object manufacture DMD Direct metal deposition LMD Laser metal deposition SDRS Single direction raster scanning CDRS Cross direction raster scanning FEM Finite element method FVM Finite volume method HVOF High-velocity oxygen fuel TIG Tungsten inert gas GTAW Gas tungsten arc welding GMAW Gas metal arc welding LENS Laser engineered net shaping PDAS Primary dendrite arm spacing SDAS Secondary dendrite arm spacing AISI American Iron and Steel Institute HAZ Heat affected zone VOF Volume of fluid ALE Arbitrary Lagrangian Eulerian CFD Computational fluid dynamics UDF User defined functions TEM Transverse electromagnetic modes CSF Surface force mode NGLW Narrow gap laser welding

Abstract

Name of university: The University of Manchester

Submitted by: Heng Gu

Degree title: Doctor of Philosophy

Thesis title: Computational fluid dynamic modelling of laser additive manufacturing process and effect of gravity

Date: 30th September 2017

Laser additive manufacturing (LAM) is based on selectively adding materials, layer by layer, to form a 3D part using one or multiple laser beams to fuse or solidify the materials. Although considerable amount of work can be found on investigating the LAM processes, little is known on the effects of gravity and dynamic fluid flow characteristics in different material growth orientations. Along with the advancement of laser manufacturing techniques, LAM is increasingly being used in a wide variety of environments including cylinder body, surface cladding on turbine blades, offshore drilling heads, and side walls of sleeve and mould where various deposition orientations are normally required. In addition, for space additive manufacturing, the operating environment will experience very low or zero gravity. Many efforts have been made by the previous research on developing numerical methods to model the LAM process. However, most of the previous modelling work has been focusing on development of the melt pool dynamics without taking into account of the free surface formation. Only a few investigations include the analysis of material addition into the dynamic flowing melt pool. No work has been found on developing a model which can simulate the deposition process with all complex features and taking gravity effects into consideration when performed in various material deposition orientations and zero gravity effects. In this research, a three-dimensional transient computational fluid dynamic model was built for the LAM process which took into account compound process factors including material addition, surface tension, melting and solidification, gravitational force, temperature dependent material properties, free surface formation and moving heat source. A better understanding of laser metal deposition process was achieved numerically and experimentally. The research covered the deposition of single layer, multiple adjacent passes and full three-dimensional geometries with overhanging features. Effects of gravity during the deposition process in various deposition orientations and for zero gravity and very low gravity were investigated and process parameters were optimised to minimise the effect. The research also extended to understanding of basic phenomena in a laser narrow gap welding process, based on layer by layer material addition, and how gravity would affect the melt pool formation inside the groove when the welding process was performed in different orientations. By analysing the melt pool development history and temperature distribution, the influence of surface tension coefficient was discussed during the process. With the help of the current model, various defects including deposition irregularity, bulge and slope at both ends of the deposition, lack of fusion, step effect, surface waviness and collapsing as a result of change of gravity were explained. Corresponding solutions to eliminate all these defects were presented. The work on zero-gravity laser additive manufacturing discovered several new phenomena not reported before, paving the way for future laser 3D printing in space.

Declaration

I hereby declare that no portion of the work referred to in this 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.

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 (the “Copyright”) and s/he 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, trademarks 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://documents.manchester.ac.uk/DocuInfo.aspx?DocID=2442 0), in any relevant Thesis restriction declarations deposited in the University Library, The University Library’s regulations (see http://www.library.manchester.ac.uk/about/regulations/) and in The University’s policy on Presentation of Theses.

Acknowledgement

My deepest gratitude goes first and foremost to my entire family for their selfless dedication throughout this whole journey. I am indebted to my parents, who give endless support and encouragement, shared all my happiness and sorrow throughout this study. I also owe a special debt of gratitude to my uncle and aunt who provide me with this precious opportunity to improve myself. I would like to thank my cousin who brings me the warmth of family in this foreign country.

I would like to give the sincere gratitude to my supervisor Professor Lin Li for his great support and inspiring advice but left sufficient freedom for me to discover this sea of knowledge. I am extremely gratefully to Dr Wei Guo, Dr David Whitehead, Daniel Wilson and Damian Crosby for their scientific advice and technical support.

I would like to thank all of my friends in the University of Manchester, especially to Tapio Väistö, Dr Isuamfon Edem, Dr Ahmad Wael Al Shaer, Dr Wei Guo in the student village and Dr Anastasia Vasileiou, Jacqui Grant from MTRL. I would like to thank all of these friends for leaving such a beautiful memory in this chapter of my life.

Finally, I would express my gratitude to my beloved wife, Lili, for her unconditional love and understanding, tremendous encouragement and patience, and all that I have learned from her. She is the most invaluable gem I have gained throughout this whole journey.

Chapter 1 Introduction

1.1 Research rationale

Additive manufacturing, based on material addition to form a 3D component, is a type of manufacturing technology integrating precision machinery, material science, computer science and numerical control. In comparison with other traditional manufacturing methods, the additive manufacturing process has its advantages of higher material usage rate, reduced total cost and the capability of producing workpiece with very complex geometry, in short lead time.

With the rapid development of laser technology, the laser additive manufacturing (LAM) process is becoming increasingly applied which can be attributed to the laser beam characteristics including high energy density, small heat affected area and high process efficiency. Laser metal deposition (LMD) process has its superiority of producing metallic products with larger dimensions and much higher production rate when compared with other laser additive manufacturing methods such as selective laser melting.

Although considerable amount of work can be found on investigating the laser metal deposition process, little is known on the effects of gravity during the LMD process. However, deposition in non-flat positions is increasingly required now, which includes surface processing on a cylindrical body, surface cladding on turbine blades, offshore drilling heads, and side-walls of sleeve and mould. Apart from these, gravity also has a substantial influence on producing overhanging structures since there is no supporting material for the samples during the deposition process. In addition, for additive manufacturing in space or other planets, zero or very low gravity is encountered. Performing such experiments in space would be very costly.

Process parameters should be optimised to achieve good deposition quality. However, experimental trials are normally time-consuming and costly. Also, equipment requirement is high for conducting such research under different process directions. Numerical modelling can be a both time-saving and cost-effective approach to obtaining a full understanding of how gravity will affect the deposition processes.

Chapter 1 Introduction

Despite the fact that many numerical models can be found on simulating the laser additive manufacturing process, few have focused on simulating the real-time depositing process with complex features. Melt pool development with or without free surface tracking, and simple geometries like single depositing line were typically chosen as the research target during the modelling work. Effects of gravity have always been largely ignored and most of the practical problems in real experimental processes during depositing complicated geometries are often neglected.

A considerable amount of research has been found on the narrow gap welding process, based on material addition, layer by layer, for its advantages including a smaller heat- affected area, lower total heat input, reduced product deformation and higher welding efficiency in comparison with other traditional welding methods. With the rapid development of laser technology, a laser beam can be applied as the heat source in laser narrow gap welding process which can reduce the gap width and heat input.

In certain industrial applications, structures with large dimensions would be difficult to be built with a single additive manufacturing process. Narrow gap welding can be applied to join additive manufactured parts.

During the narrow gap welding process, it is difficult to monitor the real-time melt pool formation inside the deep groove, even with the help of high-speed camera. Numerical simulation can be a useful tool to predict the weld bead formation in the groove under different process conditions. Most of the existing models concentrate on stress distribution by analysing the thermal behaviour of the process.

Up to now, no work has been found on developing a model which can simulate the laser metal deposition process with complex geometry and at the same time takes gravity effects into consideration when the process is performed in various operating directions and when the process is operated in zero gravity. No previous model has been built to investigate the melt pool formation during the narrow gap welding process taking account of the dynamic flowing condition inside the groove. Therefore, models for these two manufacturing processes need to be developed in order to get a better understanding of the included phenomena.

Chapter 1 Introduction

1.2 Aim and objectives of the project

Laser additive manufacturing processes use a laser beam as the heat source to melt the material which is supplied in the form of wire or powder. After the molten material cools down and solidifies, another layer of material will be processed on top of the solidified layer. Therefore, a fully dense three-dimensional product can be manufactured layer by layer with designed path. The aim of this project is to further understand the characteristics of laser additive manufacturing process, particularly the laser metal deposition process (LMD), involving various geometries with the consideration of the influence of gravity using computational fluid dynamic (CFD) modelling and experimental validation.

To achieve this aim, the objectives of this project include:

⚫ To build a CFD model for single-layer LMD process and validate the model with experimental data. ⚫ To develop a CFD model to simulate the LMD process with multiple adjacent passes and understand the melt pool development history as well as the development history of temperature distribution. ⚫ To develop a CFD model which is capable of simulating depositing full 3D geometries with overhanging features. ⚫ To investigate the effect of gravity and process parameters on melt pool development in non-flat deposition positions. ⚫ To understand various deposition surface irregularities including waviness, bulge at the beginning and collapsing at the end with the help of simulation. ⚫ To develop a CFD model to simulate a laser narrow gap welding process with the consideration of the gravity effect, surface tension gradient and strategies for eliminating some of the welding defects.

Chapter 1 Introduction

1.3 Thesis structure

This thesis covers various aspects of the laser additive manufacturing process with experimental and theoretical analysis. The thesis consists of seven chapters which can be described as follows and Figure 1.1.

Chapter 2 first gave some fundamental introductions on laser basics. Then a discussion of laser additive manufacturing was presented, after which the focus was placed on laser metal deposition processes. Process parameters which would influence the final deposition geometry were reviewed. Mechanical and thermal phenomena involved in the process were discussed. Since melt pool behaviour is the most significant factor in determining the thermal history and deposition formation, the mechanism of the melt pool development was reviewed. Various numerical methods for modelling the LMD process especially on melt pool formation were reviewed. Despite the fact that many studies have been conducted on understanding the melt pool mechanism, current literature review reveals that there has been little discussion on the gravity effects during the deposition.

Chapter 3 presented a full description of a three-dimensional numerical model which included free surface tracking, melting and solidification process, interfacial surface forces, various boundary conditions and the user-defined source terms. Self-adaptive mass and energy addition sources applied on the changing free surface were introduced. The gravitational force was added into the conservation equations and the effect of gravity could be fully uncovered in this model.

Chapter 4 presented the experimental details of the coaxial powder laser deposition process and the effect of process parameters on final deposition dimensions were investigated. By applying the optimised process parameters, depositions with complex features were successfully produced. Effect of deposition patterns and re-melting process on surface improvement was discussed. A three-dimensional model was built to simulate the single layer deposition process and the modelling result was compared with the experimental data. More models including re-melting and real 3D geometries were created to simulate the various phenomena during the deposition process which have not been successfully modelled previously. The current model is capable of simulating complex geometries with overhanging structures, which takes into account the effect of gravity, surface tension, mass addition and other thermal behaviour during the deposition.

Chapter 1 Introduction

The influence brought by the gravity effect could be seen from the modelling results with different depositing orientations.

Chapter 5 first presented the modelling strategy of a three-dimensional model, followed by the description of wire-based experimental setup equipped with high-speed camera system. The effects of gravity on deposition result in various positions were successfully demonstrated and the influence of different process parameters was discussed. Melt pool formation history and temperature development were explained to help understand how gravity would affect the material flowing conditions inside the melt pool. The model also successfully simulated the bulge formation at the deposition beginning and the collapsing at the end. Apart from the factor caused by thermal behaviour, system instability brought by the wire feeder and robot arm was also taken into consideration and added to the model. The gravity effects on unevenness at both ends were investigated. As an important factor in the melt pool, the influence of surface tension-temperature coefficient on melt pool formation in non-flat deposition orientation was discussed and how this coefficient could affect the unevenness at both ends of the deposition was also investigated. Gravity level was gradually decreased from 1g to zero g in order to investigate the effect of reduced gravity on deposition result for potential LMD process in space. Process parameters were optimised to minimise the surface irregularity under zero gravity condition.

Chapter 6 first gave a brief overview of current research status of laser narrow gap welding process. A three-dimensional model was then introduced for laser narrow gap welding process. Effect of gravity on weld bead formation in various welding positions was discussed. Lack of fusion inside the groove between the deposition and side walls could be observed from the model and process parameters were optimised to eliminate this phenomenon. Influence of gravitational force and surface tension coefficient on weld pool formation were discussed. Modelling of thick section multi-pass narrow gap welding process was presented and the phenomena included during the building up were discussed. Based on the laser narrow gap welding model, simulation of laser narrow gap dissimilar metal welding process was introduced by including one gas phase and three metallic phases. The influence of re-melting and gravity effect during dissimilar welding was also discussed.

Chapter 7 presented the overview of the main findings and conclusions from this research, followed by recommendations for the future work.

Chapter 1 Introduction

Figure 1.1 Diagram for thesis structure

Chapter 2 Literature review of laser additive manufacturing

2.1 Introduction

In this chapter, fundamentals of high-power laser were introduced first, which included the basic construction of a laser system, laser beam generation principle, laser beam characterises and some widely used laser types. Next, a brief description on laser additive manufacturing was presented and the focus was placed on laser metal deposition process which covered different methods of delivering material. Effect of process parameters on the deposition result was then discussed, after which mechanical and thermal behaviour included in the process were reviewed. Considering the importance of melt pool mechanism during the laser metal deposition process, literature reviews on the development of melt pool were presented, and the effect of gravity was emphasized especially in non-flat operating positions. Last, different methods of numerical modelling LMD process were investigated, where advantages and disadvantages of each method were discussed.

2.2 Laser basics

As an important symbol of technological development in the 20th century, laser which is the abbreviation for Light Amplification by Simulated Emission of Radiation is highly valued by many areas. Laser-based manufacturing processes cover laser welding [1], laser cutting [2], laser cladding [3], laser drilling [4], laser surface modification [5], laser additive manufacturing [6], etc. Enormous economic and social benefits have been derived with the rapid development of laser technology in the last few decades [7].

Although there is a variety of laser types, the elementary components of a laser system are the same. In general, a laser is composed of three basic components including active medium, excitation source (also known as pumping source) and optical resonator which is shown in Figure 2.1. The active medium is the source of optical gain inside a laser system. The appropriate active medium should be chosen to generate the laser which nowadays covers gas, fluid, solid and semiconductor. In order to realize the population

Chapter 2 Literature review of laser additive manufacturing inversion in active medium, methods should be adopted to excite the atom system and to increase the number of atoms of higher energy level. The common excitation sources include electrons, flash lamps, chemical reactions, lasers, ion beams etc. Light can be amplified by using an optical resonator which is composed of two mirrors with certain geometry and reflection characteristics. It is commonly a cavity composed of a 100% reflective mirror at one end and a partially transmitting mirror at the other end. The main functions of optical resonator include the provision of optical feedback capability which enables excited photons reflecting inside the cavity back and forth to obtain continuous coherent oscillations. Optical resonators also restrict the direction and frequency of the light oscillating inside and ensure the unidirectionality and monochromaticity of the output laser.

Figure 2.1 Basic construction of a laser system [8] Distribution law of black-body radiation was explained by Max Planck in 1900 using quantisation of radiation. In 1913, Niels Bohr presented the assumption of quantisation in the electron motion state, based on which in 1917 Einstein [9] deduced equations for black-body radiation from the concept of photons and proposed two important concepts for the first time: spontaneous emission and stimulated emission. The concept of stimulated emission was widely used in the development of laser technology 40 years later in the 1960s.

2.2.1 Spontaneous emission and stimulated emission

Basic assumptions for atomic theory should be introduced to understand the physics of the generation of laser light. All matter is made up of atoms, and the atomic system is in the status of discontinuous energy level. To simplify the model, a radiative process in a two-level system including a lower state E1 and a higher state E2 is proposed. When an electron on the orbit of E1 obtains some certain amount of energy from outside, it will jump to the outer orbit with higher energy level. The atom is now in an excited state and

Chapter 2 Literature review of laser additive manufacturing the energy of atom will increase. This phenomenon was identified by Einstein as absorption process which is schemed in Figure 2.2(a). The energy required for the transition is given by [10]

∗ ℎ 휈 = 퐸1 − 퐸2 (2.1)

Where ℎ∗ is Planck’s constant, 휈 is frequency of the electromagnetic radiation and 퐸 is the state energy. An atom will be unstable after being excited and will return to the lower energy state if no outside influence exists. At the same time, a photon with energy hν=E2-

E1 will be emitted spontaneously. This process is called spontaneous emission shown in Figure 2.2(b). When an atom in higher energy level decays spontaneously to a lower energy level, the photon emitted is random and the direction of radiation also differs. The radiation of different atom is diverse and hence the initial electromagnetic radiation is incoherent. When an atom in high energy level is stimulated by a photon with energy hν before spontaneous emission occurs, the atom will transit from excited state to lower energy field. Simultaneously a photon which has the same frequency and propagates in the same direction with the stimulating photon will be emitted. This phenomenon is defined as stimulated emission illustrated in Figure 2.2(c) and it is the reverse of absorption. One incident photon can stimulate the atom in excited state to emit one identical photon. Then these two photons can continue stimulating and bring about two more photons. With the action of one incident photon, an atomic system can gain a large number of identical photons, which makes amplification of light possible and hence coherent laser light can be obtained. The light amplification brought by stimulated emission is the basic concept of generation of laser.

E2 E2 E2 2 2 2 hν hν hν =E2-E1 hν hν =E2-E1 hν

1 1 1 E1 E1 E1 (a (b) (c)

) Figure 2.2 Three major processes of electromagnetic radiation: (a) absorption, (b) spontaneous emission, (c) stimulated emission

Chapter 2 Literature review of laser additive manufacturing

2.2.2 Population Inversion

It can be seen from the definition of spontaneous emission and stimulated emission, the former governs luminescent mechanism of common light source while for the generation of laser light the latter is the dominant method. To keep the laser running continuously, methods should be applied to change the distribution of the atomic system when in thermal equilibrium. Population inversion happens when a system exists in a state of which more members of the system are in higher states than in lower energy states. To realize population inversion, energy from outside should be delivered into the system to guarantee more atoms can transmit from a lower state to a higher state by absorbing energy. The process is called excitation or pumping.

2.2.3 Laser Cavity

The dominance of stimulated emission in an atomic system can be achieved by realising population inversion. However, the early radiant field must have at least one photon to trigger the active medium which is in excited state. The photon may come from the spontaneous emission of which the radiation is random and consequently the light after amplification is unordered and incoherent.

A laser cavity is generally composed of two parallel reflecting mirrors located at both sides of the working material. The output mirror is partially reflecting while the other is totally reflective to the laser radiation. Apart from the photons travelling along the axial direction, other photons will diverge away from the laser cavity quickly. Light is amplified when oscillating inside the laser cavity. Amplification continues until the monochromatic light with the same phase is emitted through the partially reflecting mirror. Although amplification of the light can be realized during oscillation, energy loss also occurs when mechanisms like absorption, refraction and projection happen at the reflecting mirrors.

2.3 Laser beam characteristics

In general, laser beam has four characteristics including monochromaticity, coherence, directionality and brightness. It can be concluded that laser has high photon degeneracy which means high coherent light intensity exists within a wide coherent volume.

Chapter 2 Literature review of laser additive manufacturing

Characteristics of laser can be attributed to the principle of stimulated emission and function of the laser cavity.

2.3.1 Monochromaticity

Laser cavity not only can amplify the stimulated emission but also limit the frequency of laser output. The light resonating in the laser cavity should comply with equation [11]

퐿 = 푛휆/2 (2.2) where 퐿 is the length of cthe avity, 휆 is the wavelength of the laser radiation and 푛 is an integer known as the mode order (푛=1,2,3,4….). Therefore, only light which satisfies the conditions can resonate in the laser cavity. Light from common source contains all kinds of wavelength and hence is a mixture of light with different wavelengths. Owing to the laser cavity, selected laser concentrates in a narrow spectral bands or frequencies within a certain range. For an electromagnetic wave, the spectral band is narrower with a longer coherence length and the monochromaticity is better.

2.3.2 Coherence

The coherence of light source is an important characteristic which refers to the correlation degree of two different vibrating points in space. Light coherence can be classified into two categories including temporal coherence and spatial coherence. Temporal coherence reflects the monochromaticity of light source, namely, the dispersion degree of the emitted photon frequency. The narrower discrete degree of the frequency is, the better monochromaticity of light source will be, and hence resulting in better temporal coherence. A simple relationship exists between coherence time 휏 and frequency range 훥휈 given by [11]

1 휏 = (2.3) 훥휈

Spatial coherence reflects the impact of light source size on coherence. Since the light is emitted from different parts of the source, source size plays an important role in light coherence.

Chapter 2 Literature review of laser additive manufacturing

2.3.3 Directionality

The resonating cavity will amplify those photons traversing in one certain direction while other photons will decay gradually which results in the excellent directionality of laser. The minimum beam divergence is constrained by diffraction effect and cannot be smaller than diffraction limit angle 휃푚 defined by [11]

휆 휃 = (2.4) 푚 2푎∙1000 where λ (nm) is beam wavelength and 푎 (mm) is beam waist radius. The directionality of laser greatly depends on the laser type and is closely related to the uniformity of active medium, length of the laser cavity and excitation method.

2.3.4 Brightness

In general, brightness is defined as the power transmitted per unit area per solid angle by the light source in a certain direction. The brightness of laser mainly depends on the highly concentrated light in the emission direction. Owing to the amplification effect of laser cavity, amplitude as well as the brightness of radiation wave will increase. The monochromatic brightness of laser 퐵휆 is given by [11]

푃 퐵휆 = 2 (2.5) 푎훥푣푠(휋휃0 ) where, 푃 is laser power, 푎 is laser beam waist area, 훥푣푠 is laser line width and 휃0 is divergence angle. In comparison with common light source, laser has better directionality

(smaller divergence angle 휃0) and better monochromaticity (smaller laser line width 훥푣푠). Therefore, laser has greater photon degeneracy and monochromatic brightness.

2.4 Types of lasers

Since the first ruby laser was invented in the 1960s, technology and application of laser have undergone rapid development. There are hundreds of active media available for laser systems which include crystal, glass, fibre, gas, liquid, semiconductor and free electron. Diversity also lies in the excitation method including flash lamps, electric excitation, thermal excitation, chemical reaction, etc. In general, different types of lasers can be categorized by three methods: active medium, operating mode including continuous wave

Chapter 2 Literature review of laser additive manufacturing

(CW) and pulsed wave, output characteristics which include atomic laser, ion laser and molecular laser. Solid state laser, gas laser, dye laser, fibre laser and diode laser are the most widely used lasers for material processing, among which fibre laser, diode laser and disk laser (one type of solid-state laser) are the most common for metal additive manufacturing process.

2.4.1 Diode Lasers

Diode lasers known as semiconductor lasers are becoming increasingly important applied in practical laser material processing. By the inducement of current flow, photons are generated when electrons from conduction band transmit to valence band combining with the holes. Population inversion which contributes to the number of electrons in conduction band exceeds the number of holes in valence is realized by applying a forward bias in the junction region. Optical cavity is composed of two opposite facets of the semiconductor formed by split and the reflectivity is around 35%. The advantages of diode lasers include small volume, long lifetime and that the electric current can be used directly as pumping source. Semiconductor lasers have high energy conversion efficiency which can reach up to 50%. Working voltage and current of diode laser system are compatible with integrated circuit and thus it can be integrated into the monolithic. High frequency can be adapted directly to modulate the current so that output laser with high- speed modulation can be achieved. On account of the advantages listed above, diode lasers have been widely used in laser communication, optical storage, optical gyro, laser printing, laser ranging, radar and many other fields. In comparison with the other type of lasers, oscillation mode of diode laser is waveguide dielectric resonator which is different from the optic open resonator. There are two material systems widely used in diode lasers.

One is based on GaAs and Ga1-xAlxAs, of which the output wavelength depends on the replacement percentage of Al and is approximately 0.85 µm. The other is based on InP and Ga1-xInxAs1-yPy and the common output wavelengths are 1.3 µm, 1.48 µm and 1.55 µm, among which wavelength of 1.55 µm is the most popular. Transmission loss of fibre optical towards light with 1.55 µm wavelength is small around 0.15 dB/km which makes high-speed fibre optical communication possible. Figure 2.3 shows the structure of a diode laser. Excitation-emission and amplification will take place in the active region. The electrons and the holes are restrained in the active region which provide advantages for the process of stimulated emission.

Chapter 2 Literature review of laser additive manufacturing

Figure 2.3 Schematic of a diode laser system [12]

2.4.2 Fibre lasers

Fibres mixed with rare earth element are used as active media in fibre laser systems. Fibre laser is essentially a special form of solid-state lasers. In comparison with conventional solid-state lasers of which active medium is in the block, fibre lasers have many advantages listed below. Firstly, pumping light is constrained in the fibre which contributes to a concentrated pumping energy and hence the threshold value for pumping source is relatively low. Secondly, though the gain per unit length is low, high single-pass gain can be achieved by applying fibres with long length and low loss rate. Thirdly, laser cavity used in single mode fibre laser has the characteristic of waveguide and hence the mode can be easily controlled. Lastly, heat dissipation of fibre material is excellent which is attributed to the high ratio between surface area and volume. Semiconductor lasers also known as diode lasers are the most popular pumping source used for fibre laser. Essentially, fibre laser system functions as a converter transforming pumping light to another laser light with a different wavelength and thus laser beam quality is much better compared with semiconductor lasers. Fibre lasers are widely applied in optical communications, optical sensor, laser material processing and laser printing. Figure 2.4 presented the principle of a cladding pumped fibre laser system.

Chapter 2 Literature review of laser additive manufacturing

Figure 2.4 Principle of a cladding pumped fibre laser [13]

2.4.3 Disk lasers

Disk lasers are solid-state lasers using a thin slice or disk of Ytterbium-doped Yttrium- Aluminium Garnet crystal as active medium. The disk laser system produces laser beam at 1030 nm wavelength, very close to that generated in the more common solid-state Nd:YAG laser of 1064 nm wavelength.

The active medium of a traditional Nd:YAG laser system is in the shape of a cylindrical rod of several millimetres in diameter, and hundreds of millimetres in length. Flash lamps or diodes are normally used as the pumping source. The optical efficiency of the pumping source converting the pump light into laser beam is low which will result in the temperature increase of rod, and hence a cooling system is required. The radially parabolic temperature profile will happen which is caused by the coolant flowing along the outside of the rod, which gives rise to the thermal lensing problem (active media being hotter on the beam axis). As a result, divergence of the laser beam and a reduction in beam quality will happen.

In 1993 Adolf Giesen and his group first developed disk laser system to overcome this problem. A very thin crystal disk with higher surface to volume ratio rather than cylindrical rod was used as the active medium. The thickness of the disk is normally only a few hundred microns and the diameter is several millimetres. Since the diameter of the pumping source is much greater than the disk thickness, axial cooling can be achieved, and thermal lensing can be reduced to a minimum promoting good beam quality. Despite the fact that the cooling effect of disk lasers are excellent because of the thin disk

Chapter 2 Literature review of laser additive manufacturing thickness, for the same reason, only a small fraction of the pump light can be absorbed when it passes through the thin disk. To increase the absorption of the pump light, a set of mirrors and retro-optics are installed around the disk to increase the number of times of the pump light passing through the disk.

Figure 2.5 Concept of a thin disk laser [14] As illustrated in Figure 2.5, the pump light is imaged first by a parabolic mirror onto the thin disk. The pump light which has not been absorbed by the disk will be reflected and recollimated by the parabolic mirrors and then shifted spatially by a mirror pair and again imaged to the disk.

2.5 Laser additive manufacturing

Laser additive manufacturing (LAM) is a breakthrough in advanced manufacturing technology field during the last few decades. With the assistance of computer-aided design (CAD) technology, a full density 3D product can be manufactured layer by layer in a laser system. The additive manufacturing is also known as instant manufacturing and free-form fabrication of which the principles are similar but process routines and functions of the facilities diverse.

Stereolithography, Selective Laser Sintering (SLS), Selective Laser Melting (SLM), Laminated Object Manufacture (LOM), Direct Metal Deposition (DMD), Laser Metal Deposition (LMD) are some of the main laser additive manufacturing technologies used during the industrial process. In comparison with other technologies, LMD has a significant advantage and tremendous potential owing to its high manufacturing efficiency and low fabricating cost. LMD also has its superiority of producing samples

Chapter 2 Literature review of laser additive manufacturing with large dimensions in comparison with SLM of which the product size is constrained by the size of facility chamber.

The following sections will further introduce the LMD process and the parameters applied during the process affecting the quality of final product. In addition, defects which may occur in the sample are discussed and solutions are put forward.

Different deposition methods will influence the final performance of the product including surface finish, microstructure and mechanical properties. Deposition efficiency also varies by applying different deposition ways.

In general, there are two techniques for LMD listed below [12]:

1. Two-step process 2. One-step process

The two-step process is also known as pre-placed deposition process. Wire or powder is placed on the substrate before laser scan the surface and form clad.

In the one-step process, metal material in the form of wire or powder is fed simultaneously with the motion of moving laser beam. There are three categories in one-step process including powder injection, wire feeding and combined wire and powder.

2.5.1 Pre-placed wire and powder

Pre-placed wire or powder deposition is a method that alloy cladding material is placed on the substrate surface by some means in advance. The laser beam then scans the alloy material in order to melt the metal together with some portion of the substrate. Because of the high cooling rate, molten metal will solidify rapidly and form a metallurgical bond on the substrate surface. Powell et al. [15] categorised several deposition stages with the increase of the energy input. The powder was melted rapidly by the laser before the heat transfered to the substrate. A great amount of the heat would be conducted to the substrate contributing to a partial solidification of the molten part. If laser source provided enough energy, heat would eventually transfer into the body substrate and form the final solidification. Pre-placed deposition is suitable for single-layer cladding since it will become difficult to place another layer of material on already deposited free-form geometry. Figure 2.6 illustrates the pre-placed powder deposition process.

Chapter 2 Literature review of laser additive manufacturing

Figure 2.6 Pre-placed powder direct metal deposition [16]

2.5.2 Powder injection

During laser metal deposition by powder injection, the metal powder is fed into heat zone directly where substrate and alloy material are melted simultaneously to form a layer of deposition after cooling. The powder is delivered by the inert gas which functions not only as a delivery carrier but also as the shielding gas preventing the deposition from oxidation. Adjustment of powder feed rate plays a significant role in controlling the quality of the final deposition product. If the feed rate of carrier gas is too small, friction between powders, powder and tube wall as well as the resistance of motion within powders cannot be overcome, which will lead to congestion of powder in the delivery device and tube. If the powder flow rate is too big, convection flow inside the pool will increase as a result of the momentum brought by the high-velocity gas flow. Heat dissipation increases while powder utilization rate will drop with unutilized powder adhered on deposition surface deteriorating the surface finish. Apart from above, carrier gas also can act as a source to protect the laser optics.

Material feeding angle is defined as the angle between the substrate and material entering the melt pool. In comparison with wire feeding deposition, powder feeding has a wider material feeding angle varying from 0° to 180° while it is limited approximately from 10° to 75° for wire feeding [17]. Lin [18] found that powder utilization could be increased by reducing the particle size, process velocity and bonding temperature.

In general, two categories of powder feeding concepts are used during manufacturing including lateral powder injection and coaxial powder injection.

Chapter 2 Literature review of laser additive manufacturing

2.5.2.1 Lateral powder injection Powder nozzle is positioned off-axial to the laser beam shown in Figure 2.7. Material feeding angle, as well as the distance between nozzle and workpiece, determines the position of the powder injection direction which is a vital role in lateral powder deposition. An increased angle leads to a decreased cross-section of the powder stream and hence increases the powder utilization [19].

Figure 2.7 Lateral powder injection [12] In lateral powder injection, feeding direction is another key factor influencing the deposition result. Powder feeding from the rear which means the feeding direction is away from the scan direction results in less oxidation and better surface finish while for front feeding direction, better powder utilization can be achieved since more powder is injected directly into the melt pool [17].

2.5.2.2 Coaxial powder injection Coaxial powder injection uses a coaxial nozzle where powder streams surround the laser beam shown in Figure 2.8. The temperature of powder is relatively high because of the interaction with laser beam before impinging on the substrate. In coaxial nozzle, several individual powder streams are involved and hence achieving a homogeneous distribution of powder particles is important to the final quality of surface finish [12].

Chapter 2 Literature review of laser additive manufacturing

Figure 2.8 Coaxial powder injection [20] Zhu and Li [20] found in their work that the difference of the deposited layer’s building height could be compensated automatically with powder focused below the substrate and laser focused above the substrate. They called it self-regulation effect. Pinkerton and Li [21] concluded from their experiments that deposition point standoff played a significant role in determining final part geometry. They found that better consistency in layer width could be achieved with no substrate movement and thus adjustment of the substrate was not always necessary between layers. In comparison with the lateral delivery method, coaxial nozzle was preferred when high precision deposition of sensitive materials was required while lateral nozzle was suitable for rotationally symmetric parts of which the accessibility is limited [22].

2.5.3 Wire feeding

During wire feeding deposition process, the wire material is normally fed from the side of the laser and delivered into the melt pool on the substrate or already deposited layer. With the moving of laser head in a desired path on the substrate, a fully dense metal particle can be produced.

Figure 2.9 Laser additive manufacturing using wire, (a) front feeding, (b) rear feeding, wire placed at (c) leading edge, (d) centre and (e) trailing edge of melt pool [23, 24]

Chapter 2 Literature review of laser additive manufacturing

Syed and Li [23] investigated the effect of wire feeding direction and location on deposition result. They concluded that better surface finish could be achieved by using front feeding and placing the wire at the leading edge of the melt pool, while for rear feeding the best results were obtained by placing the wire at the trailing edge. They also found that surface roughness could be reduced by decreasing the feeding angle for front feeding and increasing the feeding angle for rear feeding.

Mok et al. [25] carried out deposition of Ti6Al4V wire with high power diode laser and they found that the better deposition consistent could be realised when using front wire feeding with 45° feeding angle. They also discovered that the size of epitaxial columnar grains in the sample would become larger when higher laser power was applied. Syed et al. [17] concluded from their experiments that the surface roughness of powder-feed samples was between 70 and 90 μm (Ra) which was 20%-30% higher than wire fed samples ranging from 40 and 60 μm (Ra). This was due to the fact that adherent powder particles might stick on the solidified surface and coarsen the surface. Another reason was the instability of melt pool caused by varying mass flow distribution would change the final geometry of track. Surface overlaps and protrusions could be avoided by elevating the defocusing distance together with the filler wire in order to melt the wire before touching the previous deposition surface [26].

2.5.4 Combined wire and powder

Combined wire and powder deposition is a method that integrates coaxial powder injection with wire feeding. A wire feeding tube is tilted on one side of the coaxial powder nozzle delivering the metal wire into the melt pool. Shielding gas like Argon can be sent either through the coaxial nozzle or from the side shielding tube (shown in Figure 2.10). Syed et al. [27] concluded from their experiments that deposition efficiency and surface finish could be improved due to a higher energy absorption by applying the combined method. Porosity was found to be 20%-30% less in comparison with using powder injection alone. By controlling the feed rates of wire and powder separately using two different materials, samples with gradient change of chemical composition could be obtained in a single pass. Deposition profile and dilution, material distribution as well as dendrite spacing (distance between the dendrite arms) could be controlled to certain degrees by varying the feed rates [28].

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Figure 2.10 Combined wire and powder deposition [27] 2.6 Process Parameters

2.6.1 Laser power and specific energy

The melt pool temperature is significantly influenced by the selection of laser power. Without enough laser power input, discontinuity and uneven surface may appear because of incomplete fusion between the substrate and deposited material. With the increase of laser power, more metal will be melted and dilution will increase which leads to less discontinuity. When the laser power keeps rising, excessive power density will result in high dilution and hence resulting in deformation of material surface, cracking and vaporization of the material where porosities are more likely to appear.

Microstructures and mechanical properties of deposition samples were compared by Zhang et al. [29] when varying laser power value while keeping other process parameters unchanged. They found that the microstructure became coarser with the increase of laser power owing to a bigger melt pool and hence lower cooling rate and longer time for the dendrite to grow. Mechanical properties were closely related to the deposition microstructure. Refined grains could result in better grain strength and consequently improved structure hardness. Thus, finer grain size could be formed by applying a lower laser power and faster cooling rate could contribute to a higher hardness, yield strength and tensile strength.

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In general, laser power is always considered together with spot diameter and process velocity. Specific energy is defined as 퐸 = 푃/(퐷푉) [30], where P is laser power, D is spot diameter and V is process velocity. Jang et al [31] showed that the ratio between penetrated depth and height deposited could be used to estimate the criterion for specific energy. Zhang et al [32] pointed out that the deposition height firstly increased and then decreased with increasing of the specific energy. They concluded that with a low specific energy, metal powder as well as the previous deposition layer could not be sufficiently melted while excessive specific energy may overheat the substrate preventing powder from absorbing laser energy. Zhu et al [33] also explained that some part of liquid metal would be vaporized and then ionized into plasma which might stop metal powder from absorbing energy effectively if the specific energy was too big. Deposition height would decrease and metal powder unutilized would adhere on the surface resulting in poor surface finish which was also mentioned by Ma et al [34]. Excessive specific energy would enlarge the melt pool and irregularity of deposition geometry might occur. To avoid the difference between contour height and inner height of the sample, an appropriate range of specific energy should be selected [35].

2.6.2 Process velocity

During laser metal deposition, the process velocity of the laser beam is another key factor influencing the final quality of the component. When other process parameters are kept constant, process velocity determines the energy absorption efficiency and thus affects the surface quality of the cladding layer, melt pool temperature as well as the dilution ratio. When applying an excessive process velocity, the main defect that may occur is the discontinuity of layer track because the laser beam energy cannot be sufficiently absorbed by the substrate and the metal powder which also leads to a poor fusion between them. When the process velocity is too low, energy absorbed by the surface per unit time is high which will contribute to a severely diluted substrate. Melt pool is overheated and porosity may appear attributing to the vaporised metal of low boiling point.

It was concluded by Mahmood et al. [36] that physical, mechanical, metallurgical and tribological properties could be affected and properly controlled by adjusting the process velocity. They pointed out that the height of deposition initially increased and then decreased when increasing the process velocity. It was due to the fact that when low process velocity was applied, the substrate was overheated which would contribute to

Chapter 2 Literature review of laser additive manufacturing high dilution and low deposition height while less powder would be utilized when process velocity was too high. They found that wear resistance and surface roughness both increased with the increase of process velocity till 0.065 m/s, after which they both began to drop. A model was developed by Cheikh et al. [37] to predict the geometrical characteristics including layer height and width when varying process parameters. The 1/4 height was found proportional to (푃/푄푚) (푄푚/푉), where P is laser power, 푄푚 is powder mass rate and V is process velocity. Layer width was affected by laser power together with process velocity and found to be proportional to 푃(3/4)푉−(1/4) , which indicated that thermal effects could determine the deposition width. Liu et al. [38] investigated the effect of process parameters on wall thickness, powder efficiency and deposition rate. They concluded that a decrease in process velocity would lead to an increase in wall thickness, powder utilization and over all deposition rate. Kelbassa et al [39]. found that by placing the laser beam focal point as well as powder focus slightly above the substrate material, up to 90% of laser energy could be absorbed and thus high- speed laser metal deposition could be realized.

2.6.3 Gas-powder flow rate

Carrier gas plays an indispensable role in transportation of the metal powder. The flow rate of carrier gas directly determines the powder delivery quality. Congestion in the tube and discontinuous delivery of powder may occur if the gas flow rate is not high enough which finally results in a poor surface finish. However, Powder utilization will decrease when using an excessive gas flow rate.

An observation system was developed by Tan et al. [40] to investigate powder feed behaviours. They observed from their experiments that the velocity of powder was largely dependent on the flow rate of the carrier gas instead of the powder flow rate. With the increase of powder particle speed, smoother surface finish could be obtained because it was more likely for the powder to adhere on the deposition surface when a low particle velocity was applied. The deposition height would decrease when increasing the powder feed rate. They also pointed out that the powder mass concentration was closely related to the distance below the nozzle and the peak value appeared above the powder focal point. Zekovic et al. [41] showed that flow in the nozzle was a combination of two phases including carrier gas as primary phase and powder particles as secondary phase. They reported the importance of an additional gas flow for protection of laser lens especially 44

Chapter 2 Literature review of laser additive manufacturing when turbulence and powder bouncing were prominent. By modelling and comparing with experiment data, they noted that selection of standoff distance was important to powder concentration distribution and was pivotal to the final component geometry. Powder concentration was found to be closely related to gas flow velocities [42]. In his study, Lin proposed that the powder concentration would reduce, and the location of peak value would shift towards the nozzle exit when increasing the powder flow velocity. Wen et al. [43] investigated powder concentration distribution and classified it into three stages, including pre-waist stage, waist stage and post-waist stage shown in Figure 2.11. An ideal deposition plane could be predicted by considering both powder concentration and laser beam distribution. Temperature would reach its peak in the centre of converging point and decrease in the radial direction. Temperature distribution underwent alteration from di-modal to single-modal with powder stream becoming diverging.

Figure 2.11 Powder stream structure (a) pre-waist stage, (b) waist stage, (c) post-waist stage [43]

2.6.4 Overlap

During multi-pass deposition process, a small offset distance exists between two adjacent passes. Overlap is defined as 푅표푣푒푟푙푎푝 = (1 − 퐿/퐷). 100%, where 퐷 is the laser spot size and 퐿 is the offset distance of laser spot centre between the two adjacent passes. Selection of the overlap is essential in additive manufacture because it is closely related to surface finish, residual stress distribution and microstructure of the final component.

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Cao et al. [44] concluded from their experimental work that the residual stress in the overlap area was much higher and more intensely fluctuating in comparison with the inner-pass area. The residual stress range including the peak value and mean value was much wider with the increase of overlap. Finer and more evenly distributed grains could be obtained with a higher overlap due to the locational change of nucleation from overlap area alone to both overlap area and bottom area of the deposited layer. It could be observed from Figure 2.12 that when increasing the overlap from 20% to 50%, the average of grain size would decrease from 76.5 µm to 62.7 µm.

Figure 2.12 Average of grain size vs. Overlap [44] Li and Ma [45] observed from their experiments that when increasing the overlap, the surface roughness would decrease in an oscillating manner fluctuating between two lines of ΔH=h(1-k) and ΔH=0.24h(1-k), where h is the height of one single layer, k is overlap, H is the minimum height of the overlapped layer and ΔH is the difference between the maximum and minimum height of the overlapped layer. With the increase of overlap, melting ratio would firstly decrease and then increase [46], where melting ratio is defined as the ratio of energy required to melt the material to the energy incident on the workpiece [47]. Zhang et al. [29] mentioned that an ideal overlap could be achieved when area A1 (shown in Figure 2.13) which was the concave groove between two adjacent beads equalled to area A2 which was the overlapping part, where S is the offset distance of laser spot centre, W is laser spot size and H is bead height. From their study, 31% overlap was used and samples could be deposited with uniform clad geometry.

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Figure 2.13 A schematic view of overlap between adjacent bead [29]

2.6.5 Deposition pattern

During laser metal deposition process, both mechanical and metallurgical properties are affected by selection of deposition trajectory. In order to obtain a uniform surface finish and reduce the accumulation of residual stress, it is necessary to choose an appropriate deposition pattern.

He et al. [48] compared two deposition ways to investigate how deposition pattern would influence the collapse of edges: inside-to-edge method referred to depositing inner passes before the pass on the sample edge, while edge-to-inside method referred to depositing the edge pass prior to the inner passes. They concluded that the reason for the inclination on edges was the overlap effect between adjacent tracks. In comparison with the inside- to-edge deposition pattern, the inclination of melt pool was larger for edge-to-inside deposition pattern. In inside-to-edge deposition pattern, inclination would increase with the addition of the layer height and the edge began to collapse when the inclination of melt pool increased from 25° to 32°. A new deposition pattern (d in Figure 2.14) was proposed by Yu et al. [49] to minimise the part distortion. Among the four deposition methods, least distortion was obtained when using fracture deposition pattern owing to the lowest temperature gradient. An improved interior structure, as well as mechanical properties, could be achieved when overlap was increased to be 50%.

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Figure 2.14 Deposition patterns (a) raster, (b) offsetout, (c) offsetin, (d) fracture [49] By introducing a modified temperature gradient mechanism, Gao et al. [50] pointed out that the deformation of substrate on the building up direction was much larger than traverse and longitudinal directions. When fabricating samples with symmetrical geometries, deformation was closely related to the dimension of the sample in traverse and longitudinal directions. Distortion of the substrate was more likely to happen at first few layers of deposition owing to the strong heat conduction between the substrate and cladding layers in initial stages. They also concluded that by applying a to and fro deposition pattern along the short dimension direction, minimum deformation and reduced movements of the work table could be achieved. Liu et al. [51] compared two deposition patterns to investigate the influence on microstructures and mechanical properties: single direction raster scanning (SDRS) and cross direction raster scanning (CDRS). CDRS had finer deposited grains in comparison with SDRS and recrystallized grains after heat treatment was more uniform when using CDRS. CDRS sample had a better ductility than SDRS owing to a more homogeneity of grain size. Lu et al. [52] compared three different deposition methods including offset contour line mode, parallel line mode and hybrid path method. They found that high precision could be achieved when using the hybrid mode when building particles with complex structures.

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2.7 Mechanical and thermal behaviour

2.7.1 Stress

In general, there are two categories of stress existing in laser metal deposition process: one is transient stress and another is residual stress. Transient stress includes thermal stress and structural stress which result from the thermal behaviour brought by high laser beam energy. In particular, thermal stress is mainly caused by thermal gradient after the metal material is heated while structural stress attributes to phase change arising from the melting of metal after being heated. Melt pool region will be rapidly heated and partially melted owing to the short interact time between the laser beam and metal material. Elastic thermal stress will occur because thermal expansion arising from the heating effect is restricted by surrounding area with lower temperature. On the other hand, yield strength will decrease with the rising of temperature and will be exceeded by compound stress. Thus, plastic thermal compression will appear around the melt pool. After the deposition process is stopped and temperature gradient no longer exists, residual deformation cannot disappear by itself which contributes to a small shrinkage and it is also restrained by substrate material. This is how residual stress occurs in the material during laser metal deposition process. The most common way to eliminate residual stress is heat treatment including stress relief annealing, during which the component is heated to a certain temperature and kept for some period. With the increase of temperature, yield strength decreases and will be exceeded by residual stress resulting in local deformation and hence residual stress can be reduced.

Klingbeil et al. [53] pointed out that warping caused by residual stress was a fundamental obstacle in laser metal deposition process. After modelling and performing some experiments, they suggested that by preheating the substrate before deposition process and insulating the substrate during the process could substantially provide the payoffs in reducing warping effect. Subtle changes in 3D part constraints, as well as optimisation of deposition paths, could compensate the warping caused by residual stress. Finnie et al. [54] proposed that residual stress could be altered from tension to compression by preheating the substrate. Alimardani et al. [55] reported that melt pool temperature would increase after preheating the substrate and hence laser power could be reduced to keep the melt pool in proper condition which consequently would lead to a drop in temporal

Chapter 2 Literature review of laser additive manufacturing thermal stress. A three-dimensional finite element model was developed by Wang et al. [56] and used to predict the thermal history and residual stress distribution. They concluded that compressive stress was found to be in the centre of deposition area and tensile stress at both ends of the plate on building up direction while for longitudinal direction, stress changed from compressive to tensile when increasing process velocity. Residual stress along the building up direction would increase when a higher laser power was applied. Neutron diffraction and contour method were used by Rangaswamy et al. [57, 58] to investigate the residual stress distribution in both thin wall and pillar structure. It was found that in most parts of the samples, residual stress appeared to be uniaxial, with compression in the centre and tensile stress near the edges as shown in Figure 2.15. Along the deposition direction in the centre line, residual stress was compressive and directed to the growing direction between 5 and 80 mm from the free end. The stress would decrease rapidly when approaching the free end. The distribution of residual stress showed relatively complicated situation near the substrate with tension directed to traverse and longitudinal direction while compression directed to longitudinal direction which could be attributed to the enhanced thermal behaviour including reheating and cooling. At the middle-height plane, stress along longitudinal direction appeared to be compressive in the centre and tensile near the edges owing to the temperature difference between boundary and interior.

Figure 2.15 Stress in x, y, z-direction (a) along the centre line in deposition direction (b) at the middle height of wall in width (Y) direction [57]

Stresses were also found to be tensile near the vertical free surface by Moat et al. [59] which were balanced with the compressive stresses in the interior area. They varied the laser pulse parameters including pulse lengths (푡푝푢푙푠푒) and cycle lengths (푡푐푦푐푙푒) in order to investigate the effect on stress distribution. By increasing the pulse lengths (푡푝푢푙푠푒) and

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duty cycle (푡푝푢푙푠푒/푡푐푦푐푙푒), stress gradient and size of the maximum tensile stress region would increase in the longitudinal direction. By varying laser power and process velocity during laser metal deposition process, Pratt et al. [60] concluded that no significant trend appeared between stress magnitude and process velocity while a slight increase of stress magnitude occurred when increasing the laser power input.

2.7.2 Porosity

Porosity is defined as the ratio of volume of pores to the total deposition volume. During the laser metal deposition process, defects like porosities will not only directly influence the mechanical properties of the components like wear resistance and hardness but also will be the location where stress concentrates and thus leading to the formation of cracks. However, sometimes the internal porosity might become beneficial for functional, electrochemical, or biomedical applications, where mechanical stress is not an issue. In general, two categories of porosity may appear in the deposition, including lack of fusion and gas porosity shown in Figure 2.16. Lack of fusion with irregular shapes mainly attributes to the insufficient re-melting between layers. Before the deposition process happens, powder material or substrate may have been oxidised. Gas like O2 and H2 will be generated when oxidation and carbohydrates are being heated by the high-energy laser beam. Owing to the rapid melting and solidification rate, bubbles can be developed during the cooling stage and become porosities if cannot float up to the top surface in time. In order to control the gas porosity that might occur during the process, methods like drying should be adopted to prevent powder materials from being oxidised during delivery.

Kobryn et al. [61] discovered that porosity tended to depend on the substrate thickness. Lack of fusion was more likely to appear in a thin substrate while gas porosity occurred more when the substrate was thick. They also pointed out that the number of lack of fusion and gas porosity would increase when decreasing process velocity and laser power. With a higher laser power, more energy was provided to guarantee that the powder could be sufficiently melted while an increased process velocity would lead to less material and gas to be fed into the melt pool, which was also reported by Katayama et al. [62]. In order to reduce the interaction time between powder and the laser beam and reduce the gas entrapment, Majumdar et al. [63] suggested that the method of increasing process velocity was an effective choice to eliminate the porosity content in the microstructure.

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Figure 2.16 Lack of fusion and residual gas porosity [61] Susan et al. [64] pointed out that small pores could be found within the powders if they were gas-atomized, most of which were developed due to the entrapment of the atomising gas like argon. They concluded that powders with higher porosity content (ratio of average pore size to average powder size) would result in an increased porosity number after the deposition process. They also reported that the importance of surface re-melting process when no powder was fed into the melt pool. It not only could eliminate the porosity between layers but could also improve the surface roughness. Compared with gas atomised powders, Ahsan et al. [65] reported that three times less porosity content was found in plasma atomised powder and much lower chance of interlayer porosity was observed in the deposition. Waheed et al. [17] investigated the difference between powder feeding and wire feeding during direct laser deposition. Higher porosity was observed in powder feeding deposition which might be due to the more disturbed flow existing in the powder-feed melt pools. It was suggested by Choi and Chang [66] that layer thickness as well as powder flow rate should be optimised to ensure that sufficient energy was delivered to re-melt the previous layer and thus lack of fusion could be avoided. Pores in the same layer could be eliminated by increasing the overlap. Goswami et al. [67] discussed that porosity could be reduced by optimising process parameters like process velocity and pulse duration. Dissolved gases would increase when process velocity was low or high pulse duration was applied, which consequently would lead to the increase of porosity. With a high process velocity or short pulse duration, gas could not escape to the surface because of the rapid solidification rate.

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2.7.3 Cracking and distortion

In the solidification front, the new-born developed dendrites are closely interconnected and form a crystalline solid network enclosing metal liquid between dendrites. Not enough liquid will be provided as the supplement in the cooling stage afterwards because of shrinkage, thus contributes to the formation of cracking initiations between the dendrites where stress concentrates. With the decrease in temperature, the stress will increase and crack will propagate along the dendrites. Owing to the complexity of temperature gradient when cooling down, multi-directional growth of dendrites will occur in different areas. Intense collision appears between dendrites of different growth direction during rapid cooling and high stress will arise in the eutectic structure, leading to the formation of macrocracking. Cracking initiation is more likely to happen in three types of locations: inclusions as well as some hard phase in deposition layer, micropores existing in eutectic structure and small pores located between deposition layer and substrate. The inclination of cracking effect can be reduced by relieving the residual stress after the deposition process. In order to control and prevent the cracking from happening, substrate and deposition material of matched thermal expansion coefficient as well as heat capacity should be chosen. Wettability between substrate and deposition material can be enhanced by adding transitional coatings and consequently cracking initiation will be reduced. Fallah et al. [68] suggested that by locally preheating the substrate before deposition, a more uniform dendrite structure and even temperature distribution could be obtained and thus formation of crack could be prevented. Alimardani et al. [69] also investigated the effect of preheating on the crack formation and they concluded that lower temperature gradients, as well as lower cooling rates, could be achieved when preheating the substrate prior to the deposition process. Thermal stress could be successfully controlled and a crack-free outcome could be realized. Li et al. [70] discussed the importance of carbon content on microstructure and cracking susceptibility when using Fe-based metal powder. They concluded that with a carbon content no more than 0.3 wt.%, the formation of cracking could be resisted and a reduction of cracking susceptibility could be achieved when decreasing the process velocity. Studies from Song et al. [71-73] suggested that elements like Co, Ni and Hf within a proper amount could effectively reduce the cracking susceptibility when added to the powder material. A significant difference in thermal fatigue was found by Khalid et al. when depositing steel on copper with or without using

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316 SS as the buffer layer. They proposed that with 316 SS as the buffer layer, thermal fatigue resistance was greatly enhanced which led to a decrease in the formation of catastrophic cracks in comparison with depositing steel directly on copper. Apart from crack, distortion is another inevitable defect which is mainly caused by transient stress. Dai and Shaw [74] investigated the effect of deposition pattern on distortion of the components. They discovered that with long deposition direction parallel to a single axis, distortion of concave downward occurred in the deposition direction while concave upward appeared in the orientation perpendicular to the deposition direction. However, distortion could be compensated with around 30% less when deposition direction changed 90o at every corner frequently (offset).

2.7.4 Microstructure

Deposition microstructure was strongly affected by the ratio between temperature gradient and solidification rate [12, 75, 76]. Cheikh et al. [75] pointed out that the solidification front would tend to be stable when the ratio was high and otherwise instability would occur, which could explain the appearance of columnar dendrite in the bottom of deposit with a higher ratio while short dendrites in the upper region of deposition with lower ratio as shown in Figure 2.17. They also proposed that the dendrite zone on top of each layer would disappear owing to the re-melting effect between successive layers.

Figure 2.17 Microstructure change from bottom to the top layer of deposition [75] Research from Majumdar et al. [63] investigated the influence of process parameters on microstructure characterisations. They concluded that equiaxed dendrites appeared on the top surface owing to a uniform heat flow from all directions while columnar microstructures appeared near the substrate area. Coarse grains were also observed

Chapter 2 Literature review of laser additive manufacturing between two adjacent layers because of the re-melting effect by the conducted heat from the solidified pool of next layer. Grain size was found to be coarsened with increased power density and decreased process velocity while the grain size range tended to be bigger with a higher process velocity applied. Wu et al. [77] showed that columnar grains grew in the first ten layers indicating that heat conduction through the substrate dominated the rapid cooling process. With the increase of deposited height and rising deposition temperature, microstructure became coarse in all areas and hence the microstructure brought by re-melting effect between successive layers became negligible. They also noticed that length of columnar grains would decrease when increasing laser power which might attribute to higher substrate temperature and thus smaller thermal gradient. Dinda et al. [78] carried out a series of experiments by varying process parameters and they reported that growth direction of columnar dendrites was closely related to the laser deposition direction seen in Figure 2.18. Figure 2.18 (b) presented an identical deposition direction for each layer from left to right, while for Figure 2.18 (a) the deposition direction changed by 180° after each was deposited. They pointed out that the growth direction of dendrites was opposite to the actual heat flux direction which was approximately towards the substrate, solidification boundary and trailing end of the melt pool. They also noticed the microstructure difference between bottom part of the deposition and top layers which could be attributed to the difference in cooling rate and solidification velocity. Effect of post-processing heat treatment was investigated after the deposition process. They discovered that when increasing the annealing temperature from 1000oC to 1200oC, dendritic characters of microstructure disappeared and fully equiaxed grain structure with inhomogeneous grain size distribution was developed.

Figure 2.18 Optical photos of longitudinal section for different deposition patterns [78]

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Guan et al. [79] investigated the influence of overlapping rate on microstructure evolution during laser metal deposition. In the overlapped area, microstructure changed from cellular grains to equiaxed dendrites by increasing the overlapping rate.

2.7.5 Surface unevenness

Waviness phenomenon happening during the processes of laser metal deposition and weld bead formation has been reported by many people. Bulge or collapsing at both ends are observed during the laser metal deposition process while humping defined as a periodic undulation of the weld is one of the common defects for the welding process.

Su et al. [80] discovered some bulge effect at the beginning of their deposition and addressed that it could be explained by the fact that the temperature increased rapidly at the beginning of the process when laser beam hit on the substrate and melted the material. As a result, the temperature gradient in the melt pool became higher which at the same time increased the surface tension force. Since the surface tension force was proportional to the difference of free surface energy between substrate and melt pool, a spherically shaped deposition was formed to reduce the surface tension. Substrate temperature rose after laser beam started to move which would contribute to a lower thermal gradient. Therefore, the bulge effect would disappear once the process became stable. Similar result was found by Morgan et al. [81] when a re-melted pre-placed powder bed process was investigated. To improve the surface quality of laser cladding, Liu and Li [82] developed an in-time motion control system to detect the deposition dimension and make some compensation correspondently. They noticed that sometimes bulge would occur on the deposition surface where newly added material would accumulate resulting in an increased melt pool width and total collapsing of the clad bead.

Figure 2.19 Bulge effect at the beginning of deposition [80] 56

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A closed online monitoring system was developed by Heralic et al. [83] to improve the bead quality during laser metal wire deposition process. In order to keep the whole process stable, they pointed out that hand-tuned adjustment was required during the starting and end of the deposition. Wall height inconsistency was also observed by Martina et al. [84] where bulge was found at the beginning and collapsing towards the end. From their perspective, this phenomenon might due to the fact that base material was cold when deposition started, and more heat was built up at the end caused by weaker thermal dissipation to the surrounding material. To solve this problem, they mentioned that the depression issue could be addressed by decreasing the energy input towards the end. Humping and collapsing phenomenon could also be compensated by swapping the starting and end sequence of the deposition between two layers. Lee et al. [85] concluded from their numerical work that bugles at the beginning of deposition could be reduced by changing the surface tension gradient of the material. A similar mushroom-shaped deposition shown in Figure 2.20 could be found for both positive and negative gradients while a relative small bulge could be achieved if material with a mixed surface tension gradient was used.

Figure 2.20 Bead geometry at the beginning of the deposition with different surface tension gradient (a) Negative, (b) positive, (c) Mixed [85]

He et al. [48] discussed the role of gravity, shielding gas flow, supporting force from the substrate and surface tension force during formation of the melt pool when it was on the substrate edge. They pointed out that deposition pattern and depositing sequence would have a huge impact on the edge collapsing, especially when determining inclination angle of the melt pool shown in Figure 2.21, where X1 and X2 is the overlap area, 훼 and 훽 indicate the left and right side-surface of the same wall and 휃 is inclination angle. Process parameters also have a profound influence on formation of the deposition regarding the quality of the edge and side surface. Specific energy defined as 퐸 = 푃/(퐷푉) was found to be the crucial factor for height/width ratio of the cladding result [35], where P is laser

Chapter 2 Literature review of laser additive manufacturing power, D is spot diameter and V is process velocity. They also found it quite interesting that although power density was kept constant, the deposition contour was more likely to collapse when a higher laser power was applied.

Figure 2.21 Melt pool on the edge with different deposition sequence [48] Some nodulation was noticed by Pi et al. [86] at the beginning and end of their sample. They reported that it might be due to acceleration and deceleration effects of the moving stage at both ends, where total time of travelling for that short distance was longer and hence more material would be deposited. To solve this problem, they kept the laser off until the process velocity reached a stable state at the beginning and shut the laser off before the deposition actually stopped. A three-dimensional model was built by Alimardani et al. [87, 88] to investigate the temperature and thermal stress distributions during multilayer laser solid freeform fabrication process. They observed from their experimental and numerical results that the deposition height increased at both ends of the sample, which could be attributed to the higher maximum temperature in that region. As a result, a larger melt pool would form where more powder material would be caught and deposited. Height increase at both ends of the deposition was also noticed by Morville et al. [89]. They suggested that this phenomenon was due to the fact of surface tension. Melt pool length was short at the beginning of the deposition and added material would only have limited flow distance to travel within the melt pool, which contributed to the height increase when starting. Molten material tended to flow towards the rear region of the melt pool with increased melt pool length. After steady state was reached where melt pool length became constant, newly added material would form a deposition with constant height. At the end of the process, melt pool length would become larger again which would result in a higher track height in the end. Wang and Felicelli [90] demonstrated a FEM model of multiple layer laser deposition with SS410 steel. Heat transfer model was coupled with phase transfer analysis and they observed that low

Chapter 2 Literature review of laser additive manufacturing percentage of austenite was shown at the beginning of the deposition meaning that there was not enough heat input to melt all the added material. They attributed this phenomenon to the fact that this region experienced a higher cooling rate as a result of longer time of cooling between layers. They suggested that higher laser power should be used at the beginning to ensure a full melting of newly added material and a lower power applied when approaching the end.

Meng et al. [91] built a numerical model in a recent research of humping defect in high- speed gas tungsten arc welding shown in Figure 2.22. Arc shear stress was found to be the dominant driving force for deformation of the melt pool surface while Marangoni stress and gradational force had little effect due to the fact that strong convective flow caused by the arc shear stress would contribute to a small temperature gradient in the melt pool.

Figure 2.22 Simulation of humping effect in high-speed gas tungsten arc welding [91] Raleigh’s theory of instability model was applied by Gratzket et al. [92] to explain the humping phenomenon. They described that the molten metal would form into a liquid cylinder due to the effect of surface tension on the elongated melt pool, reducing the surface area of liquid metal. From their point of view, the humping could be attributed to the break of the liquid cylinder when the balance between surface tension and gravity force was interrupted. They also reported that width to length ratio of the melt pool should be bigger than 1/2π to avoid the humping effect.

Chapter 2 Literature review of laser additive manufacturing

2.7.6 Re-melting

Laser re-melting is a process using the laser beam to melt the sample surface without adding any feeding material in order to improve the surface properties. Rombouts et al. [93] mentioned that the appearance of “staircase effect” during additive manufacturing process was determined by the combined factors of process parameters, sample geometry and characteristics of additional material. A reduction of 5μ in surface roughness was achieved when sufficient heat input was used. They also discussed the relationship between surface roughness on the side wall and slope angle of the geometry. The result indicated that the surface quality could be improved more than one order of magnitude and the dramatic increase of surface roughness as a result of slope angle could be avoided by re-melting process. Yasa et al. [94] investigated the impact of re-melting on sample density, surface roughness, hardness and microstructure during SLM process. They pointed out that the relative density could be improved from 93.5% to 99.2% when re- melting was conducted between two layers. Porosities were less likely to appear after re- melting which could lead to a fully dense shell. Vickers microhardness was also improved (from 403 ± 11 to 444 ± 8 HV) on the sample surface as a result of microstructure transformation into fine-cell size after re-melting. From another research by Yasa and Kruth [95], stair effect was found to be prominently eliminated by re-melting. The first step was to re-melt structure contours between two layers. After the sample was fully built, the second step was applied where re-melting process was targeted on the inclined surface. Apart from additive manufacturing, re-melting was also widely used in other processes. Frostevarg et al. [96] reported that undercut (appears in the form of a groove or crater near the weld toe) on the sample surface could be reduced and weld root profile could be improved with the help of re-melting. They recommended that a wider but shallower melt pool should be used by applying a defocused laser beam for both top and bottom surface since there would be no need to achieve a deep penetration. Gerritsen et al. [97] concluded from their work that re-melting could be performed on weld toes with a certain range of heat input. Welding performance was found to improve in the aspect of hardness, residual stress and fatigue near the region of weld toes after re-melting. During laser cladding process of Fe-based material, Gao et al. [98] compared sample properties before and after re-melting. Better surface quality, higher microhardness and porosity- free between overlapping regions were achieved after re-melting. When laser cladding Ni–Cr–Mo alloy on Q235 steel, Wang et al. [99] found that microstructure of the coating

Chapter 2 Literature review of laser additive manufacturing could be refined and become uniform after re-melting where the number and scale of the defects could both decrease. Corrosion resistance was also improved with higher corrosion potential and lower corrosion current density in comparison with the coating without re-melting.

In addition to experimental research, numerical modelling is widely implemented to help understand the re-melting process. A three-dimensional model was built by Wang et al. [100] for predicting the temperature field and melt pool formation during plasma-based re-melting process. The simulation results indicated that the melt pool formation was determined by the combined factors of jet-inlet velocity, the relative speed between heat input and sample and Marangoni convection. Influence of shear stress on fluid flow brought by impinging plasma jet was found to be predominant only when the velocity of inlet plasma was high. Yilbas et al. [101] developed a FEM model to investigate the change of residual stress distribution after re-melting during HVOF coating process. The residual stress level was observed to increase both from experimental and modelling results despite the fact that defects including porosities, cracks and unevenness on the coating surface were mostly eliminated after re-melting. Cellular microstructures were found in the processed region and the microhardness was enhanced attributed to re- melting. Another FEM model was demonstrated by Vastola et al. [102] where phase transformation during the process of re-melting and solidification was investigated. The effect of process parameters including process velocity, beam size and laser power were discussed. They pointed out that in order to achieve a better quality of interlayer re- melting, high power and high process velocity should be adopted rather than using low power and lower process velocity. Liu et al. [103] presented a numerical model to analyse the melt pool formation and residual stress induced deformation after laser re-melting. More deformation was found while the total stress level was decreased and transforming from compressive to tensile in the unre-melted region after the re-melting process. Gusarov et al. [104] proposed a FEM model to investigate the evolution of residual stress during laser re-melting and additive manufacturing process. They concluded that the residual stress was more dependent on shape rather than size of the re-melted region, where tensile stress was found twice as high in longitudinal than in traversal direction.

Chapter 2 Literature review of laser additive manufacturing

2.8 Melt pool behaviour

During laser metal deposition, a melt pool will be developed when the laser beam hits the substrate surface within a short interaction time. Metal material in the form of powder or wire is fed into the pool and then melts after absorbing the heat. Because of the small interaction area laser beam is capable of producing, cooling rate of the process is high and thus the molten material can solidify quickly. The dimensions of deposited bead including width, height and contact angle are highly dependent on the melt pool shape. Therefore, getting a full understanding of the melt pool behaviour is essential and each factor affecting the pool penetration should be investigated.

When the arc is used as the heat source, Marangoni stress, electromagnetic (Lorentz) and aerodynamic drag forces are normally included when talking about the flow mechanisms happening in the pool [105]. For the laser-based process, Marangoni stress induced convection flow is generally taken into consideration [106, 107]. However, gravitational force is assumed to have a negligible effect during the melt pool formation since most of the previous study is focused on the process performed in the flat position. However, gravity may become a significant impact during the process conducted in horizontal, vertical and upside-down positions.

2.8.1 Marangoni flow

Surface tension arises as a result of different attractive forces between molecules on the fluid surface. In laser processing, Marangoni induced flow is driven by surface tension difference caused by the temperature difference on the melt pool surface. Fluid will always flow from low surface tension area to high surface tension area and hence a continuous flow will be formed inside the melt pool before solidification.

Su et al. [108] and Cho et al. [109] pointed out that convection flow driven by surface tension in melt pool played the most important role determining the bead dimensions including width and penetration depth. Koo et al. [106] concluded that shallower penetration in melt pools exhibited a radially outward surface flow pattern while deeper penetration in melt pools exhibited a radially inward surface flow pattern. Surface tension was highly dependent on the temperature and different surface-active element content especially sulphur. One study by Heiple and Roper [110] mentioned that a large surface-

Chapter 2 Literature review of laser additive manufacturing tension gradient would exist across the surface due to the reason that a large temperature gradient existed between the centre and the edges of the melt pool. They also claimed that when the content of surface active element like sulphur was high enough, surface tension temperature coefficient would be positive, and an inward convection flow would be produced which would contribute to a narrow and deep melt pool. Conversely, materials with low concentrations of surface active elements would produce an outward surface fluid flow, as a result of which a wide and shallow melt pool would be generated. In this case, the surface tension temperature gradient appeared to be negative. Similarly, research from Shirali & Mills [111] showed that an increase in sulphur content would produce an increase in the depth/width ratio. For some other elements like oxygen, Robinson & Gould [112] proposed that although oxygen was almost as surface active as sulphur, it did not always have as great impact on melt pool formation as sulphur because the actual soluble content of oxygen normally was quite low.

Keene [113] pointed out that in reality the surface tension would decrease when approaching the critical temperature, indicating that the surface tension temperature coefficient would go through a maximum at the critical temperature and a reverse direction of convection could be produced. This view was also supported by Lee et al. [107] who used a model which had a critical temperature where the surface tension gradient would change from positive to negative. Mills et al. [114] applied an equation (2.6) shown below to calculate the values of surface tension gradient which had a transition temperature.

훥퐻0 −( ) 0 0 푅푇 γ = γ푚 − 퐴(푇 − 푇 ) − 푅푇Г푠푙푛 [푙 + 푎푖푘 푒푥푝 ] (2.6)

0 0 where, γ푚 is the surface tension of pure metal at reference melting temperature 푇 , 퐴 is a coefficient for the variation of surface tension at temperature 푇 above the liquids, 푅 is 0 gas constant, Г푠 is surface excess at saturation, 푘 is entropy segregation constant, 훥퐻 is enthalpy of segregation, 푎푖 is the activity of species 𝑖 in solution. From the equation, the transition temperature of surface tension gradient for each material can be calculated.

Burgardt and Heiple [115] proposed that the primary mechanism, where variations of process parameter could affect the deposition quality, was achieved by changing the temperature gradient on the melt pool surface. In their research, all the bead geometry changes caused by varying the process parameters including power density and process

Chapter 2 Literature review of laser additive manufacturing velocity were found to be consistent with the surface tension driven fluid flow model. Their result was also complemented by the work done by Mills and Keene [105]. From their analysis, the change in thermo-capillary forces was attributed to the change in temperature gradient which was directly affected by the heat input and process velocity.

Collectively, these studies outline a critical role in investigating the impact of convection flow driven by the surface tension difference, which will determine the melt pool geometry. Therefore, if desired deposition quality needs to be achieved in various operating positions, Marangoni flow should be taken into consideration.

2.8.2 Gravitational force

Gravity has not been regarded as having a significant role during the melt pool formation due to the fact that most of the reported researches were conducted on flat position. Only a few efforts were made to achieve good deposition quality on vertical position [116, 117].

Figure 2.23 Cross-section of bead when applying different process velocity (a) 200 mm/min, (b) 400 mm/min, (c) 700 mm/min, (d) 1000 mm/min [118]

Paul et al. [118] pointed out that laser rapid manufacturing process on the vertical surface substrate was important for many engineering applications, such as surface cladding of turbine blade shroud and interlock, offshore drilling heads, cylinder bodies, sleeve and

Chapter 2 Literature review of laser additive manufacturing mould side walls etc. From their modelling and experimental work, they concluded that more deposit would shift downwards when lower process velocity was applied as shown in Figure 2.23. The reason was that shorter interaction time would result in a shorter period for the deposition to remain in the liquid state where higher effective viscosity was able to counter the flow of molten material under gravity.

During laser welding process, Guo et al. [119] discovered that the sagging and undercut defects could be avoided by performing welding in horizontal position where the pressure in the centre of melt pool was relatively low when compared with the pressure in flat welding position. The low pressure would make it possible for the surface tension to balance the hydrostatic pressure brought by the gravity force.

Due to the existence of temperature gradient in the melt pool, density difference would occur resulting in buoyancy force. The direction of this kind of convection flow induced by thermal expansion was vertically upward which would promote a wider and shallower melt pool which is illustrated in Figure 2.24 [120]. Tsai and Kou [121] developed a model to discuss the effect of density change on melt pool formation. They pointed out that free surface development was influenced not only by the superheating induced density variation but also related to density decrease caused by melting. Thus far, most of the research in investigating the impact of buoyancy force during the process were performed based on the condition when the melt pool was on flat position. All the results showed that the buoyancy force played an insignificant role in determining the melt pool behaviour. Koo [106] mentioned that buoyancy was treated with no great circumspection in his research due to the reason that gravitation was assumed to be a negligible effect on laser welding of thin metal sheets. Results of a mathematical modelling from Mills and Keene [105] showed that for TIG welding the influential factor for four convective forces sequence from the most to least: Marangoni, electromagnetic, aerodynamic and buoyancy, among which the buoyancy force played the least important role. Similarly, another research from Lee et al. [107] decided not to include melt pool convection induced by buoyancy force because it was typically negligible in comparison with the flow created by surface tension gradient.

Chapter 2 Literature review of laser additive manufacturing

Figure 2.24 Buoyancy force and convection flow driven by buoyancy force [120] However, for the situations in non-flat positions, only a few studies have been done on the impact of buoyancy force in melt pool behaviour. Traidia et al. [122] argued that although buoyancy forces were not of high importance in flat-position GTAW (where the fluid flow was mainly governed by the Marangoni effect), it played an important role during horizontal gas tungsten arc welding (GTAW) process, since it would act in a different direction than the Marangoni stress. They concluded that the buoyancy effect could sometimes be neglected in flat-position welding simulations to speed up the convergence. However, this assumption was no longer true for non-flat position welding system. From their modelling work shown in Figure 2.25, they also pointed out that Marangoni shear stress played a beneficial role in limiting the asymmetry of horizontal GTA welds while the buoyancy forces were the leading cause for the appearance of asymmetry inside melt pool. As shown in Figure 2.25 (a), the shape of the melt pool appeared to be symmetric. However, the convection flow driven by buoyancy force seemed to create an uprising flow making the upper side of the melt pool much bigger as illustrated in Figure 2.25 (b). Figure 2.25 (c) presents the combination of two convection flows which resulted in a melt pool with upper area larger than the bottom area.

Given all that has been mentioned so far, the effect of buoyancy force has not been deeply investigated since most of the research was done on flat position where the influential factor is trivial. However, it seems that buoyancy force also plays a pivotal role in the non-flat position process which will change the melt pool behaviour making it asymmetric. Considering that there is no specific study which has been done on the effect of buoyancy force during laser metal deposition process in different positions, the role of buoyancy force should be further investigated.

Chapter 2 Literature review of laser additive manufacturing

a b c

Figure 2.25 (a) Melt pool shape formed by Marangoni stress only, (b) Melt pool shape formed by gravity force only, (c) Melt shape formed by the combination of those two forces together [122]

In recent years, additive manufacturing has been increasingly mentioned for space-based applications. National Research Council (NRC) [123] evaluated the prospects of in-space additive manufacturing examining both the benefits and challenges. The potential benefits included the possibility of creating replacement components during long-term space flight, recycling trash materials collected from space, creating structures difficult to produce on or transport from Earth, creating sensors & sensor systems & satellites, fully printing spacecraft and using the resources on planetary surfaces. Apart from all the benefits, technical challenges also existed when compared with AM process on the earth, which included consideration of space environment (zero-gravity and vacuum), difference in thermal environment (lack of convection), infrastructure and platform building up and power supply. In their report they mentioned that with the absence of gravity, surface tension forces would become more important determining the system behaviour. They also pointed out that the process which relied on the controlling of fluid or flow conditions required further research. Another report from NASA [124] demonstrated the integration of an FDM 3D printer into the ISS Microgravity Science Glovebox (MSG) and produced the first 3D printed sample in space. Comparison of mechanical and microstructure properties between samples obtained from ground and space were performed. A technology assessment of laser-assisted materials processing in space from Karthik and Karen [125] suggested that wire could be used as feeding material to eliminate the handling of powder feedstock in a zero or low gravitational atmosphere. However, little is known on the possible result of metal-based additive manufacturing process in space with reduced gravity or zero gravity environment. Therefore,

Chapter 2 Literature review of laser additive manufacturing investigations should be performed to help adapt the knowledge of additive manufacturing from the earth to the space environment.

2.8.3 Melting and solidification

In most of the cases, it is rather difficult to measure the temperature gradient and solidification rate inside the melt pool during the process due to the small size of melt pool and rapidly changing flowing condition of the molten fluid. Therefore, numerical models are typically applied to help get a better understanding of the solidification process inside the melt pool. Based on the mathematical correlation between cooling rate and microstructure formation, the final microstructure of deposition like dendrite spacing can be used to validate the model. Kumar and Roy [126] developed a theoretical process map for predicting the effect of operating parameters on deposition height, width and dilution during laser cladding process. With the assistance of FEM modelling, they were able to explain the thermal distribution inside the melt pool. In general, the formation of microstructure and morphology were determined by temperature gradient G and solidification velocity Vs. Cooling rate 푇̇ = 퐺푉푠 would determine the length of dendrite along the solidification front, where higher cooling rate would result in finer microstructure grains. On the other hand, microstructure formation was governed by ratio

G/Vs, where a rise of the ratio would contribute to a transformation of microstructure from equiaxed dendrites to columnar dendrites. They also discovered that since maximum solidification velocity lied on the highest point of solid-liquid line while minimum on the bottom of solidification front, cooling rate 푇̇ appeared to be highest on top of the melt pool while highest ratio G/Vs lied on the bottom of melt pool. Neela and De [127] presented a three dimensional model for the analysis of heat transfer during laser metal deposition process. They pointed out that cooling rate reached the highest on the solidification interface and would decrease with the distance getting far from the melt pool centre. They also mentioned that laser power should be lowered for upper layers in order to keep a constant melt pool size due to the heat accumulation effect. Lee et al. [128] built a numerical model to first obtain thermal status in different regions of melt pool, followed by prediction of grain growth like primary dendrite arm spacing (PDAS). Surface tension gradient dependent on temperature variation was included in his model. Calculated temperature gradient and solidification rate were used to analyse microstructure formation by correlating the obtained results with theoretical models.

Chapter 2 Literature review of laser additive manufacturing

Solidification rate was found to be maximal along the centreline while decreasing from centre to solid-fluid interface and top to bottom inside the melt pool. Solidification rate difference was also observed in different depths of the melt pool, where at the tailing end there was a big decrease from top to bottom surface. This was attributed to dramatic drop of the angle between solidification boundary normal and laser beam moving direction. On the contrary, temperature gradient was found to be minimal in the melt pool centre and would increase along the depth and width of the melt pool. The cooling rate was then calculated based on the distribution of temperature gradient and solidification rate. They reported that the highest cooling rate was found in the deepest region of melt pool due to less effect from convection flow. Convection flow induced by Marangoni stress also changed the distribution of cooling rate and temperature gradient in fluid mixing region. Primary dendrite arm spacing (PDAS) was found to increase with the decrease of cooling rate inside the melt pool.

Figure 2.26 Impact of temperature gradient and solidification rate on microstructure formation [129] Kou [129] mentioned that microstructure formation was determined by solidification rate and temperature gradient, where cooling rate determined structure grain size and the ratio between temperature gradient (G) and growth rate (R) determined morphology of the microstructure. Cooling rate is defined as the product of G and R while ratio G/R is related to a criterion given by [129]

퐺 ∆푇 ≥ (2.7) 푅 퐷퐿

Chapter 2 Literature review of laser additive manufacturing where ∆푇 is the difference between solidus and liquidus temperature of alloy material,

퐷퐿 is diffusion coefficient. This criterion should be met in order to ensure the planar solidification interface to stay in a steady condition. For a certain material with fixed freezing range and diffusion coefficient, higher growth rate or smaller temperature gradient would break this balance and hence resulted in grain growth shift from planar to cellular. It is shown in Figure 2.26 that with the decrease of ratio R/G, microstructure would change from planar to cellular, columnar dendrite and equiaxed dendrite. Planar structure occurred when growth rate was very small while equiaxed dendrite happened when growth rate was extremely high. Kou [129] also pointed out that with smaller cooling rate, more time would be given for the large dendrite arms to grow at the cost of smaller dendrites during solidification process. Therefore, higher cooling rate would result in lower cell spacing and hence finer microstructure. David and Vitek [120] confirmed that growth rate was highly related to the process velocity. Melt pool appeared to be elliptical when heat source was applied with small velocity shown in Figure 2.27(a). Minimum temperature gradient would lie on the melt pool centreline where the growth rate reached the maximum. Higher growth rate would contribute to more fusion heat to be released during the solidification process. Therefore, a critical growth rate was put forward where the latent heat could just be dissipated through the temperature gradient along the centreline. For higher heat source velocity or in other words higher growth rate, melt pool shape would get elongated and became teardrop shaped. Maximal solidification rate would be lower than heat source speed because the minimal angle between maximal thermal gradient and process velocity could not reach 0 at tailing region of the melt pool shown in Figure 2.27(b).

Figure 2.27 Growth rate and temperature gradient on solidification boundary with different melt pool shape [120] 70

Chapter 2 Literature review of laser additive manufacturing

They also mentioned that the melt pool would be further elongated with more solute trapped in the tailing edge which could result in lower solidification temperature. Apart from considering the grain growth along temperature gradient direction, impact of crystallography should also be included during the analysis of microstructure formation. ‘Easy growth’ directions would exist during the solidification process, and the dendrite growth rate was related to the growth rate of solid-liquid interface. An FEM model was coupled with cellular automation method by Yin and Felicelli [130] to understand the relationship between cooling rate and microstructure formation during LENS process. Primary dendrite arm spacing (PDAS) and secondary dendrite arm spacing (SDAS) were calculated based on the thermal distribution history obtained from FEM. Both PDAS and SDAS were found to decrease with increase of cooling rate. With higher laser process velocity, PDAS and SDAS would become smaller. Due to the high cooling rate, secondary dendrite might not be observed when extremely high process velocity was applied which would result in microstructure morphology transformation from dendrite to cellular. They also discussed the effect of layer thickness, substrate size and different locations inside melt pool on dendrite morphology.

The microstructure difference brought by all the factors mentioned above can be attributed to different cooling conditions this local region has experienced. A high-speed visible imaging system was adopted by Griffith et al. [131] to investigate the thermal behaviour during direct deposition of 316 stainless material. Temperature distribution along the laser scanning direction was measured, and the maximal temperature was obtained in the centre of melt pool region where temperature gradient was zero. Maximum temperature gradient 160 K/mm was located approximately 1mm away from the melt pool centre. A sharp decrease in temperature gradient was observed with the increase of measurement distance from heat source until reaching the solidification boundary where thermal gradient stayed around 30K/mm. Kobryn and Semiatin [132] presented a solidification map for Ti-6Al-4V regarding the effect of temperature gradient and growth rate. Different microstructure morphology was observed when different laser system was applied. Grain width was found to decrease with the increase of laser process velocity while the influence of laser power on grain width was much less significant when compared with laser process velocity. During single-crystal laser deposition process, Gaumann et al. [133] developed a microstructure map correlated with various process parameters to ensure epitaxy and columnar grain growth avoiding the appearance of the

Chapter 2 Literature review of laser additive manufacturing equiaxed microstructure. Specific critical ratio G3.4/V was decided for the adjustment of process parameters correspondent to different thermal conditions in order to achieve control of the desired microstructure type.

2.8.4 Overhang

One challenge in layer-based additive manufacture process is to produce overhang structures as illustrated in Figure 2.28(a), where some part of the final product is without supporting material. Different methods are applied for depositing the overhang material between two adjacent layers shown in Figure 2.28(b) during different manufacturing processes. The maximum inclination angle of the geometry is relatively limited for laser metal deposition in comparison with powder bed-based laser rapid manufacturing process. This is due to the lack of supporting material which will bring difficulties to produce complex geometries with overhang and big-inclination angle features.

Figure 2.28 Illustration of overhang, L is layer thickness, St is layer top surface and Sb is layer bottom surface [134] Zhang et al. [134] investigated the relationship between layer thickness, maximum overhang distance and inclination angle. They addressed that the allowable overhang distance for every layer was dependent on the material properties, system characteristics and working conditions. They also mentioned that for a certain material, constructible overhang geometry was determined by layer thickness due to the fact that more deposited material would mean that more molten mass needed to be self-supported by surface tension. Several methods are commonly used to overcome these problems in order to achieve a decent quality of depositions on non-flat surfaces. The first way is to create some small increments of overhang region between each layer to form a staircase-like geometry, which is shown in Figure 2.29(a). In this case, there is no need for laser head

Chapter 2 Literature review of laser additive manufacturing and substrate to rotate while the overhang performance is highly related to material properties and process parameters. The second method illustrated in Figure 2.29(b) is to tilt the substrate or heat input direction in between each layer to keep the sample growing direction and depositing surface tangent. Melt pool is formed on the previously deposited layer and hence staircase phenomenon will not occur. However, deposition collapsing still happens when the gravity of the molten material exceeds the surface tension in the gravitational direction. In addition, this method requires more flexible working conditions like robot arms and rotating table. The third method is to create some supporting structures and remove them after the process [135].

Figure 2.29 Two different methods to produce overhang structures[136] A “vase” shaped geometry with 80° maximum inclination was successfully deposited by Shi et al. [136] by utilising a hollow laser beam which was mounted on a 6 axis robot. Deposition head was kept tangential with the growth direction and hence staircase effect was eliminated. Weiss et al. [135] mentioned that copper could be used as the supporting material for stainless steel due to the fact that these two materials would experience a small difference in melting temperature but a big difference in thermal conductivities. Newly input energy could only re-melt the early deposited steel material but not copper which would dissipate heat faster. After the process, nitric acid was applied to remove the supporting material. Overlap ratio was found by Shang et al. [137] to be the main factor limiting slope angle of the deposited geometry with overhanging features. Zhu et al. [138] discussed the influence of substrate-inclined angle on laser deposition quality by utilising an inside-beam powder feeding system. From their experiments, the powder feeding nozzle was kept constantly perpendicular to the substrate while inclination angle of the stage was varied between each layer. Deposition on non-horizontal position was achieved by Zhu et al. [138] although convergence of the powder became worse with an increased inclined angle. By tilting the arc and wire feed angle relative to the substrate, walls with

Chapter 2 Literature review of laser additive manufacturing different inclination angles were successfully deposited by Kazanas [139] et al. A 6-axis robot arm coupled with two additional tilt and rotatory positioning system was developed by Ding et al. [140] to realise deposition of a revolved geometry. Nassar and Reutzel [141] mentioned that overhang structures were more likely to be produced if a pulsed or power- modulated beam was applied. Continuous wave laser instead would result in a large melt pool where surface tension and rapidly solidified region could not hold gravity of the melt pool with a big volume.

Contact angle hysteresis was introduced by Wu et al. [142] shown in Figure 2.30 to explain the dripping phenomenon during a pulsed arc additive manufacturing process. The difference of the contact angle would increase with a larger droplet and the droplet would start to flow down when the gravity force exceeded the hysteresis force. In order to keep the balance between gravity and surface tension, they suggested that a smaller heat input could attribute to a reduced melt pool and droplet size which hence could reduce the impact of gravity force.

Figure 2.30 Contact angle of a water droplet adhering on a glass window [142] A hybrid system was developed by Ruan et al. [143] to combine the laser additive manufacturing and material removal process. Paths for deposition and milling were defined automatically and complex geometry with 45º overhang features was produced. Hensinger et al. [144] discussed that gravity was able to drive the melt pool towards the supporting substrate if laser remained vertically aligned with the working surface. Several different methods were presented and compared by Prabhu et al. [145] to investigate the effect of process parameters on overhang structure qualities. Extension mode with 5 axes controlling (by tilting the substrate after each layer) was found to be the most qualified method attributing to the maximum support for melt pool from the previous layer.

Chapter 2 Literature review of laser additive manufacturing

2.9 Modelling

In general, dynamic flow in the melt pool is difficult to be clearly captured and analysed with experimental results. Modelling provides an effective way to help fully understand the process and it can also be used to improve the production sequences and to minimize production problems with less investment in experimental trials.

There are mainly two categories of modelling which are widely used to simulate the complicated process [146], namely: (1) Thermal-mechanical modelling, (2) Multi- physical modelling. The first method is the fundamental modelling for various laser applications since all the laser-based processes undergo a process of heat transfer. After achieving the thermal history of the process, residual stress can be predicted mostly done by using FEM techniques. The second method concentrates on the simulation of fluid flow inside the melt pool. Both FEM (Ansys, COMSOL, Abaqus) and FVM (Fluent & Flow 3D) methods are capable of simulating the thermal distribution of the process. Considering that most of the current work concentrates on the investigation of melt pool behaviour which is highly related to the second area, the literature review below will focus on the multi-physical modelling of melting and solidification.

2.9.1 Thermal-mechanical modelling

2.9.1.1 Thermal residual stress and strain Temperature field and deposition profiles were estimated by Manvatkar [147] using a heat transfer model. They found that the cooling rate would decrease while maximum temperature would increase with more layers being deposited. Distribution of residual stresses obtained from a 3D model was compared by Labudovic and Kovacevic [148] with x-ray diffraction results. They concluded that residual stresses would increase during multiple layer process in comparison with a single layer deposition. A 3D heat transfer model was built by Chirag and Patel [149] to investigate the temperature history, total deformation, von-misses stress and shear stress during LMD process. The model developed in ANSYS software utilized element birth and death technique to achieve addition of the new material. Steady-state calculation based on COMSOL Multiphysics software was used by Peyer et al. [150] to simulate the wall geometry during LMD, after which a transient thermal model was created to predict the temperature history of the whole process. Wang et al. [56] developed a model using SYSWELD to study

Chapter 2 Literature review of laser additive manufacturing temperature history and residual stress state during LMD process. Experimental results from neutron diffraction mapping were compared with the simulated stress distribution. Process velocity was found to have little effect on residual stress along the building up direction while higher laser power could contribute to an increased residual stress. For longitudinal residual stress, higher process velocity could result in tensile residual stress while compressive residual was obtained when low process velocity was applied. An FEM model was built by Farahmand and Kovacevic [151] to calculate temperature gradient, residual stress history and development of the melt pool during multiple layer laser cladding using a high power diode laser. They reported that the level of residual stress differed with the cladding region and the high-stress region appeared in the area where the last tracks were deposited. This might due to be the high cooling rate and stress relaxation from previous tracks. Stresses were relieved during the re-melting process which lowered the residual stress of previous layers. They also observed that smaller process velocity would result in higher residual stresses both in longitudinal and transverse directions.

Figure 2.31 Stress components of a single track laser deposition (a) x-direction, (b) y- direction, (c) z-direction, (d) von Mises equivalent stress [151]

Michaleris [152] presented a model combining quiet and inactive method to realise mass addition in an FEM model. Temperature history was analysed and results could be

Chapter 2 Literature review of laser additive manufacturing obtained more efficiently by utilizing this new hybrid method. Zhu et al. [153] produced a sample which contained curving geometries with various radius. They built a transient temperature model using ANSYS to investigate the temperature history during this multiple layer deposition especially between different deposition layers and around the curvature region. Higher temperature was observed with more layers building up and higher temperature with a larger melt pool was also found around curvatures with a smaller radius value. They suggested that decreased laser power should be applied to the mentioned situations to achieve a stable melt pool and deposition with even layer height. For multi-material deposition, Dai and Shaw [154] introduced a three-dimensional FEM model to gain a better understanding of distortion, temperature field, transient and residual stress during the process. They concluded that deposition sequence, deposition pattern and process velocity had a considerable influence on stress state as well as temperature history. Deposition sequence would determine which material to deposit first among multiple materials with different thermal conductivity and thermal expansion coefficient. Deposition pattern referred to the strategy of changing process direction between each depositing pass or feature. Deposition of a straight wall and a cylindrical wall was analysed numerically and experimentally by Zekovic et al. [155]. A heat transient model was developed by Zhang et al. [156] to help understand the role of preheating on thermal history and residual stress distribution. The resultant stress value obtained from numerical model was found to decrease with increase of the preheating temperature, hence distortion and cracking defect could be prevented. Alimardani et al. [87] presented a three-dimensional model to investigate the influence of preheating and clamping on stress state. Thermal distribution was first calculated using COMSOL followed by a combined code which was created in MATLAB to achieve prediction of the deposition geometry. They concluded from their work that preheating reduced the time required for melt pool to reach steady state. Residual stress could be lessened by both preheating and applying a clamping to the system.

A one-dimensional model was adopted by Bruckner et al. [157] to understand the residual stress distribution during cladding process. Thermal expansion and contraction were considered as a result of thermal cycle, which in turn contributed to the density difference in different sample area. Thermal strain was related to the local material density where stress could be calculated. Further on, a three-dimensional FEM model was built by them to investigate the influence of process parameters on residual stress. A three-dimensional

Chapter 2 Literature review of laser additive manufacturing

FEM model was built by Zhao et al. [158] to investigate the temperature and stress distribution during a multiple layer laser cladding process. Temperature field was first simulated based on which the stress field was calculated. Result of residual stress distribution indicated that the longitudinal stress was the highest, followed by transverse stress and stress along the thickness direction. Cracks were found more likely to appear in the intersection region between layers.

2.9.1.2 Metallurgical model for phase transformation Bailey et al. [159] pointed out that temperature field, microstructure and stress condition were closely related during laser hardening of AISI 4140 steel. FVM method was firstly applied to solve a thermal/strain model, after which temperature and microstructure histories were imported in ABAQUS to analysis the hardness field and stress state. Neil et al. [160] presented a numerical model shown in Figure 2.32 which covered processes of heating, cooling and tempering to obtain a deep understanding of solid phase transformation, hardness and residual stresses distribution during laser metal deposition process for AISI H13 tool steel. Phase transformation between martensite, ε-carbide and ferrite, cementite and ferrite were taken into account according to how temperature history of the process went through. Residual stress model was built based on the assumption that elastic and plastic strains were dependent on temperature but independent of material phase. Melting and solidification were considered using a MODEL CHANGE command in ABAQUS. Hardness was also calculated based on the phase fraction.

Figure 2.32 Phase fraction of martensite during laser metal deposition [160]

Chapter 2 Literature review of laser additive manufacturing

A thermos-kinetic laser powder deposition model was developed by Costa et al. [161] to investigate the role of substrate size and idle time in between depositing layers on the distribution of the microstructure and hardness for AISI 420 steel. Heat transfer model was coupled with transformation kinetics and temperature dependent material properties. Fresh martensite and austenite were found on the upper section with higher hardness while tempered martensite appeared in the lower region of the deposition with lower hardness. They also found that smaller substrate size or shorter time interval in between layers would contribute to a bigger heatsink and reduced amount of tempered martensite. The solidification process and residual stress were investigated by Ghosh and Choi [162] utilizing a finite element model in three steps. Temperature and phase composition history was first analysed using ABAQUS, followed by calculation of residual stresses. Development of microstructure was lastly modelled by applying a subroutine to predict TTT (time-temperature transformation) curved for low alloy steels. The microstructure evolution including austenite decomposition and formation rate of different phases during the cooling process could be determined by tracking the thermal history of the location of interest. From the modelling results they concluded that stress evolution and residual stress were greatly influenced by phase transformation. Distortion and cracking were observed when residual stress exceeded the yield stress.

Santhanakrishnan et al. [163] built a model to predict solidification, temperature history and thermal cycle when laser cladding H13 steel on AISI 4140 steel. Phase change kinetics were included and a coupled FE-TK method was applied to understand microstructure and hardness distribution. They found that different microstructure appeared in different regions of the deposited sample: (1) fine dendrite near the deposition surface, (2) mixed dendrite and cellular structure in between two layers, (3) cellular and coarse cellular in dilution area, (4) coarse cellular in HAZ. Hardness distribution in different sections was discussed and compared which could also be attributed to the regional thermal cycle. Kelly and Kampe [164] adopted a model to investigate the microstructure evolution when laser depositing Ti-6Al-4V material. They pointed out that the deposition microstructure was determined by thermal cycle during the building up which included peak temperature, time duration at peak temperature and cooling rate. A microstructural development map was created based on the thermal analysis. α + β-phase field was found for the starting few layers and (α ↔ β) α + β-phase appeared on the upper layers as a result of higher peak temperature and lower cooling rate. Ti-6Al-4V was also

Chapter 2 Literature review of laser additive manufacturing used as depositing material by Bontha et al. [165]. They developed a model to obtain thermal map during the process of building a thin wall, followed by prediction of solidification and microstructure development. Cooling rate and thermal gradient were thought to be the dominant factor for the microstructure formation. They concluded from their work that process parameters including laser power and process velocity would have a big impact on cooling rate and thermal gradients. Microstructure tended to change from columnar to mixed and equiaxed with an increased laser power density. Microstructure of laser deposited Ni-base superalloy in overlap area was investigated by Pirch et al. [166]. Since growth of dendrite would follow the temperature gradient direction, [010] growth was found near the border area between adjacent layers. [100] growth above the substrate surface and [001] growth closer to the substrate were observed respectively which could be verified with the simulated cooling rate and temperature gradient distribution across the bead area. A theoretical kinetic analysis on growth and nucleation of TiC particles was coupled with FEM thermal analysis by Lei et al. [167] in order to understand the phase transformation and microstructure development during a laser cladding TiC/NiCrBSiC composite process. During the process of rapid manufacturing Titanium samples, Crespo [168] pointed out that cooling rate was the main factor determining transformation from β phase to two different phases. Heat transfer model using FEM was firstly built, followed by microstructure analysis including diffusional transformation and martensitic transformation depending on the cooling rate during the manufacturing process. Other factors including re-melting, preheating, process velocity and time interval between adjacent layers were found to have influence on the cooling rate and hence would affect the phase transformation. A microstructure evolution map correspondent to different process parameters was created with the results obtained from the model coupling heat transfer and phase transformation.

2.9.2 Multi-physical modelling

Thermal modelling of melting and solidification process includes several modules. Free surface between air and substrate is firstly located either by using boundary condition or VOF method, where laser and added material source can be added. Laser beam applied on the substrate surface is treated as heat flux or volumetric heat source according to the method used. Material will be melted when reaching the melting point and solidify after cooling down. Temperature-dependent material properties including thermal

Chapter 2 Literature review of laser additive manufacturing conductivity, density and specific heat capacity are normally considered which will introduce significant non-linearity into the solution. Due to the fact that the melt pool development is highly related to the convection flow induced by surface tension difference, so temperature dependent surface tension is an important property needed to be considered. Boundary conditions including different heat transfer methods are also included in the model.

Most of the past research in this area focused on the melt pool formation, history and distribution of the temperature and the effect of process parameters on melt pool size during the process. Not too much work has been found on adding material into the pool to form a deposition, especially in various operating directions.

In general, the first step of modelling is to locate the free surface which is the interface between air and substrate where surface tension forces, heat source and mass addition are applied. Different method of free surface tracking brings different calculation process which in turn will result in different modelling accuracy and modelling result. There are four methods that are frequently used in modelling of laser metal deposition process: volume of fluid (VOF) method, dynamic mesh method, level set (LS) method and Arbitrary Lagrangian Eulerian (ALE) method. They all have their advantages and disadvantages. Literature reviews about these four different methods are presented below.

2.9.2.1 VOF method The volume of fluid (VOF) method was firstly reported by Hirt and Nichols [169]. It can simulate two or more immiscible fluids by tracking volume fraction of different phases in each cell throughout the domain. A scalar fraction function F is introduced which is defined as the integral of a fluid's characteristic function in the control volume. The volume fraction of each fluid is tracked through every cell in the computational grid, where all fluids share a single set of momentum equations. When a cell contains no tracked fluid inside, F=0; when a cell is full of tracked fluid, F=1; and when a cell locates on the interface between the tracked fluid and another fluid, 0 < F < 1.

The VOF method is widely used in modelling of laser-based process, especially when investigating formation and temperature history of the melt pool. The VOF method was employed by Frank et al. [170] to understand the thermal behaviour of melt pool due to a moving heat source. The model was used both for the preheating process and the actual laser deposition process. The melt pool configuration was defined in terms of a volume

Chapter 2 Literature review of laser additive manufacturing of fluid function, F(x,y,t), which represented the volume of fluid per unit volume and would satisfy the conservation equation shown in equation (2.8):

휕퐹 + (푉 ∙ ∇)퐹 = 0 (2.8) 휕푡

If F =1, the cell would be full of substrate phase while the cell would be located in the air phase when F=0. For those cells located on the interface between air and material, they would have a fraction value 0

Chapter 2 Literature review of laser additive manufacturing

Figure 2.33 Temperature distribution and bead geometry [172] A two-dimensional model was developed by Choi et al. [174] to study the laser cladding process including the process of melting, solidification and evaporation. The VOF method was used to follow the melt pool surface where mass material was added. Surface roughness was found to be highly related to the laser power and powder feed rate. Too high or too low laser power could both contribute to a high surface roughness while surface roughness was observed to be more sensitive to powder feeding rate. Therefore, proper process parameters should be determined in order to obtain deposition with decent quality. Ibarra-Medina et al. [175] built a laser metal deposition model which took into account the interaction between powder and gas flow, powder and laser beam as well as the melt pool formation. Realistic deposition geometry could be modelled which was attributed to the coupling of several models with few assumptions to be made. The VOF method was utilized in their work and it was mentioned that this method could be applied for modelling deposition with complex geometries such as clads with high wetting angle. The VOF method was implemented in a two-dimensional model created by Liou et al. [176] in order to understand the phenomenon during laser deposition process. The effect of powder ejection was included and some lack of fusion phenomena was noticed as a result of low total heat input. From the same research group, Fan et al. [177] presented a microscopic model utilizing the result of temperature history and solute concentration fields obtained from the macrostructure model. Grain formation inside the melt pool and along the substrate/liquid interface were modelled and compared.

Arrizubieta et al. [178] questioned the importance of melt pool dynamics during laser metal deposition process. Firstly, they confirmed both from their experimental and modelling results that convection flow brought by the temperature difference would have a big influence on the melt pool formation. However, when additional filler material was

Chapter 2 Literature review of laser additive manufacturing taken into account, not too much difference on the final bead geometry could be observed among cases with or without the consideration of convection flow. They explained that the temperature difference between newly added material and molten material inside the pool would result in a reduced temperature gradient and hence the influence of convection flow caused by Marangoni stress was decreased. The VOF method in OpenFOAM was implemented by Gürtler et al. [179] to model powder bed based deposition process. Powders were initially patched as face centred spheres with evenly distributed spacing between each other. The free surface between gas and powder material was tracked using VOF. Defects like lack of fusion and porosities between layers were observed in their model and had a good agreement with the obtained experimental results.

2.9.2.2 Dynamic mesh method The dynamic mesh method is applied by researchers to simulate the clad formation by considering the boundary motion. Deposition growth is realized by controlling the amount of mesh deforming within each time step. Amara et al. [180] applied the dynamic mesh method involving the use of user-defined functions (UDF) in the calculation procedure which allowed to follow the variation of the cell volume and then to obtain the clad profiles as a function of the operation parameters. From their model they also discussed the influence of process velocity on the height of obtained clad which is shown in Figure 2.34. Heat conduction equation, initial & boundaries conditions, laser power absorption, Marangoni convection and temperature dependent physical properties were included in their model.

Figure 2.34 Bead deposited with various process velocities

Chapter 2 Literature review of laser additive manufacturing

From another research group, Kai et al. [181] developed and implemented a new dynamic moving mesh method to accurately capture the details of temperature evolution anywhere in the build platform in an efficient manner during the simulation of selective laser melting process. They concluded that this method had been shown to provide significant computational enhancements over other solution methodologies while enabling fine-scale solutions in the domain space.

Although the dynamic mesh method provides a possibility of modelling laser manufacturing process, not too much modelling work has been found based on this method. This may be due to the fact that the amount of deforming mesh by the time is predefined and hence it cannot be used to predict the deposition geometry when a group of new process parameters are to be tested.

2.9.2.3 Level set method The level set method is a satisfying way of viewing the well-known kinematic boundary condition of the interface. A distance function φ(x⃗⃗, t) is defined in the domain as follows [182] (2.9):

φ(x⃗⃗, t) = ±d (2.9) where d defines the distance between x⃗⃗ and the free surface. Only zero level set has physical meaning, i.e., describing the free surface position. In the gas area, φ(x, t) < 0 while φ(x, t) > 0 indicate those cells located in the substrate area. The interface can be identified when φ(x, t) = 0. The level set function evolves as [182]:

휕φ + 퐹 ∙ ∇φ=0 (2.10) 휕푡 where 퐹 is a speed function evaluated at each point in the domain. Because the continuity of the level set approach, the function can be updated after each time step in order to reobtain the zero value of the level set function tracking the free surface of the deposited layer inherently.

Kong and Kovacevic [51] applied this method to achieve tracking the free surface motion of the melt pool with feeding powder material and laser beam. From the results of the model, they concluded that an increase of laser power and powder feed rate or reduction of the process velocity could contribute to an increase of the clad height and would directly influence the re-melt depth between each layers. Their numerical results shown 85

Chapter 2 Literature review of laser additive manufacturing in Figure 2.35 were found to have a qualitative agreement with the experimental measurements. Another report from He and Mazumder [183] also applied level set method to track the evolution of the liquid/gas interface. They investigated the evolution of temperature and velocity fields, free surface development and energy distribution at liquid/gas interface during coaxial laser metal deposition. Some small extent of contraction could be observed after comparing the calculated zero level set function with experimental liquid/gas interface. A comprehensive 3D self-consistent transient model for coaxial laser metal deposition process was developed by Wen et al [184]. In their model, the effect of continual addition of mass and energy due to deposited powder was rigorously considered by incorporating new source terms. The melt pool and part geometry of multilayer deposition was successfully modelled, and the deposition profiles agreed well with experimental results.

Figure 2.35 Temperature distribution of third layer after 6s [182] A double layer laser metal deposition model was developed by He et al. [185] to understand the temperature history and composition profile of the process. Convection flow was considered to be the main factor for the evolution of solute concentrating distribution. They reported that the heat accumulation after the first layer would have a big impact on formation of the second layer, where peak temperature during the second layer would become higher. Interestingly, they found the maximum temperature located in the re-melting region of the previous layer rather than on the melt pool surface. The level set method was employed by Mirzade et al. [186] to simulate the process of melting and crystallization during laser cladding. Effect of process parameters on melt pool formation and temperature distribution were discussed. They observed that temperature change was monotonic with a high cooling rate when starting, while the temperature

Chapter 2 Literature review of laser additive manufacturing change became non-monotonic when the process got stable. They suggested that it could be attributed to the decrease of cooling rate and the impact of latent heat released during solidification. Qi et al. [187] developed a numerical model for investigation of heat transfer and fluid flow in laser cladding process. Level set method was used to track the free surface between liquid and gas phase. Interaction between the laser and powder was discussed where they mentioned that some portion of the powder was already melted before reaching the melt pool. Vaporization of the powder should be avoided to keep a good powder utilization and reduce the development of porosities caused by plasma. Laser metal deposition process for repair use was modelled by Han et al. [188]. A hole was filled with melted material by two layers to achieve the target. Mass addition and frees surface tracking were realized using level set method. Some unevenness was observed at starting of the deposition and they suggested that this problem could be fixed by optimizing the laser process velocity and deposition pattern. Liu and Qi [189] built a three-dimensional model to simulate the laser deposition of Single-Crystal Superalloy. Effects of process parameters on melt pool formation were investigated especially the re- melting mechanism. They pointed out that approximately 85% of the previous layer could be re-melted when depositing the next layer. Regions with grains growing in different directions could be re-melted and hence a uniform microstructure directional growth could be achieved.

In comparison with the VOF method, the level set method has the advantages of dealing with robust geometric information (normal and curvatures) and has the ability to handle topological changes (merging and pinching). While for VOF method the fraction of the fluid on the interface is not continuous which can lead to numerical diffusion making the modelling results inaccurate. However, VOF has the advantage over level set method of being mass-conservative.

Figure 2.36 Coupled Level set and VOF method [190]

Chapter 2 Literature review of laser additive manufacturing

Recently a coupled level set method and VOF method is proposed by some people to make the most use of the advantages of these two methods in order to compensate the existing disadvantages. User Defined Scale was illustrated using Fluent software by Zhen [190]. She concluded that by reconstructing between the fraction of the fluid and level set value within a one-time step, the model could reach both mass conservation and continuity. Not too much work has been conducted in modelling laser-based process using this combined method. Recently a laser deposition model was developed by Dubrov et al. [191] coupling VOF and Level set method to achieve a good modelling accuracy on both free surface tracking and mass continuity. In their model, the volume fraction of each phase was considered when solving transport equations while level set method was applied to handle thermos-capillary and capillary forces. Deposition formation together with temperature distribution and velocity was successfully simulated. The importance of convection flow in determining the melt pool was also discussed.

2.9.2.4 Arbitrary Lagrangian Eulerian (ALE) method The arbitrary Lagrangian-Eulerian (ALE) method was firstly introduced by Hirt et al. [192] to solve Navier-Stokes equations both Lagrangian and Eulerian. The Navier-Stokes equations govern the motion of fluids and can be seen as Newton's second law of motion for fluids. The Navier-Stokes equations can be expressed as:

휕풖 2 휌 ( + 풖 ∙ ∇풖) = −∇푝 + ∇ ∙ (휇(∇풖 + (∇풖)푇) − 휇(∇ ∙ 풖)퐈) + 푭 (2.11) 휕푡 3 where 풖 is the fluid velocity, 푝 is the fluid pressure, 휌 is the fluid density, and 휇 is the fluid dynamic viscosity. The term on the left side of the equation corresponds to the inertial forces and terms on the right side of the equation correspond to pressure forces, viscous forces and the external forces applied to the fluid.

Mesh nodes are able to move together with the fluid (Lagrangian), stay fixed (Eulerian) or move arbitrarily. Many research groups have adopted ALE method to help develop numerical models to get a deep understanding of laser process. A two-dimensional transient model was built by Morville et al. [89] to investigate the effect of process parameters on melt pool behaviour. ALE moving mesh method was used to take into consideration the deforming melt pool surface as a result of surface tension and mass addition. The processes of depositing single and multiple layers were simulated respectively and they concluded that surface finish could be improved by increasing the

Chapter 2 Literature review of laser additive manufacturing dilution percentage between layers and adding less material for each track. Medale et al. [193] adopted ALE method to account for the evolution and deformation of the free surface during the laser spot welding process. Energy efficiency defined as laser absorptivity coefficient was found to increase from 50% to 80% with keyhole depth getting bigger where more reflected laser beam was trapped inside the keyhole. The moving mesh method was proved to be capable of dealing with rapid changing melt pool free surface of the keyhole. Medale et al. [194] also applied this method to build welding models with various energy input levels. Weld pool situations without keyhole, with fixed and deformable keyhole profile were considered respectively correspondent to the applied laser energy density during the process.

Frank et al. [195] pointed out that the ALE method could be applied to track the interface between the solid wire and melt pool surface during a laser cladding process using wire as filler material. Three steps were included when coupling wire feeding with deformable mesh boundary: motion equations of structures related to interaction pressure needed to be firstly found, after which structure acceleration related to the boundary pressure was built, the third step was to couple the structure motion equation with pressure solution. They addressed that the key aspect in their model was to combine VOF method with motion solution of rigid body equations. ALE method was utilized by Khairallah et al. [196] to model selective laser melting process in a micrometre scale. Explicit hydrodynamics and thermal conduction model were coupled to simulate the phenomenon of randomly distributed powder material, surface tension (Marangoni effect not included), Gaussian laser beam in microscale and temperature dependent material properties. Melt pool formation related to the process parameters and effect of surface tension were also discussed. Kong et al. [197] compared two situations during modelling of TIG welding process by using FEM method shown in Figure 2.37. In one situation top surface was fixed while the other situation includes evolution of free surface utilizing ALE method taking into account the minimization of total energy, where effect of surface energy, gravitational forces and arc pressure on the free surface formation was included. A deeper but narrower melt pool was obtained when free surface was considered. The front of top surface was depressed by arc pressure which would bring about a bigger velocity field on the surface and some part of the fluid was driven towards the rear region of the melt pool.

Chapter 2 Literature review of laser additive manufacturing

Figure 2.37 Melt pool surface with and without free deformation [197] The ALE and Phase Field methods were compared by Bruyere et al. [198] when simulating the laser spot welding process. For the ALE method, the controlled deforming mesh was achieved by solving a steady hyperelastic model while moving the correspondent boundary nodes. Level set was chosen as Eulerian method where two phases including air and liquid metal were considered. After comparison, they concluded that both methods had their advantages and disadvantages: ALE method worked well before the interface became complex. Since the effect of gas phase was only regarded as pressure boundary condition which was applied on the surface, the simulation process was less time consuming but on the other hand the accuracy of the model was more or less compromised. Due to the existence of gas phase, thermal effect including conduction and convection between gas and metal material were considered, making the model more realistic. Keyhole formation was found out to be highly influenced by the gas flow when using Level-Set method in comparison with the ALE method. However, phase field method was generally more time consuming as a result of more complicated coupling process between models. Dal and Fabbro [199] mentioned it in their review work on welding simulation that high distortion of mesh, which was a quite crucial factor in FEM modelling, could be avoided by applying the ALE method with self-improving mesh quality within each time step. However, they also mentioned that the drawback of this method also existed which was the limitation in modelling situations with closed frontiers like porosities.

2.10 Summary and discussion

This chapter presented some major aspects of the laser additive manufacturing. The main focus was placed on discussing how the various process factors would influence the final deposition result. Mechanical and thermal behaviour of the product could be related to

Chapter 2 Literature review of laser additive manufacturing the applied process parameters. Different numerical methods were introduced and compared.

Although a considerable amount of work can be found on investigating the laser metal deposition process, not too many studies have focused on understanding the possible effects of gravity because most of the previous work is conducted in the flat position. For the modelling part, most of the previous work concentrates on the analysis of temperature distribution and melt pool formation no matter if the free surface is considered or not. Only a few research include material addition into the dynamic flowing melt pool, among which simple geometries like single depositing line are typically chosen as the research target.

Therefore, the main knowledge gaps concluded from the literature review can be listed as: 1. A full understanding of the gravity effect during laser additive manufacturing process is missing. 2. No previous work has been found on developing a model which can simulate the deposition process with all complex features and also takes gravity effect into account when performed in various operating directions.

Chapter 3 Formulation of a CFD model for laser additive manufacturing process

3.1 Introduction

This chapter first gave a brief overview of the current status of modelling the laser metal deposition process, followed by some discussions about the gravity effect on the melt pool development. Next, the formulation of a three-dimensional transient numerical model was introduced. The major deposition mechanism and the driving force of the melt pool formation as well as the melting & solidification principles were illustrated. Apart from the general conservation equations, some user-defined source terms were presented. The VOF (volume of fluid) method was introduced to track the free interface between different phases. A theoretical volumetric heat source was developed to apply the laser beam energy on the free interface, which would be added to the energy conservation equations. Mass sources in the form of powder and wire were discussed theoretically which would be added to the mass conservation equations. Momentum sources including surface tension forces, gravity forces and momentum sink during melting & solidification process were applied to modify the momentum conservation equations.

3.2 Previous work

Not too many studies have been conducted to determine the possible effects of gravity during laser metal deposition process because most of the previous work are performed in the flat position. However, along with the advancement of laser manufacturing techniques, laser metal deposition is increasingly being used in a wide variety of environments including surface cladding of turbine blade shroud and interlock, offshore drilling heads, cylinder body, sleeve and mould side walls, where various deposition directions are normally required.

A few efforts are found to achieve a good deposition quality on vertical position by designing a special powder feeding nozzle [200, 201]. Paul et al.[118] concluded from their modelling and experimental work that during vertical cladding more deposition

Chapter 3 Formulation of a CFD model for laser additive manufacturing process would shift downwards if a lower process velocity was applied. The reason was that shorter interaction time would result in a shorter period for the deposition to remain in the molten state where higher effective viscosity would counter the flow of melt deposit under gravity. They discussed the effect of laser power, wire feed rate and process velocity but neglected the role of surface tension force which was highly related to the laser spot size and the effect of multiple layer deposition was also not considered. They presented a model which was in steady-state mode coupled with enhanced thermal conductivity method representing the thermo-capillary phenomenon. Too many assumptions were made so that it could not fully cover all the happening mechanisms during the process.

Some other researchers investigated the factor of gravity during the melt pool development in the non-flat welding process. Cho et al. [202] performed their work on V-groove GMAW for various welding positions, and their results indicated that a fully penetrated weld bead could be hard to achieve without a root gap in flat and overhead position while humping and melt-through beads were more likely to appear in vertical up welding position. Thomy et al. [203] designed a fibre laser welding system which was capable of producing good welds in the construction of pipelines. Muhammad et al. [204] discussed the gravity effect during laser welding at eight different positions and concluded that gravity had little impact on the bead shape and flow structure but could change the pore structure considerably. The little influence on melt pool shape might attribute to the lack of material addition. Kumar et al. [205] suggested that the workpiece orientations would significantly affect the free surface profile in GMA fillet welding. They also suggested that the average cooling rate for vertical down welding was higher than vertical upward or horizontal positions, while the cooling rate would decrease with the increase of heat input per unit length. Guo et al. [119] discovered that the sagging and undercut defects could be avoided by applying welding in horizontal position where the pressure in the centre of melt pool was relatively low in comparison with the pressure in flat welding position. The low pressure would make it possible for the surface tension to balance the hydrostatic pressure brought by gravity force. However, although a considerable amount of research can be found on welding in different orientations, very little attention has been paid to laser metal deposition in non-flat positions.

With the rapid development of computational industry, there has been more and more software coming to the market and increasing number of people begin to use numerical

Chapter 3 Formulation of a CFD model for laser additive manufacturing process methods to study the law for engineering experiments. In the field of laser metal deposition, two methods are widely used to simulate the complicated process. FEM method is mainly applied to model the process of heat transfer and residual stress formation while FVM method is mostly used to help investigate the melt pool behaviour and liquid flow. A recent research done by Balley et al. [160] presented a model to predict the transformation of solid phase, material hardness and residual stresses during multilayer and multitrack laser metal deposition. Long et al. [206] discussed the impact of deposition patterns on cracking failure by looking into the thermal stress using Ansys. Three patterns were introduced in consideration of the sample geometry including long- edge parallel reciprocating scanning method (LPRS), short-edge parallel reciprocating scanning method (SPRS) and inter-layer orthogonal direction-changing parallel reciprocating scanning method (IODPRS). A Coupled temperature and stress/strain field model was built by Alimardani et al. [87] to investigate the effect of preheating and clamping on reducing the thermal stresses. Yong et al. [207] established a numerical model for transient thermal analysis and their numerical results were compared with the data obtained from a vision detection system. A model conducted by Zhenguo et al. [208] using Abaqus simulated the process of laser hot-wire additive manufacturing which included temperature, stress and strain fields and distortions. For FEM, a simplified thermal assumption is normally made while residual stress, distortions and metallurgical transformation can be taken into account. The biggest drawback of FEM method is the difficulty in modelling material flowing process which is crucial during laser metal deposition process. Only solidus material is applied in most of FEM and the way to achieve material melting and flowing is to adjust the mesh parametrically (Element Birth and Death method in Ansys). However, the geometry of deposition will have to be predefined and hence cannot reflect the real purpose of modelling if geometry itself is the targeted result.

On the contrary, FVM method is more suitable to compute a more realistic velocity and thermal field inside the melt pool. Among the models built based on FVM methods, there are also two categories regarding if multiple phases are considered. One is modelled with fixed top surface assuming that the melt pool surface is flat and will not deform according to the fluid flow. The other method includes free surface tracking between different materials (normally gas and liquid) and more realistic results can be achieved in comparison with the non-deformable surface. An anisotropic enhanced thermal

Chapter 3 Formulation of a CFD model for laser additive manufacturing process conductivity approach was applied by Safdar et al. [209] for the modelling of laser melt pools. Heat transfer caused by convection flow was not considered in their model. Instead, thermal conductivity values for material higher than liquidus temperature were enhanced five times their original values. Unlike isotropic enhanced thermal conductivity method, the enhancement in thermal conductivity was only applied in the cross-section direction (no scanning and depth direction) in anisotropic enhanced thermal conductivity method. A coupled heat transfer and fluid flow model was built by Zhang et al. [210] for a laser- GMAW hybrid process with the existence of keyhole. The factor of gas phase was also neglected in their research. Pan et al. [211] proposed a three-dimensional model to investigate the transport phenomenon in melt pool and weld bead formation. Defects including humping and undercut effect were discussed. Meng et al. [212] also presented a numerical model to discuss the undercut phenomenon during the gas tungsten arc welding process.

Although FVM method can be used to simulate the fluid flow during the process, most of the previous work has been focusing on the development of the melt pool only no matter if the free surface is considered or not. Not too many studies have considered modelling the process of adding filler material into the melt pool, which will bring more difficulties to the calculation process. No report so far has been found on modelling the whole deposition process with complex geometries which also takes into account the gravity effect in various deposition directions.

This study therefore sets out to investigate the impact of gravity on laser metal deposition in different operating positions. A three-dimensional transient model with VOF free surface tracking was built using CFD software Fluent 14.5. Process parameters related to the melt pool formation including process velocity, laser power, laser spot size and material feed rate were investigated. User-defined mass, momentum and energy source terms were added to the model in consideration of the compound process phenomena covering material addition, surface tension, Marangoni stress, buoyancy force, temperature dependent material properties, heat loss and moving Gaussian laser beam heat input. History of melt pool area on the surface and total melt pool volume were extracted from the model to get a better understanding of the melt pool formation. Results of the deposited bead geometry obtained from the simulation would be compared with the experimental data. The following sections will describe the modelling principles involved in the laser metal deposition process.

Chapter 3 Formulation of a CFD model for laser additive manufacturing process

3.3 Governing equations

Finite volume method (FVM) was applied to the discretization process. The calculation domain was discretized into a finite number of control volumes and governing equations were integrated over the control volumes into which the domain was discretized. The Gauss theorem was applied to transform the volume integrals of the convection and diffusion terms into surface integrals. Figure 3.1 shown below illustrates the conservation in a discrete element.

Figure 3.1 Conservation in a discrete element [213] Mass and momentum conservation equations were solved by default from the software, and additional user-defined mass and momentum terms were added into the model. Considering that heat transfer happened during the laser metal deposition including the input of laser beam energy, additional equations for energy conservation were required.

Figure 3.2 describes the mass balance in a controlled volume element dV, where the rate of mass from different directions that enters and leaves the volume element is used to identify the mass change condition in the control volume.

Figure 3.2 Mass balance in a small volume element dV [214]

Chapter 3 Formulation of a CFD model for laser additive manufacturing process

The three-dimensional mass conservation equation is given as follows [215]:

휕휌 휕휌 휕(휌푢) 휕(휌푣) 휕(휌푤) + ∇ ∙ (휌풗⃗⃗⃗) = + + + + 푆 (3.1) ∂t ∂t ∂x ∂y ∂z 푚 where ρ is the density of the material in a calculating cell, t is time, u, v and w are the 3 component velocities in x, y and z-directions of resultant velocity 풗⃗⃗⃗, 푆푚 (kg/m ·s) is volumetric mass source addition rate.

Figure 3.3 presents the momentum transport in the x-direction in a control volume element dV, where transport acts normal and tangential to the surfaces.

Figure 3.3 Molecular transport of the momentum in the x-direction in an arbitrary small volume element dV [214]

Fluid in this study was assumed as Newtonian with laminar flow and the momentum conservation equation can be given by equation (3.2) as [215]

휕 (휌풗⃗⃗⃗) + ∇ ∙ (휌풗⃗⃗⃗풗⃗⃗⃗) = −∇푝 + ∇ ∙ (휏̿) + 휌푔⃗ + 퐹⃗ + S (3.2) 휕푡 푚표푚 where 푝 is static pressure, 휏̿ is the stress tensor due to the viscous stress on element surface as a result of viscosity between molecules, 휌푔⃗ and 퐹⃗ are gravitational body force and external body forces. In this case, buoyancy force due to expansion was included as a body force which will be described below. Gravitational body force in various directions was applied as a user-defined momentum source on the material with temperature higher than the melting point. Due to the fact that momentum brought by the filler wire entering into the pool would play important roles during development of the free melt pool surface, some user-defined momentum sources were added accordingly as external body forces.

Laser beam energy was applied as the heat source on free surface cells within the laser beam radius. The heat input was transferred through convection, conduction and radiation

Chapter 3 Formulation of a CFD model for laser additive manufacturing process to the substrate and surrounding atmosphere. Thermal energy conservation equation can be presented as [216]:

휕(휌퐻) + ∇ ∙ (휌푣⃗퐻) = ∇ ∙ (푘∇푇) + 푆 (3.3) 휕푡 ℎ where 퐻 is enthalpy, 푘 is thermal conductivity and 푆ℎ is volumetric heat source term applied on the interfacial cells between air and liquid material.

3.4 VOF model

The VOF model was firstly reported by Hirt and Nichols [169]. It can simulate two or more immiscible fluids by tracking volume fraction of different phases in each cell throughout the domain. For each additional new phase, a variable called volume fraction of phase is introduced to present volumetric percentage of this material in the cell. The volume fraction of each phase in a computational cell is all known and the sum of these fractions equals to unity. The liquid flowing variables are shared by all phases while the material property of this cell is the average volume values of each phase. Since only a pure phase or a mixed phase can occur in a certain computational cell, for phase 푞 with volume fraction of 퐹푞 only three situations can happen shown in Figure 3.4:

(1) 퐹푞 = 0: no phase 푞 in the cell;

(2) 퐹푞 = 1: phase 푞 occupies all the cell;

(3) 0 < 퐹푞 < 1 : phase 푞 mixed with other phases on the interface.

Figure 3.4 (a) Interface between two phases. (b) VOF function of the two liquids [217]

If a cell is occupied by two phases with volume fraction 퐹1 and 퐹2, density 휌1and 휌2 respectively, density of this cell can be calculated as:

Chapter 3 Formulation of a CFD model for laser additive manufacturing process

휌 = 퐹1휌1 + 퐹2휌2 = 퐹1휌1 + (1 − 퐹1)휌2 (3.4)

Hence density of a certain phase material in a cell can be described as:

휌 = 휌0퐹 (3.5) and volumetric mass source addition rate of this phase can be given as:

푆푚 = 휌0퐹̇ (3.6)

By combining the mass conservation equation (3.1) and equation (3.5), (3.6) shown above, conservation equation for volume fraction can be derived as [109]

휕퐹 + ∇ ∙ (푣⃗퐹) = 퐹̇ (3.7) 휕푡

Two phases including air and metal were considered in this research so that time-varying interface between air and substrate could be tracked which moved accordingly to the melting and fluid flow condition inside the pool. Therefore, development of the melt pool and deposition formation could be tracked and predicted.

3.5 Surface tension

Surface tension arises as a result of different attractive forces between molecules in a fluid. During laser processing, Marangoni flow is driven by surface tension difference attributing to the temperature difference on melt pool surface. Fluid will always flow from low surface tension area to high surface tension area, and hence a continuous convective flow will be formed inside the melt pool before solidification.

훔풕

훔풏 훁푻

Free interface

Figure 3.5 Normal stress and tangential stress on the free interface

Chapter 3 Formulation of a CFD model for laser additive manufacturing process

The continuum surface force mode (CSF) firstly proposed by Brackbill et al. [218] is implemented in Fluent to simulate the surface tension forces normal to the interface. The pressure drop on both sides of the liquid can be defined as:

1 1 푝2 − 푝1 = 훾( + ) (3.8) 푅1 푅2 where 푝2 and 푝1 are the pressures on two sides of the fluid, 훾 is the surface tension coefficient and 푅1 푅2 are the two radiuses defining the surface curvature. In Fluent, the surface curvature is calculated as surface normal gradient which is also defined as volume fraction gradient of a certain phase on the interface shown as [169]:

∇퐹 κ = ∇ ∙ 푛̅ = ∇ ∙ (3.9) |∇퐹| where κ is surface curvature, n is surface normal and F is phase volume fraction. For a simulation between two phases i and j with surface tension coefficient 훾푖푗, the surface tension force term added to the momentum conservation equation can be expressed as:

휌푎푣푔κ푖∇퐹푖 푆푛표푟푚푎푙 = 훾푖푗 1 (3.10) (휌 +휌 ) 2 푖 푗 where 휌푎푣푔 is the average density of the interfacial cell related to the phase volume fraction, 휌푖 and 휌푗 are the density of phase i and j respectively. Marangoni stress on the tangential direction of the surface induced by the temperature dependent surface tension coefficient can be presented as:

휕훾 휕푇 휏 = (3.11) 푀푎푟푎푛푔표푛푖 휕푇 휕풏

Mills et al. [114] introduced that the surface tension coefficient was a function of temperature and sulphur concentration which is shown as follows:

훥퐻0 −( ) 0 0 푅푇 γ = γ푚 − 퐴(푇 − 푇 ) − 푅푇Г푠푙푛 [1 + 푎푖푘 푒푥푝 ] (3.12)

0 0 where γm is the surface tension of pure metal at reference melting temperature T , A is a coefficient for the variation of surface tension at temperature T above the liquids, R is gas constant, Г is surface excess at saturation, k is entropy segregation constant, ΔH0 s is the enthalpy of segregation, ai is the activity of species i in solution.

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3.6 Melting and solidification

Enthalpy-porosity technique firstly introduced by Voller et al. [216, 219] was applied in this work to simulate the material melting and solidification process. Liquid fraction 푓 with a value between 0 and 1 indicating the fraction of liquid cell volume was applied in this model. The enthalpy of the material can be described as:

푇 H = ℎ푟푒푓 + ∫ 퐶푝푑푇 + 푓퐿 (3.13) 푇푟푒푓 where ℎ푟푒푓 is reference enthalpy, 푇푟푒푓 is reference temperature, 퐶푝 is specific heat at constant pressure, 퐿 is latent heat of the material. Liquid fraction 푓 can be determined as:

0 𝑖푓 푇 ≤ 푇푠표푙푖푑 푇−푇푠표푙푖푑 푓 = { 𝑖푓 푇푠표푙푖푑 < 푇 < 푇푙푖푞푢푖푑 (3.14) 푇푙푖푞푢푖푑−푇푠표푙푖푑 1 𝑖푓 푇 ≥ 푇푙푖푞푢푖푑 where 푇푠표푙푖푑 and 푇푙푖푞푢푖푑 are the material solidus and liquidus temperature respectively. The mushy region was regarded as a porous medium in enthalpy-porosity model where liquid velocity was dramatically decreased to zero when transforming from liquidus to solidus phase in order to simulate the process of solidification. A momentum source for decreasing the velocity in mushy zone [220] can be expressed as:

(1−푓)2 푆 = 퐴 푣⃗ (3.15) 푑 (푓3+휀) 푚푢푠ℎ where 푓 is the liquid fraction, 휀 is a small number 0.001 in case 푓 reaches 0, 퐴푚푢푠ℎ is mushy zone constant which indicates the damping amplitude for the liquid velocity to zero during solidification. Considering that laser metal deposition was a high cooling rate process which was combined with big temperature gradient and high solidification rate, a relative big 퐴푚푢푠ℎ value (1e7) was chosen to realize the rapid cooling.

3.7 Buoyancy force

Due to the temperature gradient in the melt pool, density difference would occur which would generate buoyancy force. The direction of convection flow induced by thermal expansion was vertically upward against the gravity direction which would promote a wider and shallower melt pool on a flat deposition surface. Not too much work has been

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Chapter 3 Formulation of a CFD model for laser additive manufacturing process done on the effect of buoyancy force in non-flat melt pool formation especially when the filler material is added to the melt pool. A buoyancy source term was added into the momentum conservation equation to simulate the thermal expansion induced convection flow. Material density was regarded as constant in other conservation equations apart from the buoyancy term in the momentum equation to get a faster convergence during the whole calculation. Boussinesq approximation [221] model can be expressed as:

푆푏푢표푦푎푛푐푦 = (휌 − 휌푚)푔 ≈ −휌푚훽(푇 − 푇푚)푔 (3.16) where 휌푚 is density of the material at melting temperature 푇푚, 훽 is the thermal expansion coefficient. An approximation was made where 휌 = 휌푚(1 − 훽∆푇) to eliminate the temperature dependent material density property.

3.8 Numerical implementations

The commercial CFD software ANSYS Fluent 14.5 was employed for the numerical solution. Fluent solver is based on the Finite Volume Method (FVM). The calculation domain was discretized into a finite number of control volumes and general conservation equations were solved on the set of control volumes. The flow of liquid metal in the melt pool was treated as incompressible laminar flow.

Gambit 2.4.6 was used to build the three-dimensional model and set the grid. After the model was imported into Fluent, a pressure-based transient solver with PISO (Pressure- Implicit with Splitting of Operators) pressure-velocity coupling method was applied to solve the governing equations with boundary conditions. The velocity and temperature fields were discretized with a second order upwind scheme, and the pressure field was discretized with a PRESTO! scheme. The convergence criteria for residuals of continuity and momentum equation was 10-3 and 10-6 for energy equation.

The domain size, grid size, and time step size were investigated to obtain high accuracy of the numerical solution. The domain size of the numerical model was shown to be large enough to eliminate the boundary effect on the formation of melt pool. A series of tests were conducted in the model by using finer grid sizes and smaller time steps while applying the same material properties, boundary conditions and solution types. Courant 푢∆푡 number defined as 퐶 = was used to check how fast the fluid could traverse a ∆푥 computational grid cell in a given time step, where 푢 is the magnitude of the velocity, ∆푡 102

Chapter 3 Formulation of a CFD model for laser additive manufacturing process is the time step and ∆푥 is the grid size. The grid and time step were refined until the numerical solution became independent of both the grid size and time step, while keeping the courant number 퐶 smaller than 1.

The enthalpy-porosity technique was applied to track the solid-liquid surface. The liquid fraction was computed in each iteration based on the enthalpy balance. The liquid-gas face was tracked by explicit VOF method with Geo-Reconstruct discretization scheme.

A compiled user-defined function (UDF) was introduced to calculate the position of laser heat source at a given time as a function of the process velocity and the Gaussian heat distribution in terms of spatial coordinates. The moving laser beam was applied as a volumetric heat source on the interfacial cells between air and metal with 0.05 cut-off value of metal volume fraction (0.05 < 퐹푚푒푡푎푙 < 1). The UDF was written in C language and linked to the FLUENT solver.

3.9 User-defined energy, mass and momentum sources

3.9.1 Energy source

Gaussian heat distribution of the laser beam was applied as the heat source on interface surface between air and metal material which is defined [222] as

2 푞푛 = 푞0 · exp [−퐶푟푛 ] (3.17)

where C is a concentration coefficient, 푟푛 is a radial distance from the heat source centre on the circular disk, 푞0 is the maximum heat flux at the centre of heat source model. Assume that there was no power loss when focusing the laser, total power Q can be obtained as

∞ 푞 휋 Q = 푞 ∫ exp [−퐶푟2] 2휋푑푟 = 0 (3.18) 0 0 퐶 hence

푄퐶 푞 = (3.19) 0 휋 hence

푄퐶 푞 = exp [−퐶푟2] (3.20) 푛 휋 푛

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Equation shown above indicates that heat flux tends to be zero when radial distance 푟푛 is at infinity. Some agreement is needed in order to determine in which range values of the Gaussian distribution curve can be seen as negligible. According to the work from Tseng and Aoh [223], 푞푛 = 0.05푞0 when 푟푛 = 푟0 is introduced. The relationship between concentration of Gaussian distribution and the laser beam radius 푟0 can be obtained as

3 C = 2 (3.21) 푟0

Therefore, the heat flux equation can be expressed as

2 3푄 3푟푛 푞푛 = 2 exp [− 2 ] (3.22) 휋푟0 푟0 where maximum heat flux 푞0 at the centre of the laser beam is

3푄 푞0 = 2 (3.23) 휋푟0

Heat profile (surface heat) and heat source (volumetric heat) are two different ways in Fluent to apply heat addition. Fixed wall is normally used as a boundary type for heat profile while for heat source method heat input is directly applied to a group of selected interior cells. Heat profile can be defined as heat conduction, radiation, convection in the unit of W/m2. The volumetric energy source in the unit of W/m3 is added to the selected cells for heat source method. VOF method was used in this research which was attributed to its capability of simulating the flow of two immiscible materials between air and liquid metal. The free interface between these two materials could be tracked and in this case the interface was the targeted deposited geometry which would undergo melting and solidification process. Considering the complex phenomenon included in the process especially the factor of surface tension, free surface tracking was necessary and hence a second phase air should be considered. The volume fraction of metal phase in a certain cell was always changing with the running time due to the varying interface between two phases. Therefore, the heat profile which could only be applied on a fixed wall was not applicable. Heat source added to cells on the interface was the better option.

During the laser manufacturing process with conduction mode, heat flux from the laser beam is impacted on the sample surface and the heat is transferred into the sample by conduction. Considering that the mesh in the model was relatively small with the length

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Chapter 3 Formulation of a CFD model for laser additive manufacturing process scale of 150μm, volumetric heat source applied on the first interface layer could still be regarded as a surface heat source.

Figure 3.6 Comparison between cells on the curved interface and flat surface Centre of the heat source was firstly represented by a time-varying function which was related to the laser moving velocity. Then cells within the radius range were selected and the heat source was applied to these cells. With a concave shaped deposition profile shown in Figure 3.6, more cells on the interface would be included within radius range in comparison with cells on the flat bottom. If the Gaussian distribution was applied arbitrarily on the free surface, total heat input would be higher than the actual laser generated energy. With changing cell numbers on the interface, the total heat input was changing accordingly. Therefore, some solution was introduced to keep the total energy input conservative.

Assume that the Gaussian distributed laser beam is applied to interfacial cells which have a total number of n. The conservation equation for the total heat input can be expressed as

푛 퐻푆 ∗ 푇 ∗ 푉 = 푃 ∗ 푇 ∗ 휂 (3.24) ∫1 푛 푡푖푚푒푠푡푒푝 푐푒푙푙 푙푎푠푒푟 푡푖푚푒푠푡푒푝

3 where 퐻푆푛 (W/m ) is the heat source term on a certain cell which will be added to the energy equation in Fluent solver, 푇푡푖푚푒푠푡푒푝 (s) is the actual time of each calculation 3 interval, 푉푐푒푙푙 (m ) is the cell volume on which 퐻푆푛 is added, 푃푙푎푠푒푟 (W) is the laser power used for the laser metal deposition process and 휂(%) is the laser power efficiency considering energy loss during absorption and reflection. The total heat source input 3 퐻푆푡표푡푎푙 (W/m ) can be described as

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푛 푃푙푎푠푒푟∗휂 퐻푆푡표푡푎푙 = ∫ 퐻푆푛 = (3.25) 1 푉푐푒푙푙

3 According to the equations deducted above, 퐻푆푛 (W/m ) has a relationship with the 3 maximum heat source 퐻푆0 (W/m ) in the centre of the laser beam and it can be shown as

3∗푟2 퐻푆 = 퐻푆 ∗ exp (− 푛) (3.26) 푛 0 푟2

3 Hence 퐻푆푡표푡푎푙 (W/m ) turns into

푛 3∗푟2 푛 3∗푟2 퐻푆 = ∫ 퐻푆 ∗ exp (− 푛)=퐻푆 ∗ ∫ exp (− 푛) (3.27) 푡표푡푎푙 1 0 푟2 0 1 푟2

By combining equation (3.25) and (3.27)

2 푛 3∗푟푛 푃푙푎푠푒푟∗휂 퐻푆0 ∗ ∫ exp (− 2 ) = (3.28) 1 푟 푉푐푒푙푙

3 Maximum heat source 퐻푆0 (W/m ) in the centre then can be expressed as

푃 ∗휂 퐻푆 = 푙푎푠푒푟 0 푛 3∗푟2 (3.29) 푉 ∗∫ exp(− 푛) 푐푒푙푙 1 푟2 and heat source 퐻푆푛 applied on a cell with a certain radial distance 푟푛 can be presented as;

푃 ∗휂 3∗푟2 퐻푆 = 푙푎푠푒푟 ∗ exp (− 푛) 푛 푛 3∗푟2 푟2 (3.30) 푉 ∗∫ exp(− 푛) 푐푒푙푙 1 푟2

Since laser beam is assumed to be moving on xy plane in the direction of y, 푟푛 can be calculated as follows:

2 2 2 푟푛 = [푦푛 − (푣푙푎푠푒푟푡 + 푦0)] + (푥푛 − 푥0) (3.31) where 푣푙푎푠푒푟 (m/s) is laser process velocity, 푡 is the calculation flowing time, 푥0 and 푦0 is the location of starting point on the substrate.

For top-hat mode laser profile, the heat source is evenly distributed within the effective laser radius which can be expressed as

푃푙푎푠푒푟∗휂 퐻푆푛 = (3.32) 푛푉푐푒푙푙

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In conclusion, due to the fact that free surface cells are changing accordingly to the changing melt pool geometry after each calculation time step, the heat source is redistributed depending on the number of cells which is included. Therefore, the heat source model can self-adaptively change with the development of melt pool and deposition profile to meet the conservation of total heat input.

3.9.2 Mass source

There is a notable paucity of studies which focus on adding mass material into the melt pool, especially when it is coupled with free surface development. Most of the modelling work performed in the field of laser deposition/welding process was contributed to investigating the temperature history on a fixed flat surface. Like laser energy addition, wire material feeding into the melt pool can be realized in Fluent with two different methods. Mass flux on a customised boundary profile is implemented on a surface wall where targeted mass flow rate can be added as a function of physical flowing time and spatial position. One drawback of this method is that the surface wall is predefined and mostly fixed. Although one research from Amara et al. [224] indicated that dynamic mesh method combined with heat flux on the surface wall could be utilised to model the mass addition process, the amount of deforming mesh by the time was predefined and could not be used to predict the deposition result before realistic experiments were performed. Considering the melt pool surface is always evolving with the flowing time, the volumetric mass source should be added to the selected cells in between each real-time step. The number of free surface cells will be determined after each time step and mass source term will be added on these cells accordingly during the next time step.

For the deposition process using wire as filler material, the mass addition rate can be regarded as evenly distributed in each cell within the wire radius range. Assume that there are cells with a total number of n located on the interface between air and liquid material, total mass addition during each time step can be presented as

푛 ∗ 푀푆 ∗ 푇푡푖푚푒푠푡푒푝 ∗ 푉푐푒푙푙 = 퐹푤푖푟푒 ∗ 푇푡푖푚푒푠푡푒푝 (3.33) where 푀푆 (kg/m3·s) is the mass source term for each selected cell which will be added to 3 the continuity equation in the solver and 퐹푤푖푟푒 (kg/m ·s) is mass addition rate of the wire. Mass source then can be expressed as:

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Chapter 3 Formulation of a CFD model for laser additive manufacturing process

퐹 푀푆 = 푤푖푟푒 (3.34) 푛∗푉푐푒푙푙 where 퐹푤푖푟푒 can also be calculated as

2 퐹푤푖푟푒 = 휌푤푖푟푒 ∗ 휋푟푤푖푟푒 ∗ 푓푤푖푟푒 (3.35)

3 where 휌푤푖푟푒 (kg/m ) is the density of wire material, 푟푤푖푟푒 (m) is the radius of wire and

푓푤푖푟푒 (m/s) is wire feed rate. Mass source can be determined as follows:

휌∗휋푟2 ∗푓 푀푆 = 푤푖푟푒 푤푖푟푒 (3.36) 푛∗푉푐푒푙푙

When powder is used as the filler material, powder can also be regarded as Gaussian distribution within the melt pool when reaching the substrate, which can be expressed as:

푓 ∗휂 3∗푟2 푀푆 = 푝표푤푑푒푟 푝표푤푑푒푟 ∗ exp (− 푛) 푝표푤푑푒푟 푛 3∗푟2 푟2 (3.37) 푉 ∗∫ exp(− 푛) 푐푒푙푙 1 푟2

3.9.3 Momentum source

If mass enters the calculation domain without additional momentum, the mass will have to be accelerated by the liquid flow and this may cause a drop in the velocity field. The titled angle between sample substrate and wire feeding nozzle as well as the wire tip distance above the sample surface plays a crucial role during the development of the melt pool and deposition profile. Therefore, the addition of momentum brought by the wire should be taken into consideration. Assume that the mass source is added directly to the interfacial cells within the wire radius range, wire material entering into the melt pool has a velocity with two components which is shown in Figure 3.7. One is the wire feeding velocity coming from the wire feeder and the other is in the same direction with the laser beam. Since the wire feeder nozzle is attached on the laser head, the wire feed speed in the laser moving direction is the same as laser process velocity.

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Chapter 3 Formulation of a CFD model for laser additive manufacturing process

Figure 3.7 Schematic of the laser metal deposition process Therefore, momentum source (kg/m2·s 2) on the z-direction can be determined as follows:

2 2 휌∗휋푟푤푖푟푒∗푓푤푖푟푒∗푠푖푛훼 푀표푚푧 = − (3.38) 푛∗푉푐푒푙푙 while momentum source on the y-direction can be expressed as:

2 휌∗휋푟푤푖푟푒∗푓푤푖푟푒 푀표푚푦 = ∗ (푓푤푖푟푒 ∗ 푐표푠훼 − 푣푙푎푠푒푟) (3.39) 푛∗푉푐푒푙푙 where α is the angle between wire feeder and sample surface, and 푣푙푎푠푒푟 (m/s) is laser process velocity. The momentum source term applied on the interface cells will be added to the momentum conservation equation (3.2).

3.10 Boundary conditions

Two phases were created illustrated in Figure 3.8 considering that free interface tracking between air and substrate material was essential throughout this whole research. The boundary condition was categorized into two groups including internal and external boundaries. Internal boundary referred to the free surface of melt pool while external boundary was the boundary of the whole computational domain.

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Conduction and radiation Conduction and radiation

Conduction and radiation Conduction and radiation Figure 3.8 Boundary conditions of the model

3.10.1 Internal boundary condition

VOF model was updated every time step to track the liquid-air free surface while mass, momentum, and energy source were added on this interfacial boundary. Mass source in the form of wire or powder feeding was added to the melt pool within wire/powder radius range which can be presented as:

휕퐹 푀푆 + ∇ ∙ (푣⃗퐹) = (3.40) 휕푡 휌 where F is volume fraction of the cell, 푣⃗ is velocity vector, 휌 is material density and 푀푆 is the self-adaptive mass addition rate applied on the free surface cells.

Energy boundary on the melt pool interface included laser heat input and heat loss which covered heat conduction transferring into the deep substrate, heat radiation into the air and heat convection between the hot surface and flowing air, which can be expressed as follows:

휕푇 푘 · = 퐻푆 − ℎ (푇 − 푇 ) − 휎휖(푇4 − 푇4 ) (3.41) 휕풏 푛 푐 푒푛푣 푒푛푣 where 푘 is the material conductivity, 퐻푆푛 is a self-adaptive moving heat source, ℎ푐 is heat convection coefficient, σ is Boltzmann constant, ε is Radiation emissivity, and T푒푛푣 is environment temperature. During the deposition, on the free interface conduction was the dominant method of heat dissipation compared to convection and radiation to the surrounding environment. A convection coefficient of 25 W∙m-2 was applied to model convection heat loss by argon shielding gas [173].

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Chapter 3 Formulation of a CFD model for laser additive manufacturing process

Volumetric forces implemented on the free surface included Marangoni stress tangential to the surface, surface tension force in the normal direction of the interface and momentum brought by the feeding wire velocity, which can be calculated as:

S푚표푚 = 퐹푠푢푟푓푎푐푒 푡푒푛푠푖표푛 + 퐹푀푎푟푎푛푔표푛푖 + 푀표푚푦 + 푀표푚푧 (3.42) where 푀표푚푦 and 푀표푚푧 are the momentum brought by wire feeding. These four terms are described in equations 3.10, 3.11, 3.38 and 3.39 respectively. Apart from the internal source terms applied on the free interface mentioned above, body forces including gravity and buoyancy in the form of Boussinesq approximation on liquid material were also added to the equation.

3.10.2 External boundary condition

For external boundaries of the whole calculation domain, boundaries AB, AF and EF (Figure 3.8) were treated as outflow. Thermal conduction and thermal radiation were applied on substrate side (BC, ED) and bottom (CD) walls. The energy balance equation is shown as:

휕푇 푘 · = −휎휖(푇4 − 푇4 ) (3.43) 휕풏 푒푛푣

Considering that the calculation domain was selected as part of the whole workpiece, heat loss caused by thermal conduction from side walls towards the surrounding substrate material should also be taken into account. Therefore, a negative heat flux was applied on the external steel boundaries assuming that heat was conducted to a wall with constant temperature 300 K in order to realize the effect of heat conduction loss. The distance between the wall and steel external boundary was set to be 3mm on both the sidewalls and bottom considering that the substrate was mounted on a stage made of stainless.

3.11 Temperature gradient and solidification rate

As has been discussed in the previous literature review sections, temperature gradient G and solidification rate R together play significant roles in determining formation of the deposition microstructure. Cooling rate G·R determines the dendrite spacing while ratio G/R decides morphology of solidification structure. It is rather difficult to measure the exact thermal conditions inside the melt pool due to the fact that melt pool size is very

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Chapter 3 Formulation of a CFD model for laser additive manufacturing process small and the rapidly flowing molten material inside pool also add difficulties to obtain an accurate measurement. Therefore, numerical methods are normally applied to help get a better understanding of these two important factors.

Due to the high energy density of laser beam and relatively small heat affected zone, temperature gradient inside the melt pool can be extremely high during laser metal deposition process. Temperature gradient on the solidification boundary also varies with the location on the interface and direction of the temperature gradient is pointed normal to the heat flux on solidification front. In this model, the temperature gradient is obtained firstly by calculating components Gx, Gy, Gz on x, y and z-direction, where the derivative of temperature with respect to the distance along three axes are calculated respectively. Resultant temperature gradient G can be obtained as:

2 2 2 |푮| = 퐺 = √퐺푥 + 퐺푦 + 퐺푧 (3.44)

Solidification rate also known as growth rate is defined as the speed of grain growth on the liquid-solid interface. Solidification rate is in the same direction with temperature gradient towards the maximal side which is normal to the solidification front. It is highly related to the laser process velocity and can be defined as [225]:

|푹| = 푹푳 ∙ 풏 = |푹푳| ∙ cosθ (3.45) where 푹푳 is laser process velocity, 풏 is normal vector and θ is the angle between laser process direction and liquid-solid interface normal. θ varies with the specific location on the interface and it is influenced by different melt pool shape as a result of different process parameters. Growth grate reaches the highest at tailing region of the melt pool where θ is minimal and the minimum occurs at outmost of the melt pool where heat flux is perpendicular to laser process velocity. In this model, moving heat source is along y direction and cosθ can be expressed as [226]:

퐺 cosθ = 푦 2 2 2 (3.46) √퐺푥 +퐺푦+퐺푧 and cooling rate 푇̇ :

퐺푦∙푅퐿 2 2 2 푇̇ = 푅 ∙ 퐺 = ∙ √퐺푥 + 퐺푦 + 퐺푧 = 퐺푦 ∙ 푅퐿 2 2 2 (3.47) √퐺푥 +퐺푦+퐺푧

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Chapter 3 Formulation of a CFD model for laser additive manufacturing process and ratio:

퐺 퐺2+퐺2+퐺2 = 푥 푦 푧 (3.48) 푅 퐺푦∙푅퐿

Effect of gravity on these factors including distribution of temperature gradient, solidification rate, cooling rate and ratio G/R were compared between different gravitational conditions and various locations inside the melt pool.

3.12 Summary

This chapter presented a full description of a three-dimensional model which covered free surface tracking, melting and solidification process, interfacial surface forces and various boundary conditions where the user-defined mass, energy and momentum source terms were applied. A knowledge gap was also found from the previous study on the lack of research in investigating the gravity effect during deposition process in different operating directions. The gravitational source term was included in this model which provided a good numerical environment for getting a better understanding of the gravity effect. Some experimental work will be presented in the next chapter to validate this model.

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4.1 Introduction

In this chapter, the material and equipment used in coaxial powder deposition process were first introduced, followed by the description of experimental details to investigate the influence of different process parameters on deposition results including bead height, bead width, dilution and powder utilisation. Laser power, process velocity and powder feed rate were varied during the experiments. Multiple layer deposition was successfully performed and the effect of deposition patterns was discussed when depositing geometries with complex features. The effect of the re-melting process on deposition surface finish was also investigated.

A three-dimensional CFD model was developed to simulate the single layer deposition process with different process parameters and the obtained modelling results were compared with the experimental results. Simulation of the re-melting process applied on already deposited geometries was first put forward, and the surface improvement brought by re-melting was validated. Apart from some simple geometries, deposition with complicated structures in full three dimensions with overhanging features was also first modelled and the effect of gravity was taken into consideration.

4.2 Material

Stainless steel is widely used in all industrial fields owing to its characteristic of excellent corrosion resistance, thermal stability, polishability and weldability. The content of chromium in stainless steel is generally more than 11% due to the fact that chromic oxides closely attached to the surface can prevent the material from being further oxidized. The oxide layer is thin enough through which the metallic lustre can be observed from the surface. Once the surface layer is destroyed, exposed steel surface will react with air and new oxidation layer will be generated to keep on functioning as the protection. According to the different phase content, stainless steel can be categorized into several types

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Chapter 4 Study and prediction of laser metal deposition with various structures including austenite, ferrite, martensite and duplex stainless steel. In addition, stainless steel can also be divided into classes by different elemental composition.

316L stainless blocks with dimensions of 50 mm × 50 mm × 10 mm were used as the substrate during the deposition process. Stainless steel 316L powder from Höganäs Belgium shown in Figure 4.1 was used as the deposition powder in the experiment.

Figure 4.1 Stainless steel 316L powder The diameter of the powder ranges from 50 μm to 150 μm. The chemical composition of 316L is listed in Table 4.1 [227].

Table 4.1 Chemical composition of 316L powder [227]

Chemical C Ni Fe Cr Si O N Mo S Contents (%) 0.02 13.0 base 16.8 0.85 0.20 0.04 2.20 0.03

Owing to the existence of element Mo which will increase the corrosion resistance in chloride solutions, 316L has a broad range of uses in the marine environment. Moreover, good gloss on the products can be achieved after the cold rolling process. Excellent working hardening ability with weak-magnetic property and good high-temperature strength are some other advantages of 316L stainless steel.

4.3 Experimental equipment

4.3.1 Laser applied

The laser used for the experiments was provided by Laserline LDL 160-1500 diode laser with maximum 1500 W power as shown in Figure 4.2. Fibre cable with 1mm diameter and 5 m long was used to deliver the laser beam. In order to compensate the energy loss

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Chapter 4 Study and prediction of laser metal deposition with various structures during the delivery of laser beam through the fibre cable, actual laser power was measured using a Gentec UP55N-300F-H9 power meter. The measured power corresponding to the nominal power is shown in Figure 4.3. The output laser beam comprised duo wavelength of 808 nm and 940 nm (infrared). The mode of the laser beam was top-hat distribution with a circular shape. The diameter of laser spot size was measured by radiating the laser beam for 1 s with 25 W laser power on a ceramic plate. Due to the characteristic of low thermal conductivity for ceramic material, spot size could be obtained by measuring the heat affected zone using an optical microscope. Figure 4.4 was plotted by adjusting the distance between the nozzle tip and deposition plane.

As can be observed from Figure 4.3, energy loss due to the fibre delivery took up approximately 35% of the total nominal energy. It is shown in Figure 4.4 that with the increase of the distance between nozzle tip and substrate, laser spot size would firstly decrease reaching the minimum point at around 7.5 mm, after which diameter of the laser started to become bigger. Therefore, the laser beam focal point was located 7.5 mm below the nozzle tip.

Figure 4.2 Laserline LDL 160-1500 Diode laser

Measured power output and nominal power 1200

1000

800

600

400

200 Measured Measured Power (W)

0 0 200 400 600 800 1000 1200 1400 1600 Nominal Power (W) Figure 4.3 Measured power output and nominal power

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Laser spot size variation 2.3

2.1

1.9

1.7

Laser Laser spot size (mm) 1.5 4 5 6 7 8 9 Distance between nozzle tip and deposition plane (mm)

Figure 4.4 Variation between spot size and the distance from nozzle tip

4.3.2 Powder delivery system

The powder was delivered by an FST powder feeder shown in Figure 4.5. Powder feed parameters determining the powder delivery rate includes carrier gas pressure and flow rate, stirrer speed and the rotation speed of powder disk which were controlled by a panel manufactured by Siemens. The feeding system was based on the volumetric principle and equipped with two 1.5 ltr hoppers which provided precise and reproducible powder rates [228]. During the experiments, the flow rate of carrier gas and the stirrer velocity were kept constant. Therefore, the powder feed rate was dependent on the disk rotation speed. In order to investigate the variation between the disk rotation speed and powder flow rate, a container was firstly weighed and then placed on the substrate collecting the powder coming out from the nozzle for 1 minutes. The container together with powder was weighed again and the powder feed rate could be calculated by subtracting the original container weight.

Figure 4.5 FST powder feeder

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Chapter 4 Study and prediction of laser metal deposition with various structures

The coaxial nozzle together with the inner schematic design [229] of the nozzle used for the experiments is shown in Figure 4.6. It was manufactured by DeBe Lasers Ltd composed with main body made of aluminium and nozzle tip made of copper. Argon would function both as the powder-delivering gas and shielding gas protecting the deposition process from oxidation. An annular coolant channel was used to decrease the temperature generated during the process through thermal conductivity. The laser beam was focused through the central passage which was protected by argon gas preventing the optical lens from the contamination by powder bouncing from the melt pool. Four gas flows carrying metal powder would mix and travel through an annual passage converging into a consolidated powder stream after the nozzle outlet.

Figure 4.6 Coaxial nozzle used for the experiments [98] 4.4 Modelling strategy

Figure 4.7 illustrates the schematic of laser metal deposition with coaxial powder feeding. Powder was fed into the melt pool coaxially with the moving laser beam.

Figure 4.7 Schematic of laser metal deposition with coaxial powder feeding

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Chapter 4 Study and prediction of laser metal deposition with various structures

The dimensions of the whole calculation domain were 6 mm × 30 mm × 6 mm and the mesh was generated using GAMBIT which contained 370440 (42×210×42) hexahedral cells with uniform grid spacing 142 µm shown in Figure 4.8. The calculation time step 푢∆푡 was 3×10-4 s and tested by the Courant number which is defined as 퐶 = , where 푢 is ∆푥 the magnitude of the velocity, ∆푡 is the time step and ∆푥 is the grid size. Fluid cannot travel more than one cell in one time step meanning that the courant number cannot bigger than one. It took approximately 24 hours to simulate a single deposition line which ran 2.2 s in real time. With time step bigger than 3×10-4 s, more iterations would be required to reach the residual criteria during every calculation time step and hence the total modelling time would be longer. Grid and time-step independence test were carried out by reducing the grid spacing to 100 µm with 1080000 (60×300×60) hexahedral cells and decreasing the time step to 1.5×10-4 s. Little difference in simulation result was obtained with the finer grid and/or smaller time step, while it would take much longer time to complete simulating the same deposition length.

6mm

Figure 4.8 Mesh for the calculation domain A 2 mm thickness 316L substrate region was initially patched for the domain, on top of which air region with 4 mm thickness was created shown in Figure 4.9. PISO algorithm was used as the pressure-velocity coupling solution. The velocity and temperature fields were discretized with a second order upwind scheme, and the pressure field was discretized with a PRESTO! scheme. The convergence criteria for residuals of continuity and momentum equation was 10-3 and 10-6 for energy equation. Considering that the modelling of the whole deposition process was required, a transient time solution with pressure-based solver was used. VOF model combined with Geo-Reconstruct method

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Chapter 4 Study and prediction of laser metal deposition with various structures was applied to track the interface between base material and gas phase, where user- defined mass, momentum and energy source terms were added.

outflow (side and top surfaces of the upper domain)

wall (side and bottom surfaces of the lower domain)

Figure 4.9 Boundary conditions and initial phase patch of the calculation domain

(outflow)

(outflow) (outflow)

(wall) (wall)

(wall)

Figure 4.10 Illustration of boundary conditions on cross-section plan The top and side surfaces of the upper domain were gas (air) phase boundaries treated as outflow shown in Figures 4.9 and 4.10. The bottom and side surfaces of the lower domain were substrate phase boundaries defined as wall conditions. The calculation domain was chosen as part of the whole workpiece, and hence heat conduction and heat radiation were applied on the wall boundaries.

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Chapter 4 Study and prediction of laser metal deposition with various structures

Heat conduction, heat radiation and heat convection were applied on the free interface between air and substrate phases, which was tracked by the VOF method. As discussed in Chapter 3, a user-defined self-adaptive Gaussian distributed laser beam was applied as a volumetric heat source (W/m3) added to the energy conservation equations on the free interface as shown in equation (4.1).

푃 ∗휂 3∗푟2 퐻푆 = 푙푎푠푒푟 ∗ exp (− 푛) 푛 푛 3∗푟2 푟2 (4.1) 푉 ∗∫ exp (− 푛) 푐푒푙푙 1 푟2

3 where 푉푐푒푙푙 (m ) is the cell volume, 푃푙푎푠푒푟 (W) is the laser power used for the laser metal deposition process and 휂 (%) is the laser power efficiency, 푟 is laser beam radius, 푟푛 is a radial distance from the laser beam centre. 푟푛 can be expressed as:

2 2 푟푛 = √[푦푛 − (푣푙푎푠푒푟 ∗ 푡 + 푦0)] + (푥푛 − 푥0) (4.2) where 푣푙푎푠푒푟 (m/s) is process velocity, 푡 is the calculation flowing time, 푥0 and 푦0 are the locations of starting point inside the groove on the subtract. Mass source brought by the powder feeding was applied to the free interface. Powder material was fed into the melt pool travelling together with the laser beam. Powder feeding was applied on the free interface as mass source and added to the continuity equations as:

푓 ∗휂 3∗푟2 푀푆 = 푝표푤푑푒푟 푝표푤푑푒푟 ∗ exp (− 푛) 푝표푤푑푒푟 푛 3∗푟2 푟2 (4.3) 푉 ∗∫ exp(− 푛) 푐푒푙푙 1 푟2 where 푓푝표푤푑푒푟 is powder feed rate and 휂푝표푤푑푒푟 is powder utilization. Solidification and melting process was taken into consideration with the enthalpy-porosity method by applying a damping force to the mushy material to decrease the fluid velocity. Gravity force was added to the cells with temperature higher than the melting point. The Boussinesq approximation was used to present the buoyancy force to improve the calculation speed.

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4.5 Experimental results and discussion

4.5.1 Influence of laser power

A series of experiments were implemented to investigate the influence of different process parameters on deposition quality. To understand the effect of laser power on final deposition result, laser power was increased from 250 W to 1000 W while keeping process velocity and powder feed rate constant as 8 mm/s and 0.21 g/s. The distance between deposition point and nozzle was adopted as 7 mm, and the diameter of the laser beam was 1.7 mm. After deposition, samples were grinded and polished on a polishing machine. 10% aqueous oxalic acid was then used to electrolytic etch the cross-section of the samples to see the deposited microstructure. Dimensions of the track including the height and width were measured using an optical microscope. Melted area of the substrate and area of the track above were also measured to calculate the dilution.

Figure 4.11 presents the results of the single deposited track with increased laser power from layer number one to layer number nine. Process parameters and the results of measured dimensions are shown in Table 4.2. Unevenness in width could be obviously seen from Figure 4.11 when lower power was used. Powder delivered from the nozzle could not be completely melted within the melt pool which would contribute to the irregularity of final deposition. Higher deposition temperature was the main reason for the increasingly severe oxidation when applying higher laser power.

10 mm Figure 4.11 Single layer deposition using different laser power

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Table 4.2 Relationship between laser power and deposition results

Process velocity: 8 mm/s, Powder feed rate: 0.21 g/s Test number Laser power (W) Height (mm) Width (mm) Dilution 1 250 0.622 0.947 0.00% 2 300 0.689 1.063 0.00% 3 370 0.715 1.179 0.00% 4 460 0.702 1.381 8.31% 5 530 0.732 1.424 15.70% 6 600 0.875 1.541 17.54% 7 700 0.911 1.587 25.97% 8 800 0.892 1.650 32.45% 9 1000 1.018 1.796 36.71% Variation between laser power against layer height, width and dilution are plotted later in Figure 4.19, 4.22, and 4.25 respectively. Layer height, layer width and dilution would increase when increasing the laser power. More powder would be melted and deposited with higher power density defined as I=P/A, where P is the laser power and A is spot size area. No dilution between deposition and substrate could be observed when laser powers of 250 W, 300 W and 370 W were used. This could be the same as the reason for the cause of irregularity and discontinuous clad geometry, on account of the lack of sufficient power density to form a uniform melt pool. For multiple layer deposition, low dilution might lead to poor mechanical properties resulting from lack of fusion between adjacent layers.

4.5.2 Influence of powder feed rate

For the purpose of finding out how deposition result would change with the amount of powder delivered into the melt pool, powder feed rate was increased from 0.07 g/s to 0.49 g/s while keeping laser power and process velocity constant as 800 W and 8 mm/s.

Figure 4.12 provides the comparison of the sample exterior when powder feed rate was increased from clad 1 to clad 7 while Table 4.3 presents the deposition dimension of the cross-section. Both deposition height and width would increase when more powder was added to the melt pool. When low powder feed rate was applied, dilution was high which was mainly due to the reason that all the metal powder could interact well with the laser

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Chapter 4 Study and prediction of laser metal deposition with various structures beam during flight. The powder could be completely melted after being delivered into the melt pool and most of the laser energy was used to melt the substrate resulting in a high dilution. When increasing the power feed rate, energy absorbed by the powder would become higher and hence contribute to the decrease of dilution.

Figure 4.12 Single layer deposition using different powder feed rate Table 4.3 Relationship between powder feed rate and deposition result

Laser Power: 800 W Process velocity: 8 mm/s Test number Powder feed rate(g/s) Height(mm) Width(mm) Dilution 1 0.07 0.337 1.551 66.91% 2 0.15 0.564 1.652 48.37% 3 0.21 0.920 1.670 31.31% 4 0.28 1.134 1.759 20.41% 5 0.35 1.482 1.915 8.99% 6 0.42 1.579 2.139 4.56% 7 0.49 1.797 2.227 1.13%

4.5.3 Influence of process velocity

In order to investigate the influence of process velocity on deposition quality, process velocity was varied from 8 mm/s to 20 mm/s while keeping laser power and powder feed rate constant as 1000 W and 0.28 g/s.

Figure 4.13 shown below illustrates the cross-section of the deposited track when process velocity was increased from sample A to G. Both deposition width and height decreased with the increase of the process velocity. Lower specific energy would be achieved with higher process velocity which led to a reduced amount of powder being deposited on the 124

Chapter 4 Study and prediction of laser metal deposition with various structures substrate. Therefore, powder utilisation rate would be low when keeping the powder feed rate constant. No dramatic change in dilution between the track and substrate could be observed when raising the process velocity which would indicate that the allocation of laser energy was similar for melting the substrate and powder.

A B C D

E F G

500μm

Figure 4.13 Cross-section of single deposition track (A) 8 mm/s, (B) 10 mm/s, (C) 12 mm/s, (D) 14 mm/s, (E) 16 mm/s, (F) 18 mm/s, (G) 20 mm/s Table 4.4 Relationship between process velocity and deposition result

Laser Power: 1000 W Powder Feed Rate: 0.28 g/s

Process velocity (mm/s) Width(mm) Height(mm) Dilution

8 2.039 1.158 22.33%

10 1.835 0.954 25.66%

12 1.581 0.808 30.49%

14 1.596 0.711 32.89%

16 1.470 0.569 37.06%

18 1.469 0.627 32.70%

20 1.354 0.558 32.51%

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Chapter 4 Study and prediction of laser metal deposition with various structures

4.6 Modelling of single track

A three-dimensional model was built to help understand the influence of process parameters on final bead dimensions. Figure 4.14 shown below presents the modelling result with 1 kW laser power, 1.7 mm beam diameter, 0.28 g/s powder feed rate and 8 mm/s process velocity.

Figure 4.14 Simulated deposition and temperature distribution using 1 kW laser power, 1.7 mm beam diameter, 0.28 g/s powder feed rate and 8 mm/s process velocity Bead morphology and temperature distribution of the simulated results are presented in Figure 4.14. Figure 4.15 shows the melt pool development history at some moment. It can be observed that at the beginning of deposition, melt pool appeared to be elliptical shape when laser started to impact on the cold substrate causing rapid temperature increase and high-temperature gradient. Heat was transferred to the surrounding substrate by conduction. Melt pool area would expand to a maximal value with the movement of laser beam until reaching a steady condition (1.038 s) where temperature gradient became relatively stable. With further movement of the laser beam getting far away from the beginning deposition region, the melt pool appeared to be teardrop shape where the maximum solidification rate was lower than the process velocity. A bump could be observed at the beginning of deposition which was due to the surface tension-temperature variation. In order to avoid the irregularity at the beginning, the actual deposition dimension was extracted from the middle plane where the melt pool had reached steady status. The cross-sections in the middle plane (y=15 mm) of the seven cases investigating the impact of process velocity on deposition dimension are shown in Figure 4.16, where a good agreement is reached comparing with Figure 4.13.

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Chapter 4 Study and prediction of laser metal deposition with various structures

0.588 s 1.2 s

1.896 s 2.4 s

Figure 4.15 Development of melt pool and velocity field 0.588 s, 1.2 s, 1.896 s, 2.4 s

A B C D 8 mm/s 10 mm/s 12 mm/s 14 mm/s

E F G 16 mm/s 18 mm/s 20 mm/s

Figure 4.16 Effect of process velocity on deposition cross section (comparison A-G with Figure 4.13)

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Chapter 4 Study and prediction of laser metal deposition with various structures

Detailed comparisons between experimental and modelling results are presented in the charts in the following section. Figure 4.17 and 4.18 show the modelling results of deposition geometries when applying different process velocities.

A B C D E F G

v v v v v v v 8 mm/s 10 mm/s 12 mm/s 14 mm/s 16 mm/s 18 mm/s 20 mm/s

Figure 4.17 Effect of process velocity on deposition top view (comparison A-G)

A 8 mm/s v B 10 mm/s v

C 12 mm/s v D 14 mm/s v

E 16 mm/s v F 18 mm/s v

G 20 mm/s v

Figure 4.18 Effect of process velocity on deposition side view (comparison A-G) 4.7 Comparison of influence factors

Experimental results were classified into three groups including layer height, width and dilution to compare the factor level of process parameters on deposition dimension. Results obtained from simulations were also listed and compared with experimental results.

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Chapter 4 Study and prediction of laser metal deposition with various structures

Effect of laser power on depositon height 2.5 Experimental 2.0 Modelling 1.5

1.0

Height Height (mm) 0.5

0.0 0 200 400 600 800 1000 Laser power (W) Figure 4.19 Effect of laser power on layer height

Effect of powder feedrate on deposition height 2.5 Experimental 2.0 Modelling

1.5

1.0 Height Height (mm) 0.5

0.0 0 0.1 0.2 0.3 0.4 0.5 Powder feedrate (g/s) Figure 4.20 Effect of powder feed rate on layer height

Effect of process velocity on deposition height 2.5 Experimental 2.0 Modelling

1.5

1.0 Height Height (mm) 0.5

0.0 0 5 10 15 20 Process velocity (mm/s) Figure 4.21 Effect of process velocity on layer height It can be observed from the three graphs (4.19, 4.20 and 4.21) that the clad height would increase when increasing the powder feed rate and laser power while decrease when the process velocity was increased. By comparing the variation trend of the three figures, process velocity and powder feed rate played more significant role in determining the

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Chapter 4 Study and prediction of laser metal deposition with various structures layer height. Design expert will be used in the next stage to find out the accurate factor level.

2.5 Effect of laser power on depositon width Experimental 2.0 Modelling

1.5

1.0 Width(mm) 0.5

0.0 0 200 400 600 800 1000 Laser power (W) Figure 4.22 Effect of laser power on layer width

Effect of powder feedrate on deposition width 2.5 2.0 1.5

1.0 Expeimental Width(mm) 0.5 Modelling 0.0 0 0.1 0.2 0.3 0.4 0.5 Powder feedrate (g/s)

Figure 4.23 Effect of process velocity on layer width

Effect of process velocity on deposition width 2.5 2.0 1.5 1.0 Experimental

Width(mm) Modelling 0.5 0.0 0 5 10 15 20 Process velocity (mm/s) Figure 4.24 Effect of process velocity on layer width It can be concluded from the three graphs shown above (4.22, 4.23 and 4.24) that clad width would decrease when increasing the process velocity and would become wider when laser power and powder feed rate were increased. In comparison with the influential

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Chapter 4 Study and prediction of laser metal deposition with various structures degree process parameters had on deposition height, layer width only changed slightly when varying the laser power, process velocity and powder feed rate.

Effect of laser power on dilution 70.00% 60.00% Experimental 50.00% 40.00% 30.00%

Dilution 20.00% 10.00% 0.00% 0 200 400 600 800 1000 Laser power (W) Figure 4.25 Effect of laser power on dilution

Effect of powder feedrate on dilution 70.00% Experimental 60.00% 50.00% 40.00%

30.00% Dilution 20.00% 10.00% 0.00% 0 0.1 0.2 0.3 0.4 0.5 Powder feedrate (g/s) Figure 4.26 Effect of powder feed rate on dilution

Effect of scanning velocity on dilution 70.00% 60.00% Experimental 50.00% 40.00%

30.00% Dilution 20.00% 10.00% 0.00% 0 5 10 15 20 Process velocity (mm/s)

Figure 4.27 Effect of process velocity dilution Dilution in laser metal deposition process (especially for laser cladding) is often defined as the amount of intermixing of the deposition and substrate, and it plays a major role in

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Chapter 4 Study and prediction of laser metal deposition with various structures determining mechanical properties like tensile strength and yield strength. As can be observed from Figure 4.25 to Figure 4.27, no dilution would be obtained if low laser power or high powder feed rate was applied. Dilution would increase with the increase of laser power and process velocity. Process velocity had a slight influence on dilution while dramatic change would occur on dilution when varying powder feed rate and laser power.

Design-expert software was used to investigate the influence of factors on the deposition height and width. Three factors with five levels were taken into consideration. The experimental results are shown in Table 4.5 and the analysis of the results is presented below in Figures 4.28 and 4.29 and Table 4.6.

Table 4.5 Experimental results of design expert

Factor 1 Factor 2 Factor 3 B: Process A: Laser C: Powder Feed Height Width Type Velocity Power(W) rate (g/s) (mm) (mm) (mm/s) Factorial 600 6 0.07 0.393 1.604 Factorial 600 6 0.35 1.627 1.974 Factorial 600 14 0.07 0.169 1.338 Factorial 600 14 0.35 0.816 1.612 Axial 700 10 0.21 0.654 1.455 Centre 800 10 0.21 0.696 1.528 Axial 800 12 0.21 0.692 1.527 Centre 800 10 0.21 0.704 1.535 Axial 800 10 0.14 0.554 1.502 Centre 800 10 0.21 0.630 1.504 Centre 800 10 0.21 0.684 1.538 Centre 800 10 0.21 0.719 1.531 Axial 800 10 0.28 0.789 1.627 Centre 800 10 0.21 0.655 1.581 Axial 800 8 0.21 0.877 1.733 Axial 900 10 0.21 0.719 1.558 Factorial 1000 6 0.07 0.508 1.787 Factorial 1000 14 0.07 0.25 1.586 Factorial 1000 6 0.35 1.769 2.412 Factorial 1000 14 0.35 0.846 1.819

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Chapter 4 Study and prediction of laser metal deposition with various structures

Figure 4.28 Effect of factors on deposition height

Figure 4.29 Effect of factors on deposition width Table 4.6 Analysis of variance table (ANOVA)

Sum of Mean F p-value

Source Squares df Square Value Prob > F Model 2.66 9 0.3 76.43 < 0.0001 significant A-Laser Power 0.02 1 0.02 5.13 0.0469 B-Process Velocity 0.63 1 0.63 162.38 < 0.0001 Width C-Powder Feed rate 1.75 1 1.75 452.94 < 0.0001 AB 2.66E-03 1 2.66E-03 0.69 0.4255 AC 7.20E-05 1 7.20E-05 0.019 0.8941 BC 0.2 1 0.2 50.75 < 0.0001 Model 0.99 9 0.11 63.8 < 0.0001 significant A-Laser Power 0.15 1 0.15 86.8 < 0.0001 B-Process velocity 0.27 1 0.27 158.79 < 0.0001 C-PowderVelocity Feed rate 0.29 1 0.29 167.12 < 0.0001 Height AB 3.44E-03 1 3.44E-03 2 0.1878 AC 5.73E-03 1 5.73E-03 3.32 0.0983 BC 0.03 1 0.03 17.28 0.002

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Analysis of variance (ANOVA) is a statistical method used to test differences between two or more means. The ANOVA table illustrates the analysis of parameter factors on deposition width and height, where F value indicates the test for comparing the source’s mean square to the residual mean square and p-value (Prob > F) describes the probability of seeing the observed F value if the null hypothesis is true (there are no factor effects). If the Prob>F value is very small (less than 0.05 by default) then the source has tested significant.

For width, the model F-value of 76.43 implied that the model was significant. Values of “Prob > F” less than 0.0500 indicated that model terms were significant. In this case, laser power, process velocity and powder feed rate were all significant model terms, among which, powder feed rate played the most important role followed by process velocity and laser power. For the case of deposition height, model F-value of 63.8 implied that the model was significant. Similar conclusions could be drawn that powder feed rate played as the most obvious influential factor on layer height, followed by process velocity and laser power. In summary, the coefficient for powder feed rate was the highest in both width and height equations indicating that powder feed rate was the most influential factor in comparison with process velocity and laser power. Moreover, laser power played the weakest part in determining the layer height and width which was in agreement with the previous experiments.

4.8 Multiple layer deposition

For single layer deposition, it can be concluded from the literature review presented in Chapter 2 that both powder feed rate and process velocity play significant roles in single layer cladding height, width and dilution. Therefore, process velocity and powder mass flow rate should be taken into consideration when determining the final deposition geometry and mechanical property during multiple layer deposition. The following experiments were implemented using maximum laser power 1kW in order to achieve a high deposition rate. Six tracks were deposited with 35% overlap side by side. Five more layers were added on top of the previous layer. The working stage was manually adjusted after deposition of each layer to keep the standoff distance constant.

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Figure 4.30 Cross-sections of 5 samples Table 4.7 shows the dimensions of deposition applying different process velocity and powder feed rate.

Table 4.7 Deposited dimension by applying different process parameters

Process velocity Powder feedrate Sample Power (W) Height (mm) Width (mm) (mm/s) (kg/hour) 1 1000 14 0.50 2.336 5.627 2 1000 16 0.75 2.945 5.820 3 1000 18 1.00 3.608 5.537 4 1000 20 1.00 3.298 5.614 5 1000 22 1.25 3.529 5.506

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Chapter 4 Study and prediction of laser metal deposition with various structures

No obvious crack or porosity can be seen from Figure 4.30. The directional growth was pointed from the bottom towards the surface of deposition attributing to the direction of thermal gradient. Grain size between layers was refined because of the re-melting effect.

The powder delivered from powder feeder could not be utilized completely. Thus, actual deposition rate should be calculated by using the equation 4.4.

푽풐풍풖풎풆 ×푫풆풏풔풊풕풚 푹 = (4.4) 푻풊풎풆

Volume could be regarded as bulk of the deposited material which in this case equaled to the product of width, height and length. Time could be calculated by adding up the time required for each layer. The density of melted 316L powder is approximately 6.9g/cm3.

Volume×Density 5.627×2.336×40×6.9 Sample1: R = = 40 = 35.26 g/s = 0.127 kg/h Time 6× ×6 14

Volume×Density 5.820×2.945×40×6.9 Sample2: R = = 40 = 52.58 g/s = 0.189 kg/h Time 6× ×6 16

Volume×Density 5.537×3.608×40×6.9 Sample3: R = = 40 = 68.93 g/s = 0.248 kg/h Time 6× ×6 18

Volume×Density 5.614×3.298×40×6.9 Sample4: R = = 40 = 71.00 g/s = 0.256 kg/h Time 6× ×6 20

푉표푙푢푚푒×퐷푒푛푠푖푡푦 5.506×3.529×40×6.9 Sample5: 푅 = = 40 = 81.90 g/s = 0.295 kg/h 푇푖푚푒 6× ×6 22

The actual deposition rate ranged from 0.127 kg/h to 0.295 kg/h. The utilization of the powder for 5 samples are 25.4%, 25.2%, 24.8%, 25.6% and 23.6% respectively. Some of the unused powder might bounce away from the substrate or fly off track after leaving the nozzle tip. Considering the interaction between laser beam and powder stream during flight, thermal property and geometry of the unused power might differ, and hence unutilized powder could not be collected and recycled.

4.9 Deposition pattern

Different methods were applied to study the effect of deposition pattern on final deposition quality. The stage was controlled by a 2-phase step motor manufactured by ISEL Automation International [230]. Deposition patterns were firstly programmed

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Chapter 4 Study and prediction of laser metal deposition with various structures manually and typed into a computer connected to the numerically controlled stage. Process velocity, powder feed rate and laser power were chosen to be 18 mm/s, 1 kg/h and 1 kW. English letters L, C and P were printed by using different deposition methods. L and C were printed in six layers width and ten layers height. P was printed in 7 layers width and ten layers height due to the difficulty in programming deposition path in a closed area continuously.

Four methods were used to print C illustrated in Figure 4.31. First it was built by connecting two separate deposited walls as indicated in Figure 4.32 (A). After the first segment was printed, working stage moved to next deposition point without laser energy input. The overlap around the connective region was also settled to be 35%. It is apparent from Figure 4.32 (A) that the irregular convex could be observed on the joint part, on account of the working stage accelerating from zero moving velocity. B, C and D illustrated in Figure 4.32 represented raster, offset out and offset in deposition pattern respectively. Steps on the surface could be seen on sample B and C, for the reason that adjacent deposited layers could not be re-melted perfectly and some layers would be squeezed up forming unevenness on the deposition surface. For deposition method D, the laser was moving from inside to outside reducing the surface unevenness arising from overlap effect. However, oxidation on the previous layer could still influence surface roughness of the following layer. Although surface roughness could be improved by applying different deposition method, convex on the corner still existed for all of the samples. More experiments were conducted to validate the result by printing letter C and P which are presented in Figures 4.33 and 4.34.

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A B

C D

Figure 4.31 Four methods to print L (A) Two segments, (B) Raster, (C) offset in, (D) offset out

Figure 4.32 Four deposited letter L by using different deposition methods

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A B

Figure 4.33 Two methods to print C, (A) raster (B) offset out

Figure 4.34 Path design for letter P and deposited sample 139

Chapter 4 Study and prediction of laser metal deposition with various structures

4.10 Re-melting

Laser re-melting is a process using the laser beam to melt the sample surface without adding any feeding material in order to improve the surface properties. Defects including oxidation, porosity, sulphide and metal compounds can generally be found between coarse dendrites of the deposition. Properties like fatigue strength, corrosion resistance and wear resistance may be weakened when these defects exist on the surface of the samples. Impurities, porosities and compound can be released by using the laser re- melting technique. Refined grain size can be achieved in the meantime because of the rapid cooling rate. Song et al. [231] concluded from their work that a large amount of porosities could be eliminated by applying laser surface re-melting and good metallurgical bond between laser re-melted layer and substrate could be achieved.

In this case, laser re-melting was applied to improve the surface roughness of the deposition surface. Same deposition path and process parameters were conducted for this sample with the sample shown in Figure 4.31 (D). Laser re-melting process was implemented between each layer to reduce the accumulation of surface unevenness. Less oxidation could be observed on the surface after re-melting. Additionally, metallic lustre and smoother surface could be obtained in comparison with the process without re- melting. However, convex on the corner arising from the change of process direction could not be avoided. Moreover, sidewall steps still existed both with or without re- melting process which is shown in Figure 4.35 (B).

A B

Figure 4.35 (A) Deposition of re-melted sample L (B) Comparison of the side wall between re-melted sample and Figure 4.31 (D)

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4.11 Modelling of deposition with multiple adjacent passes

Figure 4.36 Laser deposition process of five adjacent passes

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Chapter 4 Study and prediction of laser metal deposition with various structures

Figure 4.36 presents the laser deposition process of five adjacent passes, where the development of bead morphology, melt pool-velocity field and temperature distribution during the process are illustrated. Overlap ratio between adjacent passes was set to be 35% and 1 s idle time was applied to simulate the required moment for the stage to move between passes. It could be observed that melt pool became more elongated with more passes being processed which was due to the heat accumulation on the substrate and formed deposition. Melt pool transformed from symmetrical teardrop shape to asymmetrical tilted shape after the first pass. The laser would re-melt some portion of the last deposited pass and a titled melt pool slope would appear between the substrate and deposition during the process. Temperature distribution at the end of each pass indicated that the deposition functioned as a heat sink where heat could not dissipate quickly enough, and hence temperature of the substrate would increase after each pass.

Figure 4.37 Melt pool development history during depositing five adjacent passes Figure 4.37 describes the melt pool development during the five-pass deposition process, where the red line represents the melt pool area on the deposition surface while melt pool volume (blue line) represents the total molten material. Melt pool area was highly dependent on the laser spot size and hence not too much change occurred after five passes. However, because of the accumulated heat effect, heat was conducted to the substrate contributing to a rapid increase in total melt pool volume, especially after the third pass. It could be noticed that during the first pass, there was a drop in cell number on both melt pool area and total melt pool volume. This was caused by the bulge effect and a bump could normally be seen at the beginning of deposition. It will be discussed in the next chapter.

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4.12 Modelling of re-melting process

As can be concluded from the literature review in Chapter 1, most of the previous research on modelling of re-melting process mainly focused on the melt pool development and temperature distribution history. Comparision of stress status including analysis of deformation after re-melting has been widely conducted by utilising FEM modelling which is based on the analysis of temperature development. Experiments were performed by many other people reporting the effect of re-melting on improving the surface quality, microhardness and other surface properties. However, few have attempted to simulate the re-melting process on already deposited geometries. This may be attributed to the complicated melting and solidification process which is difficult to model, especially when the free surface between air and deposition is required to be tracked.

Figure 4.38 Side view of the deposition formation

Figure 4.39 Re-melting process with excessive heat input The re-melting process, which took into consideration the gravity effect, free surface tracking, melt pool formation and fluid flow, was simulated using this model. Improvement of the surface roughness and effect of process parameter were successfully

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Chapter 4 Study and prediction of laser metal deposition with various structures simulated. A side view of the deposition development during the process of five adjacent passes is shown in Figure 4.38. It can be noticed that when depositing a new layer, some portion of the previous layer would be re-melted and mix with the newly added material and solidify again. Owing to the surface tension, the cross-section of the molten bead appeared to be sphere shape with some certain contact angle with the substrate. Therefore, step effect on the surface was inevitable since there was always some portion of the previous layer could not be melted. It can be observed from Figure 4.38 that a slope existed on the surface when depositing the second pass next to the first one. Despite the fact that some unevenness between passes could still be seen on the surface, a relatively stable surface finish was achieved after the fourth pass was deposited. Re-melting was applied after deposition with heat input but without adding material and reduced laser power was generally used for re-melting. Excessive heat input would result in a larger melt pool because of heat accumulation and there being no added material to melt. Surface tension would no longer be able to hold the molten material, and the deposition would collapse due to the gravity effect shown in Figure 4.39. Figure 4.40 presents the comparison of surface quality before and after re-melting when proper process parameters were applied.

Figure 4.40 Surface finish comparison before and after re-melting process

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4.13 Modelling of deposition with complex shapes

Figure 4.41 Deposition of letters ‘LPRC’ and temperature distribution

Figure 4.42 Deposition of letters ‘LPRC’ with cooling between features 145

Chapter 4 Study and prediction of laser metal deposition with various structures

Deposition of letters ‘LPRC’ was modelled and the resultant bead morphology and temperature distribution are shown in Figure 4.41. Some convex shape with increased deposition height could be seen around the corner due to the longer period laser would dwell around the corner area, which brought about increased heat input and more material to be added to the melt pool. Some pre-designed geometries would disappear when the ratio between required deposition feature and laser spot size was not big enough. As shown in Figure 4.41, the circle feature of letter “P and R” could not be seen clearly because the heat input was concentrated in that region with continuous change of laser process direction. The small ratio resulted in a rapid increase of temperature and expanding melt pool size. Figure 4.42 presents the deposition with reduced laser power and cooling around the circle area, and the final deposition quality was improved.

4.14 Modelling of geometries with overhang structures

From the literature review, most of the previous modelling research on laser deposition process focused on discussing the melt pool development on flat operating directions, where the effect of gravity has always been ignored.

111.6º

100.5º

Figure 4.43 Deposition of second and third overhanging layers

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However, in real experimental environment gravity effects is a significant role in determining the deposition quality, especially when overhang structures are required or the process is implemented in non-flat operating directions. Therefore, getting a better understanding of the influence of gravity effect on deposition quality is necessary in real experimental environment.

67.5º 73.4º

97.7º 95.3º

Figure 4.44 Multilayer deposition with overhanging structures

Figure 4.45 Multilayer deposition with complex overhanging structures The current model was capable of building complex geometries with overhanging structures, which took into account the effect of gravity, surface tension, mass addition 147

Chapter 4 Study and prediction of laser metal deposition with various structures and other thermal processes during the deposition. The influence brought by the gravity effect could be seen from the modelling results varying with different depositing conditions. Real three-dimensional deposition geometries could be successfully built.

Figure 4.43 presents the second and third deposited overhanging layers. 0.3 mm offset on laser beam centre was applied between two layers to achieve deposition with 111.6º and 100.5º inclination angle after depositing the second and third layer. The relation between surface tension and gravity of the molten material should be balanced, where the excessive amount of molten material would cause collapsing of the final deposition when surface tension could no longer hold the total mass. Figure 4.44 illustrates the deposition of two circles with inward (73.4º inclination angle) and outward (97.7º inclination angle) offset distances. The collapsing problem was more likely to happen with an outward offset if no supporting structure was applied during the deposition process. Figure 4.45 shows the deposition with some complex overhanging structures. During the simulation, gravity direction was set to be perpendicular to the laser beam direction meaning that the laser was titled in order to keep a perpendicular position with the substrate. In this case, collapsing was believed to happen with more layers to be deposited without support. Another way to produce this geometry is to change the sample position by rotating the product after depositing the two-layer circle, where the operating gravitational direction is kept constant.

4.15 Conclusion

This chapter first presented the experimental details of coaxial powder laser deposition process. Influence of process parameters on final deposition quality was investigated. Within some certain process windows, layer height, width and dilution would all increase with the increase of laser power. When increasing the powder feedrate, higher and wider deposition could be achieved with decreased bead dilution. Layer height, width would both decrease when the process velocity was increased, but not much change could be observed for the dilution. The influential factor level of each process parameter was also discussed and powder feedrate was found to be the most influential one determining the deposition dimensions, under the condition that enough laser power was provided to fully melt the powder material.

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Then a three-dimensional model was built to simulate the single layer deposition, where different process parameters were applied during the modelling. The effect of thermal behaviour at the beginning of deposition was found to have a significant impact on the melt pool formation history, and hence affect the final deposition quality. The simulated deposition morphology and dimensions of the single bead were compared with the experimental results and a good agreement was achieved.

Multiple layer deposition with various process parameters was successfully performed without the appearance of cracking and porosities. By applying the optimised process parameters, letters shaped “L” “P” “C” were deposited with different deposition patterns. It was observed that the deposition pattern would have a substantial influence on the final surface roughness. Re-melting process was applied to improve the surface quality.

A three-dimensional model was built to simulate the laser deposition process with multiple adjacent passes. The temperature distribution and melt pool formation could be explained by tracking the melt pool surface area and total melt pool volume, where the heat accumulation effect on the formation of the deposition could be understood. The influence of existing layer on the new developing melt pool could also be observed from the model.

Modelling of re-melting process on improving the actual deposition surface roughness was presented and selection of the right re-melting parameters was found essential to achieve an improved surface quality. Deposition of letters “LPRC” with complicated features was successfully simulated and it was found that the spot size would play a significant role during the real deposition process.

Last, 3D geometries with overhanging features were successfully simulated for the first time and the effect of gravity could be observed from the modelling result.

With the help of numerical simulation, gravity direction relative to the operating environment could be easily altered to investigate the possibility of depositing geometries with some certain overhanging structures before experiments. The effect of gravity on melt pool formation and final deposition quality will be discussed specifically in the next chapter.

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5.1 Introduction

The literature review and earlier discussion indicated that although plenty of work could be found on modelling the laser metal deposition process, the gravity effect on the molten material was always neglected, especially when the deposition process was conducted in non-flat direction with additional filler material feeding into the melt pool. The effect of gravity should be coupled with fluid dynamics and melting & solidification process inside the melt pool in order to investigate the influence gravitational force could bring. All the complicated mechanisms included during the melt pool formation would bring many difficulties for developing such a numerical model.

This chapter first presented the modelling strategy of a three-dimensional model, followed by the description of experimental setup equipped with a high-speed camera system. The effect of gravity on deposition formation in various positions was successfully demonstrated and the influence of different process parameters was discussed. Melt pool formation history and temperature development were explained to help understand how gravity would affect the fluid flowing conditions inside the melt pool. A gap also exists from the previous work that a full understanding of the unevenness at both ends of the deposition is missing. This model successfully simulated the appearance of bulge at the beginning and the collapsing at the end. The system instability brought by the wire feeder and robot arm was also taken into consideration and added to the model. The gravity effect on unevenness at both ends was investigated. As an important factor in the melt pool, the influence of surface tension-temperature coefficient on melt pool formation in non-flat position, was discussed and how this coefficient could affect the unevenness at both ends of the deposition was also investigated. For potential space applications, gravity was gradually decreased from 1 g to zero g including the moon gravity 0.16 g to investigate the effect of reduced gravity on the final deposition. Process parameters were optimised to minimise the effect.

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5.2 Modelling strategy

The modelling strategy is similar to the principles which have been presented in Chapter 4 for laser metal deposition process using powder, except that the mass addition is in the form of filler wire and the momentum brought by the wire is also considered. The dimensions of the whole calculation domain were 6 mm × 30 mm × 6 mm and the mesh was generated using GAMBIT which contained 370440 (42×210×42) hexahedral cells with uniform grid spacing 142 µm shown in Figure 5.1. The calculation time step was 푢∆푡 3×10-4 s and tested by the Courant number which is defined as 퐶 = , where 푢 is the ∆푥 magnitude of the velocity, ∆푡 is the time step and ∆푥 is the grid size. Fluid cannot travel more than one cell in one time step meanning that the courant number cannot bigger than one. It took approximately 24 hours to simulate a single deposition line which would run 2.2 s in real time. With time step bigger than 3×10-4 s, more iterations would be required to reach the residual criteria during every calculation time step and hence the total modelling time would be longer. Grid and time-step independence test were carried out by reducing the grid spacing to 100 µm with 1080000 (60×300×60) hexahedral cells and decreasing the time step to 1.5×10-4 s. Slight difference in simulation result was obtained with the finer grid and/or smaller time step, while it would take much longer time to complete simulating the same deposition length.

6mm

Figure 5.1 Mesh for the calculation domain A 2 mm thickness 316L substrate region was initially patched for the domain, on top of which air region of 4 mm thickness was created shown in Figure 5.2. PISO algorithm was used as the pressure-velocity coupling solution. The velocity and temperature fields were discretized with a second order upwind scheme, and the pressure field was discretized

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Chapter 5 CFD modelling of laser metal deposition process and effect of gravity with a PRESTO! scheme. The convergence criteria for residuals of continuity and momentum equation was 10-3 and 10-6 for energy equation. Considering that the whole deposition process required to be modelled, a transient time solution with pressure-based solver was used. VOF model combined with Geo-Reconstruct method was applied to track the interface between base material and gas phase, where user-defined mass, momentum and energy source terms were added. Steps consisted within every calculating iteration are illustrated below:

(1) Fluid properties (surface tension coefficient, density, viscosity, specific heat, heat capacity) were updated according to the current temperature and flowing time (2) Momentum equations were solved as a result of user-defined momentum sources terms and mass source terms (gravity, buoyancy, wire feeding velocity, surface tension pressure, Marangoni stress, wire feeding material). (3) Mass, pressure and velocity fields were updated using the pressure correction equations to calculate the fluid flow in melt pool. (4) Energy equations including moving Gaussian heat source and thermal heat source were solved. (5) Free surface deformation was updated by solving the interactions between gas and metal phases.

(6) Check if the equations were converged: if converged, update the boundary condition and go to the next time step; if not, return to step (1).

outflow (side and top surfaces of the upper domain)

wall (side and bottom surfaces of the lower domain)

Figure 5.2 Boundary conditions and initial phase patch of the calculation domain 152

Chapter 5 CFD modelling of laser metal deposition process and effect of gravity

(outflow)

(outflow) (outflow)

(wall) (wall)

(wall)

Figure 5.3 Illustration of boundary conditions on cross-section plan The top and side surfaces of the upper domain were gas (air) phase boundaries treated as outflow shown in Figures 5.2 and 5.3. The bottom and side surfaces of the lower domain were substrate phase boundaries defined as wall conditions. The calculation domain was chosen as part of the whole workpiece, and hence heat conduction and heat radiation were applied to the wall boundaries. Heat conduction, heat radiation and heat convection were applied on the free interface between air and substrate phases, and the free interface was tracked by the VOF method. The laser beam was treated as a moving volumetric heat source (W/m3) added on the free interface and the position of laser beam centre at a given time was defined as a function of the process velocity. Mass and momentum sources brought by the wire feeding were applied to the free interface. The mass source can be expressed as:

휌∗휋푟2 ∗푓 푀푆 = 푤푖푟푒 푤푖푟푒 (5.1) 푛∗푉푐푒푙푙

3 where 휌푤푖푟푒 (kg/m ) is the density of wire material, 푟푤푖푟푒 (m) is the radius of wire and 2 2 푓푤푖푟푒 (m/s) is wire feed rate. Momentum source (kg/m ·s ) of the wire in the z-direction can be determined as follows:

2 2 휌∗휋푟푤푖푟푒∗푓푤푖푟푒∗푠푖푛훼 푀표푚푧 = − (5.2) 푛∗푉푐푒푙푙 while momentum source in the y-direction can be expressed as:

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2 휌∗휋푟푤푖푟푒∗푓푤푖푟푒 푀표푚푦 = ∗ (푓푤푖푟푒 ∗ 푐표푠훼 − 푣푙푎푠푒푟) (5.3) 푛∗푉푐푒푙푙 where α is the angle between wire feeder and sample surface, and 푣푙푎푠푒푟 (m/s) is laser process velocity. Solidification and melting process was taken into consideration with enthalpy-porosity method by applying a damping force to the mushy material to decrease the fluid velocity. Gravity force was added to the cells with temperature higher than the melting point. The Boussinesq approximation was used to present the buoyancy force to improve the calculation speed.

Table 5.1 Thermos-physical properties of 316L [114]

Symbol Nomenclature Value Units ρ Liquidus density 7600 kg/m3 푇푠 Solidus temperature 1670 K 푇푙 Liquidus temperature 1730 K 8 -1 ΔH0 Standard heat of adsorption -1.88x10 J kg mol μ Dynamic viscosity 0.005 kg∙m-1 ∙K-1 k Thermal conductivity 24 W∙m-1∙K-1 L Latent heat of fusion 2.47x105 J∙kg-1 -1 -1 퐶푝 Specific heat 700 J∙kg ∙K β Thermal expansion 4.95x10-5 K-1 -2 -1 ℎ푐표푛 Heat convection coefficient 25 W∙m ∙K σ Stefan-Boltzmann constant 5.67x108 W∙m-2∙K-4 ε Surface radiation emissivity 0.4 푓훼 Sulfur concentration in substrate 100 ppm R Universal gas constant 8314.3 J∙kg-1 ∙mol-1 η Laser beam efficiency 60 %

Thermos-physical properties of 316L for the simulation are listed in Table 5.1. In order to investigate the effect of process parameters on deposition result in different conditions, seventeen groups of cases were created for the simulation. Cases 1-5 were for the bigger laser spot size Ø3 mm with 3 kW laser power. For a smaller laser spot size Ø2 mm (cases 6-10), power was reduced to 1.3 kW to keep the power density constant as 4.2×104 W/cm2. To achieve the same material/laser spot ratio with case 3, wire feed rate was reduced to 0.009 m/s for the smaller spot size in case 11. The same value of the wire feed rate was applied for the bigger laser spot size in case 12. Energy input was set to be 3 kW for smaller spot size in case 13 while in case 14 wire feed rate was increased to 0.045 m/s for bigger spot size to achieve the same material/laser spot ratio with case 8. Laser process velocity was increased to 0.015 m/s in case 15. To investigate the role of surface tension

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Chapter 5 CFD modelling of laser metal deposition process and effect of gravity during the deposition process, comparative studies with positive and mixed surface tension-temperature gradient were applied in case 16 and 17 respectively.

Table 5.2 Process parameters in different cases for the simulation and experiments

Operating Laser spot Process Laser Wire feed rate position size velocity power (m/s) (mm) (m/s) (kW) Case 1 flat 3 0.01 3 0.02 Case 2 vertical down 3 0.01 3 0.02 Case 3 vertical up 3 0.01 3 0.02 Case 4 overhead 3 0.01 3 0.02 Case 5 horizontal 3 0.01 3 0.02 Case 6 flat 2 0.01 1.3 0.02 Case 7 vertical down 2 0.01 1.3 0.02 Case 8 vertical up 2 0.01 1.3 0.02 Case 9 overhead 2 0.01 1.3 0.02 Case 10 horizontal 2 0.01 1.3 0.02 Case 11 vertical up 2 0.01 1.3 0.009 Case 12 vertical up 3 0.01 3 0.009 Case 13 vertical up 2 0.01 3 0.02 Case 14 vertical up 3 0.01 3 0.045 Case 15 vertical up 2 0.015 1.5 0.02 Case 16 vertical up 2 0.015 1.5 0.02 (positive) Case 17 vertical up 2 0.015 1.5 0.02 (mixed)

5.3 Experimental procedures

A continuous wave (CW) ytterbium fibre laser (IPG YLR-16000) was used in the direct laser deposition experiments with a maximum laser power of 16 kW available and delivered with an optical fibre of 300 μm in diameter. The laser beam emitted from the optical fibre was firstly collimated with a lens with a 150 mm focal length and then focused with a 400 mm focal length. The beam parameter product (BPP) of the laser was 10 mm.mrad with a 0.8 mm focus spot diameter and 15 mm Rayleigh length. The Precitec laser processing head was mounted on a 6-axis KUKA robot.

A high-speed digital video recording system was applied during the experiments to monitor the whole deposition process without the interference of the high brightness laser light. The system included a Cavitar CAVILUX HF diode laser as the illumination light source and a high-speed CCD camera system fitted with a bandpass optical filter. The

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CAVILUX laser shown in Figure 5.4 is a 500 W diode laser modulated for up to 200 kHz at an 810 nm wavelength. The camera was attached on the KUKA robot moving together with the laser head and the camera was aimed at 45º aligning with the laser spot centreline.

The material of the substrate and wire were both 316L austenitic stainless steel. The dimensions of the substrate were 100 mm × 100 mm × 10 mm and the diameter of the wire was 1.2 mm. Argon shielding gas was delivered by two tubes attached on both sides of the substrate in order to reduce oxidation during the high-temperature process. For experiments performed in the vertical position, the laser head together with the wire feeder and work stage rotated to a position which is shown in Figure 5.5. The experimental set-up equipped with a high-speed camera system is presented in Figure 5.6.

b g a Laser unit High speed unit Cavitar laser camera illumition

Control unit Illumination optics

Figure 5.4(a) Cavitar laser illumination system (b) High-speed camera in horizontal position

Figure 5.5 Schematic diagrams of wire laser deposition process (a) flat (b) vertical

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Figure 5.6 Experimental set-up equipped with high-speed camera system Four adjacent tracks side by side with two layers were deposited and the result is shown in Figure 5.7 (a) (b). Process parameters applied for the experiments were: 3 mm diameter spot size, 3 kW laser power, 10 mm/s process velocity, 0.02 m/s wire feed rate and 40% overlap between passes.

Scanning direction b

g c g

Figure 5.7 2-layer deposition result and cross-section (a) top view, (b) experimental cross section, (c) cross-section of modelling result

Figure 5.7 (c) presents the simulated result of the two-layer deposition and some good agreement can be reached when compared with the experimental result in Figure 5.7 (b).

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5.4 Results and discussion

5.4.1 Deposition morphology

Figure 5.8 shows the comparison of experimental results with two diameters in the vertical up position for case 3 and case 8. Humps, irregular and thicker depositions and larger deposition slope angles at the end can be observed in Figure 5.8 (b) when a small spot size was used.

Figures 5.9 and 5.10 present the simulated results in different operating conditions with two spot sizes 3 mm and 2 mm. The laser power density was kept constant as 4.138×104 W/cm2 and wire feed rate 0.02 m/s was the same for these two groups. Since the deposition width was directly related to the laser spot size, and the total material added was the same, the resultant deposition was thicker for cases with a smaller beam spot size.

The simulated results have a good qualitative agreement with the experimental results showing that gravity has a distinct effect on the final deposition geometry, in particular for the cases with smaller spot sizes. The starting sections of deposited areas were slightly higher in case 1 compared with case 2. The same phenomenon could also be observed for cases 6 and 7 where the difference became more evident. Interestingly, deposition height appeared to be more uniform at the beginning in the vertical down position (cases 2 and 7) when gravity was in the same direction with laser process direction. The unevenness at the beginning of deposition is known as bulge effect and will be discussed in the later sections. In the overhead operating position, when a larger beam size was applied not too much difference of final bead geometry could be noticed between case 1 and case 4. However, gravity would have more influence when a smaller spot size was used in case 9 where significant unevenness appeared. For cases implemented in the horizontal positions, bead unevenness was not apparent for both small and large spot sizes since the deposition growing direction was perpendicular to the gravitational direction.

a b g v g v

Figure 5.8 Side views of depositions in the vertical up position with a laser spot size of (a) Ø3 mm [case 3], (b) Ø2 mm [case8]

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Case1

g v

Case2

g v

Case3

g v

Case4

g v

Case5

v ⊙ g

Figure 5.9 Deposition in flat, vertical down, vertical up, overhead and horizontal positions with Ø3 mm, P=3 kW, v=0.01 mm/s, f=0.02 m/s (2 layers)

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Case6

g v

Case7

g v

Case8

g v

Case9 v g

Case10

⊙ g v

Figure 5.10 Deposition in flat, vertical down, vertical up, overhead and horizontal positions with Ø2 mm, P=1.3 kW, v=0.01 mm/s, f=0.02 m/s (2 layers) 160

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Case11

g v

Case12

g v

Figure 5.11 Deposition in vertical up position (3 layers), v=0.01 mm/s, f=0.009 m/s, Case 11: Ø2 mm, P=1.3 kW Case 12: Ø3 mm, P=3 kW

Case13

g v

Case14

g v

Figure 5.12 Deposition in vertical up position, Case13: Ø2 mm, P=3 kW, v=0.01 mm/s, f=0.02 m/s (2 layers) Case 14: Ø3 mm, P=4 kW, v=0.01 mm/s, f=0.045 m/s (1 layer) For the cases deposited in the vertical up positions, the final bead geometry tended to show more irregularity. Different process parameters were applied in order to obtain a better understanding of the waviness formation. There were more humps for case 8 in comparison with case 3 as a result of smaller laser spot size, where power density and material feeding rate were kept the same. Relatively uniform deposition heights could be

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Chapter 5 CFD modelling of laser metal deposition process and effect of gravity achieved even after three layers when reduced wire feed rate was used. However, the bead height might become too small due to the low material addition rate (cases 11 and 12) which would result in lower deposition efficiency shown in Figure 5.11. By increasing the laser power input, deposition could be transformed from three humps in case 8 to two humps in case 13. Most of the distinct humps appeared during the deposition of the second layer, by increasing the wire feed rate and providing enough heat input, waviness could be seen after just one layer which is shown in case 14. There was no change in the number of humps between case 8 and case 15 when the laser process velocity was increased, despite the fact that deposition height was smaller with less total material fed into the melt pool. In conclusion, the unevenness at the beginning of deposition and the periodic humps seem to have a significant influence on the final surface finish. Therefore, further investigation is required to fully understand the reason for these two phenomena.

5.4.2 Melt pool development

a b c d e

g j f h i Figure 5.13 Temperature and melt pool-velocity field history for case 8, (a&f:0.36 s, b&g:1.44 s, c&h:1.80 s, d&i:1.908 s, e&j:2.196 s)

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Figure 5.13 illustrates the development of different stages of hump formation for case 8, where a-e present the temperature distribution history and f-j show the melt pool morphology and velocity field correspondently. Since solidification rate along centreline of the melt pool was the highest while temperature gradient on the solidification boundary was the biggest, with a relatively low laser process velocity the melt pool would become elliptical shape [120]. If the process velocity was high, latent heat liberated by the solidification process could not dissipate quickly around the melt pool centre as a result of the low thermal gradient. Therefore, the melt pool would become elongated and form a shape similar to the teardrop, where the maximum solidification rate was lower than the laser moving velocity.

For deposition in the vertical up position, at the beginning with the laser energy input and wire material addition some part of the existing deposition would be re-melted and form a melt pool which could contribute to an elliptical-shaped melt pool on the surface shown in Figure 5.13(f). Due to the effect of gravity, the molten material would flow towards the gravitational direction enlarging the bottom of the melt pool seen in Figure 5.13(g). The melt pool would have the potential to transform into teardrop shape when the generated fusion heat could not be transferred quickly enough which can be observed in Figure 5.13(h). This would happen if the rear material was still in liquid phase but started to get far away from the laser beam centre. Finally, the two liquid regions would lose the connected flowing channel and form into two separate melt pools which is illustrated in Figure 5.13(i). The rear liquid material which was a mixture of the newly added wire and previously deposited material would solidify and form a droplet-shaped deposition on the bead as a result of surface tension.

When starting to deposit a new layer on the existing layer after 1 s cooling time between two passes, the average temperature of the deposition was approximately 480 K. The material already deposited from the previous layer was firstly re-melted and would flow downwards joining the wire which was newly added and melted. On the other hand, with the same laser heat input, a higher maximum temperature could be obtained during the second layer deposition due to the fact that the temperature gradient was smaller, and less heat could be conducted to the substrate, which in turn would result in a bigger melt pool on the surface. These two factors could explain the reason why the dripping phenomenon happened on the second layer deposition while the first layer was still relatively flat for case 8. Once the deposition process went through the beginning region where bulge effect

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Chapter 5 CFD modelling of laser metal deposition process and effect of gravity happened, the melt pool would experience a cycle which included teardrop shape to elliptical shape to enlarged rear to two separate melt pool regions. Therefore, the waviness on the bead would appear after a certain interval of running time.

Since the evolution of melt pool was essential during the laser metal deposition process, a new method was developed in this investigation to help gain a better understanding of the melt pool development. Melt pool surface area and total melt pool volume were extracted from numerical results in the form of cell number. Liquid cell number is defined as the number of cells which are in the molten state and contain liquid fraction of metal. The development history of liquid cell number on melt pool surface (ls) and total liquid number (lv) are presented in Figure 5.14 to Figure 5.16. To make the presentation clear, melt pool development of cases with bigger and smaller spot size are firstly illustrated in Figure 5.14 and Figure 5.15 respectively. The combined comparison of cases 1-10 is then shown in Figure 5.16 with both big and small spot sizes.

Figure 5.14 melt pool evolution for cases with big spot size

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Figure 5.15 melt pool evolution for cases with small spot size

Figure 5.16 Comparison of melt pool evolution for cases with big and small spot size At the beginning of the first layer, melt pool started to grow when laser energy impacted on the substrate. The total heat input could be transferred towards several aspects: heat reflected by substrate and wire, heating up newly added wire material, heat conducted to the substrate, heat radiation to surrounding environment, heat convection between flowing protection gas and substrate surface, heat convection induced by Marangoni stress and latent heat during phase transformation. As can be seen from the Figure 5.14, melt pool condition began to stay relatively stable after around 0.5 s when the thermal balance among these aspects was reached. Cases with smaller spot size seemed to enter the stable state slightly faster than cases with bigger spot size. Melt pool area on the surface and total melt pool volume started to decrease after 2.2 s when the laser was

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Chapter 5 CFD modelling of laser metal deposition process and effect of gravity stopped. It took around 0.17 s and 0.23 s for small and big melt pool respectively to completely solidify. For cases 1-5 with higher power and bigger laser spot size, melt pool surface area and total melt pool volume were bigger than cases 6-10 of which the laser power and laser diameter were smaller. The intensity of a Gaussian laser beam smoothly decays from its maximum on the beam axis to zero, while a top hat laser beam has an intensity profile which is flat over most of the covered area. The melt pool diameter was roughly the same as the laser spot size when Gaussian distributed laser beam was used. Therefore, melt pool area on the surface was larger for the cases with bigger laser spot size. Higher laser power was applied for the larger diameter, keeping the power density the same and most of the heat would be conducted to the material substrate which as a result would lead to higher melt pool volume. The difference between melt pool areas for different laser beam diameters was small while a big gap occurred for the melt pool volumes which could also be attributed to the high total heat input. On the melt pool surface, Marangoni stress induced by surface tension-temperature gradient played an important role in determining the pool geometry. For cases with same spot size but different operating positions the difference between melt pool areas was small. This was due to the fact that the same laser energy impacted on the substrate surface would contribute to a similar temperature distribution and hence lead to a similar temperature gradient. Similar convection flow caused by temperature gradient would result in a similar melt pool area on the surface regardless of the operating positions.

For the total melt pool volume during the first layer, there lied some difference between cases with same spot size but performed in various positions. Vertical down position had the highest melt pool volume while vertical up had the lowest melt pool volume for both bigger and smaller spot size. Due to the fact that gravity direction was the same as the deposition direction in vertical down position, the newly added wire was firstly melted and would have the potential to flow together with some re-molten substrate material towards the deposition direction. Since there was heat input along the deposition direction, it would take longer time for the molten material to solidify which would contribute to a higher total melt pool volume when compared with vertical up and horizontal positions. On the contrary, for vertical up position molten material would have the potential to flow against the deposition direction towards a cooler area or a region which was already solidified. This could lead to a faster solidification and relatively smaller total melt pool volume. It can also be observed from Figure 5.14 that with a bigger spot size, the

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After the first layer was deposited, 1 s interval was applied to simulate the idle time for the robot arm to move between layers. During the second layer, melt pool surface area and total melt pool volume both increased in comparison with the melt pool condition during the first layer. Only a small portion of the heat was transferred through radiation and convection towards the surrounding environment. Most of the heat input could only be conducted to the previously deposited layer which was a two-dimensional process instead of three-dimensional. In addition, the temperature of deposited material was still higher than room temperature as a result of the short inter-pass time. Due to the fact that melt pool size was mainly determined by the effective laser spot size which was the same as the first pass, hence only a small difference could be observed between first pass and second pass on the melt pool surface area because of the heat accumulation. Instead of being stable along the time, melt pool surface area started to fluctuate during the second layer. The accumulated heat could not dissipate quickly from the deposited layer to the substrate and this, in turn, would contribute to a much larger total melt pool volume.

What stands out during the second time section in Figure 5.15 is the rapid increase, decrease and waviness in total melt pool volume for most of the cases, especially when smaller spot size was used. Unlike the development history in the first layer, a dramatic drop in total melt pool volume occurred around 0.8 s after the second layer deposition started. The steep decrease at the beginning can be related to the bulge formation and will be discussed in next section. The maximal and minimal melt pool volumes differed among cases with different process conditions. Apart from some irregularity in vertical up position (case 3), the final deposition results were relatively smooth for the cases with big spot size. This could also be confirmed by the melt pool development history. After

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Chapter 5 CFD modelling of laser metal deposition process and effect of gravity reaching the maximum at the beginning, melt pool volume decreased and tended to stay stable in case 1,2,4,5. On the contrary, undulation could be observed from all the cases 6- 10 when smaller spot size was used.

When performed in vertical up position, the molten material would flow downwards as a result of gravity and would accumulate at the end of deposition making the end bigger. The rear region of the melt pool would expand and the melt pool volume increased to the highest. The melt pool was elongated with material constantly flowing downwards before laser centre moved far away from the rear region of the melt pool. The liquid channel between front melt pool and rear region would become narrower and narrower. Two separate molten regions would finally occur and the total melt pool volume would start to decrease. Melt pool volume would reach the lowest when the rear molten region solidified completely, after which the melt pool shape would start to transfer from teardrop shape to elliptical-shape increasing the total melt pool volume and another cycle would begin. This phenomenon would happen periodically which could explain the roller-coaster shaped history of melt pool volume during vertical up process in cases 3 and 8.

For the depositions processed in the vertical down position, the history of the melt pool volume stayed steady in total for both spot size conditions because the molten material would always flow in the same direction with the direction of laser processing and hence a quick stable state without further fluctuation could be reached. By comparing case 4 and case 9, a conclusion could be reached that for the overhead deposition position, gravity would have more influence on the final morphology when a smaller spot size was used. Two humps could be observed on the final deposition of case 9 which could be related to the two heaves on the melt pool development history. Except for the small decrease at the beginning, no oscillation appeared in case 4 and the resultant deposition was relatively flat. The balance between surface tension and gravity was more likely to be disrupted when the aspect ratio between melt pool width and melt pool height was low. When the ratio became higher, surface tension could not hold the molten material due to gravity force and as a result unevenness would occur. Cases performed in flat and horizontal positions seemed to have similar melt pool development history during the first two layers with slightly higher melt pool volume on the flat position. With more layers to be built, the tendency of collapsing would happen on horizontal position if no supporting material was applied.

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5.4.3 Temperature development

Five points were allocated on the substrate surface to track the temperature history of the deposition process as shown in Figure 5.15. P1 was located in the centre of substrate surface and the interval between points was set be 0.5 mm in the x-direction.

P2 P1 P4 P3 P5

Figure 5.17 Five allocated points on the substrate surface

Figure 5.18 Temperature history of 4 selected points for Case 6 Temperature stayed room temperature 300 K until the point entered the zone affected by the moving laser beam. Heat was firstly transferred by thermal conduction to the middle point P1 before laser beam centre reached P1. Temperature increased rapidly when laser hit the point. According to the Gaussian distribution law, P2 which was 0.5 mm away from the centre would get only 47.23% of the maximum laser power. However, due to the fact that the spot size was small and the temperature difference on the melt pool surface was high, a strong convection flow would occur within the melt pool. As a result, P1 and P2 would go through a similar temperature history. For P3 which was on the edge

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Chapter 5 CFD modelling of laser metal deposition process and effect of gravity of the effective laser beam, temperature increased slowly below the melting point, where heat transferred by convection and conduction could not provide enough energy for P3 to melt. For P4 and P5 with further distances from the laser beam centre, temperature rise was caused by thermal conduction. Temperature started to decrease with the laser beam moving away from these points. During the temperature cycle of the second layer, melting point was not reached for all five locations due to the fact that most of the laser energy was absorbed by the first deposited layer to form a melt pool and melt the newly added wire material. Therefore, temperature underwent a similar history as the first layer with lower temperature. Since the substrate was already heated during the first layer, temperature at the end of process was higher than room temperature at the starting point.

Figure 5.19 Temperature history of 4 selected points for Case 1 For case 1 with a larger laser diameter, due to the fact that three of the allocated points were located within the melt pool area and would reach the melting temperature, P1 P2 P3 experienced a similar temperature history as a result of strong convection flow in the melt pool. There was a strong temperature increase for P1 when the laser beam centre passed the middle point. During the second layer, heat from the deposition could not penetrate deeply and quickly enough to melt the points on the substrate which led to a similar temperature history but with a lower peak value in comparison with the first layer.

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Figure 5.20 Temperature history of middle point P1 for different cases Figure 5.20 presents temperature history of the middle point (P1) for different cases. Because of the higher total energy input, overall temperature of the cases with bigger spot size and laser power were higher than the cases with smaller spot size. However, not much difference could be observed among cases with same spot size but performed in different positions, which indicated that gravity factor would have limited effect on temperature development under the conditions using the same process parameters.

5.5 Surface unevenness at both ends

5.5.1 Bulge formation

Unlike the vertical up and vertical down process where gravity was parallel to the deposition direction, and the overhead process where gravity had direct influence on the melt pool formation without any support from the substrate, during depositing in the flat position gravity should have less effect on the deformation of melt pool because gravity was perpendicular to the laser beam moving towards the substrate direction and the previously deposited layer would function as a support for the melt pool. Therefore, in principle, the resultant history for melt pool volume on the flat position should be stable. However, a dramatic change of the melt pool status could still be observed on the flat position when small spot size was used. This rather interesting finding could be attributed to the appearance of the bulge at the beginning of deposition which is shown in Figure 5.21.

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bulge

Figure 5.21 Bulge at the beginning of case 4 (second layer, 0.18 s) It could be explained by the fact that temperature would increase rapidly at the beginning of the process when laser hit on the substrate and melted the material. As a result, the temperature gradient in the melt pool would become greater which at the same time increase the surface tension force. Since the surface tension force was proportional to the difference of the free surface energy between substrate and melt pool, a spherically shaped deposition was formed to reduce the surface tension. Substrate temperature would rise soon after the laser beam started to move which would contribute to a smaller thermal gradient. Therefore, the bulge effect would disappear once the process became stable.

On the other hand, robot arm acceleration & deceleration was also part of the reason for the happening unevenness. Due to the smaller average process velocity at the starting and ending of the deposition process, more heat input together with more filler material would be added to the melt pool, which in turn resulted in the bulge appearance. No report yet has been found on discussing these combined effects especially when the deposition process is demonstrated in various operating directions.

It is not realistic in the actual experimental environment to neglect the influence of time interval with discontinuous velocity at the beginning. However, it can be easily simulated without acceleration and deceleration by numerical method. The significance of these two factors can be investigated separately. In this work, firstly no acceleration or deceleration was applied during the calculation.

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5.5.2 Effect of process conditions on bulge formation

Interestingly, the bulge effect would still appear especially for the case on flat position with small spot size. Deposition geometries at 0.42 s after the second layer started were extracted to show the bulge formation for different cases. Projection of the final bulge morphology was also presented and compared in Figure 5.22.

The bulge effect seemed to have a huge impact on melt pool development at the beginning of the process. Some difference could be noticed between cases with small and big laser beam size on flat position (case 1 and case 6). No staircase effect appeared in case 1 and the contact angle of the second layer was smaller than 90o, while for case 6 with smaller spot size, the staircase effect could be noticed and the contact angle was bigger than 90º. The maximal width of case 1 was 3.59 mm 20% larger than the spot diameter 3 mm while for case 6 the maximal width was 2.8 mm which was 40% bigger than the spot diameter. Due to the fact that same mass feed rate was applied on Case 1 and 6, a narrower but higher deposition was obtained for Case 6. As a result of the surface tension force, a spherically shaped deposition was formed to reduce the surface area. With a higher temperature gradient and bigger material-width ratio, bulge tended to appear more distinct for the case with small spot size. The difference of bulge formation could also be verified from the melt pool history where there was a big drop of the total melt pool volume for case 6 while it was relatively stable for case 1.

By comparing the bulge shape among cases 6-10 with same process parameters but in different operating directions, it could be observed that the gravity force had an impact on the bulge morphology. Bulge at the beginning tended to shift towards the gravitational direction and dimension of the bulge was also influenced by gravity. Vertical up position had the biggest bulge width and height, followed by overhead position, while vertical down position had the smallest bulge size. Despite the fact that bulge size of vertical up and overhead was similar, the exact location of the bulge was different. Bulge in case 8 was closer to the deposition end. From the melt pool development history in Figure 5.15, the maximal melt pool volume of vertical up and overhead position was almost the same and higher than rest of the cases in the group, which contributed to bigger bulge sizes. A prominent bulge formed at the beginning would lead to a more fluctuant melt pool volume history before the process became stable. Although the bulge was slightly asymmetric due to the gravity effect in case 10, similar melt pool history between horizontal and flat

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Case1

3.59mm

g v 1

.98mm

Figure 5.22 Bulge morphology at 0.42 s and projection of the bulge morphology at 2.2s

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According to the discussion above, a conclusion could be reached that the bulge dimension was related to the process parameters and operating directions. Different process conditions would result in different temperature development history at the deposition beginning. As a result, surface tension variation which was dependent on the temperature gradient played an important role during the bulge formation.

5.5.3 Effect of system instability on bulge formation

Apart from this, some system instability in real experimental conditions can also cause the unevenness at the beginning. Due to the fact that the laser head and wire feeder were attached on a KUKA robot arm, there was a short time interval of acceleration at the beginning and deceleration before the robot arm stopped or changed the deposition direction. This time interval was taken into consideration as shown in Figure 5.23 below.

Acceleration and deceleration of the scanning 0.012 0.01 0.008 0.006 0.004 0.002

Scanning velocity (m/s) 0 0 0.5 1 1.5 2 2.5 3 Time (s)

Figure 5.23 Acceleration and deceleration of the robot arm at starting and ending of the deposition process

2.97 mm 1.16 mm a

2.80 mm 0.76 mm b

Figure 5.24 Deposition with (a) and without (b) consideration of system instability 175

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The resultant deposition is presented in Figure 5.24 and is compared with the deposition without acceleration and deceleration at both ends. Deposition in Figure 5.24 (a) had a higher maximum height 2.99 mm and lower minimal height 1.83 mm compared with Figure 5.24 (b) which had the value of 2.77 mm and 2.01 mm, making the height difference 1.16 mm versus 0.76 mm. Diameters of the bulge were 2.97 mm and 2.80 mm which were both larger than the laser spot size 2 mm. Therefore, it could be concluded that the effect of robot acceleration at the beginning would enlarge the bulge and add to the irregularity of final deposition in the aspect of bead width and height. After melting and solidification of the process became stable, a slope existed between already deposited bead and melt pool on the front. This would result in a less uniform height at the end of deposition. The slope effect would accumulate when more layers were deposited and would lead to the collapse of deposition if it could not be compensated somehow. However, generally it is not an issue in real experimental conditions because deceleration of the robot always exists. Due to the velocity decrease at the end of process, more material would be fed into the melt pool and this would contribute to some height compensation for the slope around the end region. As can been observed from Figure 5.24, the slope angle was bigger for the case with velocity deceleration and it brought more uniform deposition quality at the end.

Figure 5.25 Melt pool and velocity field with (a) and without (b) consideration of system instability in the middle of second layer deposition (1.2 s and 1.1 s)

The history of melt pool formation could also be used to help explain the influence brought by system instability. Figure 5.25 presents the melt pool and velocity field for both cases in the middle of the deposition. The melt pool was more elongated in Figure 5.25(a) due to the fact that more total heat was delivered into the deposition at the beginning of the process with a lower average velocity. Considering that the surface

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Chapter 5 CFD modelling of laser metal deposition process and effect of gravity tension-temperature gradient was negative in these two cases, convection flow induced by Marangoni stress on the melt pool surface would move from higher temperature area to lower temperature area. Therefore, the molten material tended to flow towards the rear region of the melt pool making the pool elongated.

Figure 5.26 Comparison of melt pool development between with and without consideration of system instability

Figure 5.26 explains the development of melt pool surface and total melt pool volume during two layers’ deposition. For the case under system instability condition, it took extra 0.2 s at both ends for the robot to accelerate and decelerate, which explained the time shifting between the trend lines. When the first layer was deposited, not too much difference existed between the two cases. For the green line, maximal melt pool volume was higher at the beginning because of the acceleration effect while at the end there was a small rise due to the deceleration. The difference in melt pool volume became bigger during the second layer. The green line had a higher maximal and a lower minimal value in comparison with the blue line. The greater difference between maximum and minimum could also be related to the height difference in Figure 5.24. Because of the longer melt pool, more molten material would be pushed back towards the rear region of the melt pool making the bulge more evident.

The reason for the unevenness at the beginning and end of deposition is not always the same. Unfortunately, all of the literature failed to fully explain the mechanism or in other words explain the reasons separately. Firstly, if system instability was not considered, the deposition appeared to be higher at the beginning and lower at the end. Next, the effect of system instability was added to the melt pool development history. At the beginning

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Chapter 5 CFD modelling of laser metal deposition process and effect of gravity of the deposition, the unevenness was a combined result of bulge effect caused by surface tension variation and material accumulation because of lower process velocity during acceleration. At the end of the deposition, melt pool condition played an important role in determining the steady slope angle with the deposited bead. Robot deceleration would then influence the final deposition geometry. If the velocity decreased too rapidly, the height difference could be compensated in a relatively small scale. On the other hand, if the velocity decreased too slowly, too much material would be added to the melt pool and this would result in an increased height of the deposition at the end.

However, all the discussion made above was based on the assumption that the material addition and heat input were completely synchronized. In real deposition conditions, the time difference between material adding and heat input could sometimes lead to a lower deposition height at the beginning, if the heat was added far ahead of the material addition.

From the experiments, single bead with multiple layers was built and the results are shown in Figure 5.27. It could be seen that both the starting and ending part of the bead collapsed and became more and more serious after few layers deposition. One reason firstly came to mind was the acceleration and deceleration of the robot arm at both ends which would contribute to a more concentrated energy input. From the video, it could be noticed that another reason might be some existing synchronous lag between laser and wire feeder, which meant that the wire and laser could not start or end simultaneously. Therefore, the starting and ending part of the bead actually underwent a re-melting process without any wire material fed into the melt pool.

a start end

b start end

Figure 5.27 Top view (a) and side view (b) of 7-layer deposition

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Figure 5.28 shows some images obtained from the high-speed camera which illustrate the relation between the wire and melt pool at both ends of the deposition during the process. It can be observed from Figure 5.28(a) that the laser beam began to melt the deposited layer before wire coming out of the nozzle at the staring part of the process. Wire started to retreat before the laser beam actually stopped operating on the deposition at the end of the process which can be seen from Figure 5.28(b).

Figure 5.28 Image from the high-speed camera (a) Starting part (b) Ending Part

Figure 5.29 Modelling of deposition without synchronized heat and mass feed Figure 5.29 shows the simulated deposition result assuming that at the beginning material was fed into the melt pool 0.2 s after starting the laser, while at the end of deposition material feeding stopped 0.2 s before stopping the laser. The simulated bead shape had a good agreement with the experimental result.

This model gives a better understanding of the development of deposition unevenness, especially at both ends. It can help predict the final bead geometry covering the factor of surface tension variation and system instability.

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5.6 Effect of surface tension coefficient

5.6.1 Effect of surface tension coefficient on non-flat deposition

Surface tension plays a significant role during the development of melt pool. Apart from surface tension-temperature coefficient, same process parameters were applied in cases 15, 16 and 17 to investigate the effect of surface tension on deposition formation in non- flat positions shown in Table 5.3. Negative, positive and mixed surface tension- temperature coefficient was applied to case 15,16 and 17 respectively.

Table 5.3 Cases with different surface tension-temperature coefficient

Operating Laser spot Process Laser Wire feed rate position size velocity power (m/s) (mm) (m/s) (kW) Case 15 vertical up 2 0.015 1.5 0.02 (negative) Case 16 vertical up 2 0.015 1.5 0.02 (positive) Case 17 vertical up 2 0.015 1.5 0.02 (mixed)

Figure 5.30 presents the bead geometry for the three cases. In comparison with case 15, the deposition of case 16 and 17 was more shifted towards the gravity direction while the difference between positive and mixed case was not obvious. Positive surface tension coefficient tended to make the melt pool more spherical with a bigger difference between maximal and minimal width value. This could be explained by the fact that molten material on the surface would flow from solidification boundary towards the melt pool centre when surface tension coefficient was positive. The material distribution would be more concentrated in the melt pool centre and hence gravity would have more effect on the melt pool formation. A new balance between surface tension and gravity would be reached resulting in the deposition with more evident spherical irregularity.

Despite the fact that the laser power input was the same for three cases, the inward convection flow would slow down the heat transfer speed because the flow direction was against the temperature gradient, which would contribute to a higher maximal temperature inside the melt pool as shown in Figure 5.32. Higher temperature inside the melt pool would lead to more material to be melted.

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Figure 5.31 illustrates the comparison of melt pool development. In flat deposition position, negative surface tension coefficient would contribute to a melt pool with wider surface area but shallower depth. On the contrary, the positive coefficient value would result in a narrower but deeper melt pool. This principle seemed still working during the first layer. Although the difference was not too evident, it could still be observed that the positive case had higher total melt pool volume but lower surface area in comparison with the case under negative condition. However, during the process of the second layer, both melt pool area and melt pool volume were higher in the case of positive coefficient. Due to the heat accumulation and lower heat dissipation rate, more material was melted resulting in a higher melt pool volume. The volume difference between positive and negative also became higher. Surprisingly, the rapid increase of temperature during the second layer brought more influence to the positive case since it would become even harder to transfer the heat away due to the inward flowing fluid. As a result, the melt pool surface expanded rapidly and turned out to be larger than that of the case with a negative coefficient. This could also be observed during the deposition in the flat position and has been analysed in the bulge formation section.

For positive surface tension coefficient, the maximum surface tension value located in the melt pool centre where temperature was the highest. For the cases with negative surface tension coefficient, the maximum surface tension value lied on the solidification boundary where the temperature of molten material was the lowest. Unlike positive and negative situations, for mixed surface tension coefficient, the maximal surface tension value lied between the solid-liquid interface and melt pool centre where the surface tension coefficient transferred from positive to negative. Therefore, the molten material tended to flow from melt pool centre and solidification boundary towards the surface tension transferring region. The collided stream would contribute to a deeper melt pool and higher melt pool volume. This explained the reason why melt pool volume was the highest for the case with mixed surface tension coefficient during the first layer. When the second layer was deposited, the accumulated heat would contribute to a higher melt pool volume and reduce the volume difference between negative and positive cases.

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Case15

g v

Case16

g v

Case17

g v

Figure 5.30 Deposition in vertical up position with negative, positive and mixed ∂γ/∂T, Ø2 mm, P=1.3 kW, v=0.015 mm/s, f=0.02 m/s

Figure 5.31 Comparison of melt pool development for different surface tension coefficient

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Figure 5.32 Comparison of temperature history for different surface tension coefficient Owing to the different fluid flowing patterns within the melt pool, deposition with positive surface tension coefficient had the highest peak temperature, followed by mixed and negative. On the other hand, the peak temperature duration was the longest for mixed case because the temperature was more evenly distributed in the melt pool as a result of the mixing flowing.

Figure 5.33 Comparison of melt pool development for different surface tension coefficient with lower process velocities With lower process velocities, both the melt pool area and melt pool volume would increase and surprisingly the influence brought by surface tension coefficient decreased. As shown in Figure 5.33, not too much difference existed in total melt pool volume, especially during the second layer. Due to the dramatic increase of liquid volume as a result of higher energy input, margaroni stress induced convection flow seemed to have

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Chapter 5 CFD modelling of laser metal deposition process and effect of gravity less impact on the final deposition formation. On the contrary, the balance between gravity force and surface tension normal to the interface seemed to play more significant role in determining the unevenness on non-flat deposition positions.

5.6.2 Effect of surface tension coefficient on bulge formation

Different surface tension-temperature coefficient was applied to investigate how it would affect the bulge formation during the process. Figure 5.34 presents the final deposition geometry and bulge dimensions for cases with different surface tension coefficient. Bulge 휕훾 size in negative case was the least obvious one with smallest maximal deposition 휕푇 width and height. Although the total deposition dimension was similar for positive and 휕훾 mixed cases, more distinct bulge seemed to occur in the mixed situation with a bigger 휕푇 difference between maximal and minimal bead height at the beginning.

2.76 mm

0.761 mm 2.78

2.84 mm

0.864 mm 2.

90 90

2.84 mm

0.995 mm 2.

92 92

Figure 5.34 Surface tension on bulge formation (a) negative (b) positive (c) mixed After reaching steady state, the deposition slope angle relative to the substrate also varied 휕훾 among different cases, where positive coefficient had the biggest slope angle, 휕푇 휕훾 followed by cases with mixed and negative value. Bigger slope angle would indicate 휕푇

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Chapter 5 CFD modelling of laser metal deposition process and effect of gravity more uniform bead shape at the end and less chance to have the collapsing problem during the development of deposition with multiple layers, where the unevenness at the end would accumulate with more layers to be built.

휕훾 Temperature development history of middle point P1 with different is shown in Figure 휕푇 5.35. Apparently, the case with positive coefficient value had the highest peak temperature since the molten material would flow from solidification boundary to melt pool centre making the heat more concentrated in the middle. Due to the multi-directional 휕훾 flow generated in mixed case, the temperature was more evenly distributed inside the 휕푇 휕훾 melt pool and therefore the cooling rate for mixed case was the lowest. 휕푇

Figure 5.35 Temperature history of middle point P1 with different surface tension coefficient

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a b

c d

e f

Figure 5.36 Melt pool and velocity field for different surface tension coefficient (a) negative 4.3 s, (b) negative 5.4 s, (c) positive 4.3 s, (d) positive 5.4 s, (e) mixed 4.3 s, (f) mixed 5.4 s

휕훾 Figure 5.36 shows the comparison of melt pool and velocity field for different 휕푇 situations in the middle (4.3 s) and end (5.4 s) of the second layer deposition. Melt pool development history is presented in Figure 5.37.

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Figure 5.37 Melt pool development of cases with different surface tension coefficient Due to the fact that the melt pool area was mainly dependent on the laser spot size, not too much difference could be noticed between the cell numbers which contained liquid material on the surface. Since the liquid material contained fully molten cells and mushy area where material was between liquids and solidus temperature, it could be found that 휕훾 the case with negative value had bigger fully molten pool (red area in the liquid 휕푇 fraction graph) and less mushy zone. With negative surface tension coefficient, molten liquid tended to flow from centre to the melt pool edge making the pool wider and shallower. A big difference existed on melt pool volume between negative and other two cases at the beginning when the bulge was formed. The higher total melt pool volume contributed to bigger bulge dimension for positive and mixed cases. For mixed surface tension coefficient, molten flow moved from both the pool centre and solidification edge 휕훾 to the transferring point where changed from positive to negative. Given the situation 휕푇 that temperature was much higher during the second layer because of the heat accumulation, this multi-direction flow would lead to slightly higher maximal value in both melt pool area and total melt pool volume when compared with the other two cases. 휕훾 It could also be observed that the minimal melt pool volume in mixed case was lower 휕푇 than the case with positive coefficient. As a result, more distinct bulge appeared at the beginning when mixed surface tension coefficient was applied.

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5.7 Effect of gravity levels and zero gravity

As can be concluded from the discussion above that the balance between surface tension and gravity of liquid metal will determine the morphology of the deposition result in different process directions. In the potential space applications for component repair, with the absence or reduction of gravitational level, surface tension forces would become more important and would determine the system behaviour. However, until now there has been no studies reported on the influence of surface tension only on LMD process without the effect of gravity. In space, gravity is not available as a control variable and hence the process parameters suitable for LMD on the Earth should be optimised for the adaptation and transition to the space. With the help of current model, gravitational level can be reduced and reduced to zero to investigate the melt pool behaviour and deposition process for space applications. One of the benefits that can be brought by developing the AM process in space is the possibility of using the resources on planetary surfaces, which can help cut down the shipment of bulk materials from the Earth during launch. The Moon can be a perfect place to collect resources and manufacture some useful components. The acceleration due to gravity on the Moon surface is about 1.625 m/s2 which is approximately 0.16 ɡ [232] when compared with the gravity level on the Earth.

Gravity was decreased gradually from 1 g to zero g including the moon gravity 0.16 g to investigate the effect of reduced gravity on final deposition. An extra case with increased gravity 2 g was also considered for the possible overweight situations. The results of the first deposited layers are presented in Figure 5.38 and the results for the second deposited layers with different gravity values are compared in Figure 5.39.

Figure 5.38 Comparison of the first deposited layer under different gravity levels (Ø=2 mm, P=1.3 kW, v=0.01 mm/s, f=0.02 m/s)

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Figure 5.39 Comparison of the second deposited layer under different gravity levels (Ø=2 mm, P=1.3 kW, v=0.01 mm/s, f=0.02 m/s) As can be observed from Figure 5.38 that no clear variation among cases with different gravity values could be noticed after the first layer was deposited. This is because during the first layer, the substrate was still cold and as a result melt pools of small sizes would be formed. This can also be verified from Figure 5.40 that the total melt pool area & volume during the first layer were small and melt pool development history was stable for different gravity values. With the same heat input during the second layer, melt pool size would increase because of the heat accumulation. At the same time, surface tension in the normal direction of melt pool surface would tend to hold the liquid material as a sphere to reduce the total surface area.

Figure 5.40 Melt pool evolution for cases with different gravity values The bulge formation process could explain the fluctuation of melt pool development at the beginning of deposition during the second layer. It could also be noticed that the

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Chapter 5 CFD modelling of laser metal deposition process and effect of gravity maximum melt pool volume varied when different gravity value was applied. In flat deposition position, gravity would function as a pressing force and help surface tension hold the metal liquid better. Therefore, bigger bulges were more likely to occur with reduced or zero gravity when compared with normal or increased gravitational level.

Aspect ratio defined as width/height was reported by Pinkerton and Li [233] to be an important quality criterion to determine the appearance of porosities between two adjacent overlapping layers. Abioye et al. [234] concluded from their work that the aspect ratio would increase when increasing the laser power and process velocity while decreasing the wire feed rate. They pointed out that the aspect ratio was highly related to the wire volume deposited per unit length of track. With the decrease of wire feed rate or increase of the process velocity, wire volume deposited per unit length would drop and hence reducing the deposition height. In the meantime, deposition width was invariant with the wire volume deposition rate and as a result the ratio would become bigger. On the other hand, when increasing the laser power, the bead would be flattened enlarging the melt pool size and higher aspect ratio would be obtained with an increased layer width and decreased deposition height.

Contact angle [235] was also related to the wire volume deposited per unit length of track. Contact angle tended to become bigger with an increased wire feed rate but decreased laser power and process velocity. Because of the surface tension, spherical shaped deposition with smaller contact angle would be formed when applying lower wire feed rate and/or higher process velocity. Contact angle would also decrease if more laser power was added into the process. This is because that the melt pool width would become larger as a result of increased heat input and the liquid metal would spread towards the melt pool edge contributing to tracks with decreased height and smaller contact angles.

Figure 5.41Comparison of deposition cross-section with different gravity values As can be observed from Figure 5.39 that with the reduction of gravity value from 2g to zero g, the deposition unevenness appeared to be more prominent after experiencing the bulge formation at the process beginning. Figure 5.41 shows the comparison of the bead

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Chapter 5 CFD modelling of laser metal deposition process and effect of gravity cross-section after depositing two layers under different gravity conditions, from which width, height and contact angle of the bead were measured and aspect ratio was calculated. It is obvious that there is a dramatic increase in deposition height when decreasing the gravity value. Figure 5.42 also indicates that reduced gravity environment would result in higher deposition contact angle and lower aspect ratio. The explanation could be that in flat deposition position, the component force of gravity would help surface tension better hold the liquid metal. Without the effect of gravity, the liquid metal would have the potential to form a sphere with increased contact angle under the impact of low aspect ratio and high melt pool volume & surface area, and hence the irregularity of the deposition would become more distinct.

100 1.6 99 1.4

) 98 1.2 ° 97 1 96 0.8

95 0.6 Aspect ratio

Contact angle ( 94 0.4 93 0.2 92 0 0 0.5 1 1.5 2 contact angle Gravity value (x g) aspect ratio

Figure 5.42 Contact angle and aspect ratio with different gravity values In order to achieve better deposition surface quality under reduced or zero gravity situations, aspect ratio should be increased to produce tracks with smaller contact angle so that melt pool surface area would be lessened and the effect of surface tension force normal to the melt pool surface could be reduced. Considering that the bead irregularity happened during depositing the second layer, optimization of process parameters was conducted from the second pass to improve surface quality under zero gravity circumstances. Figure 5.43 presents the comparative results of deposition using different process parameters. Laser power was firstly increased from 1.3 kW to 1.6 kW and the results are shown in Figure 5.43 (c) and Figure 5.43 (d). It can be noticed by comparing Figure 5.43 (b) and Figure 5.43 (d) that although the deposition finish was improved by adding more heat input, the overall height was slightly decreased, given the total amount of wire fed into the melt pool was the same. The surface unevenness was improved and further improved by increasing the process velocity from 0.01 m/s to 0.015 m/s and 0.02 191

Chapter 5 CFD modelling of laser metal deposition process and effect of gravity m/s while keeping laser power and wire feed rate constant. The results are presented in Figure 5.43 (e)(f)(g)(h). While reducing the wire feed rate from 0.02 m/s to 0.01 m/s, a third layer was deposited on top of the second one to obtain the same wire feeding amount.

Figure 5.43 Comparison of deposition with different process parameters

Figure 5.44 Comparison of deposition cross-section and melt pool with different process parameters (a,g): P=1.3 kW, v=0.01 mm/s, f=0.02 m/s, (b,h): increased laser power P=1.6 kW (c,i): increased velocity v=0.015 mm/s, (d,j): further increased velocity v=0.02 mm/s, (e,k): reduced wire federate f=0.01 m/s, (f,l): reduced wire federate f=0.01 m/s and power P=1 kW Figure 5.44 compares the deposition cross section after optimizing the process parameters. As can be seen from Figure 5.44 (a) and Figure 5.44 (b) that the contact angle was reduced

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Chapter 5 CFD modelling of laser metal deposition process and effect of gravity by 20.1º after increasing the laser power. There was a noticeable drop in deposition height while the bead width was dramatically increased which was even wider than the first deposited layer. Enlarged melt pool size would result in a more flattened layer with decreased deposition height and smaller contact angle. Contact angle was decreased from 99.3 º to 84.4º and 79.5º when the process velocity was increased from 0.01m/s to 0.015 m/s and 0.02 m/s. Despite the fact that better surface finish could be obtained by increasing the process velocity, the increment of layer height was decreased with reduced wire volume deposited per unit length of track. Smaller contact angle of 79.4º was achieved by halving the wire feed rate from 0.02 m/s to 0.01 m/s. The melt pool during the third pass would become bigger due to the heat accumulation and this would result in the reduction of deposition height. A comparative case was conducted by applying a reduced laser power from 1.3 kW to 1 kW for the third layer to keep the melt pool size small. The results presented in Figure 5.43 (k,l) and Figure 5.44 (f,l) indicate that the contact angle was further reduced to 78.7 º with significantly decreased surface irregularity.

5.8 Conclusion

Gravity has a distinct effect on the final deposition result when the process is performed in non-flat directions. Melt pool development history and temperature field were utilized to help understand the deposition process. The periodical appearance of the dripping would occur more often in vertical up and overhead operating directions when gravity and movement of laser beam were not in the same direction where there was less support for the melt pool. The dripping shaped deposition would occur as a result of gravity effect on the formation of the melt pool. Gravitational force tended to enlarge the rear region of the melt pool where the molten material would accumulate until the laser beam moved far away from this region, and the channel between the rear and front region of the melt pool would break and another cycle would begin.

With the same material feeding rate and given the condition that enough heat was provided to fully melt all the added material, the process performed with bigger spot size would have less bead shape irregularity comparing with smaller spot size. This may attribute to the fact that a bigger ratio between molten material height and width would result in a more prominent bead angle where gravity would have more effect on the

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Chapter 5 CFD modelling of laser metal deposition process and effect of gravity molten material and the balance between surface tension and gravity was more likely to be disrupted.

The bulge at the beginning of deposition would occur as a result of high temperature gradient when the laser started to impact on the cold substrate surface, where a spherically shaped deposition would form to reduce the surface tension. Bigger melt pool height/width ratio would contribute to a bigger bulge size at the beginning. Gravity seemed to have an impact on the bulge formation, wherein vertical down position the gravity would decrease the bulge size while the bulge size would be enlarged in vertical up and overhead positions.

With the effect of surface tension, the deposition tended to appear bigger at the beginning and a slope would occur at the end. In the real experimental environment, apart from the factor caused by thermal behaviour, system instability including robot acceleration & deceleration, the synchronous issue between laser and wire feeder would bring more irregularity to the final bead deposition. This model was capable of simulating the whole deposition process taking into account the thermal behaviour and system instability.

The effect of gravity on deposition result in non-flat process positions would also vary with different surface tension coefficient. Different flowing patterns inside the melt pool owing to different relationship between surface tension and temperature would bring about a different balance condition of surface tension and gravitational force. With a bigger melt pool, normal surface tension force rather than tangential surface tension stress would have more impact on the overall melt pool formation.

For potential space applications, the deposition unevenness appeared to be more prominent with the reduction of gravity value. This may be attributed to the fact that the component force of gravity would help surface tension better hold the liquid metal. To improve the surface irregularity under low or zero gravity conditions, wire volume deposited per unit length of track should be decreased by reducing the wire feed rate or/and increasing the process velocity, which could contribute to a lower contact angle and higher aspect ratio. Deposition quality could also be improved by increasing the laser power which would also bring a decrease in contact angle. However, increased heat input could lead to an enlarged melt pool size with higher liquid metal volume, which in turn might cause a reduction in layer height and deposition collapsing.

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6.1 Introduction

Welding is a process where two or more of the same or dissimilar materials are joined together by bonding and diffusion between atoms or molecules. In certain industrial applications, structures with large dimensions would be difficult to be built with a single additive manufacturing process due to the limitation of the equipment. Therefore, narrow gap welding could be applied to join the separate additive manufactured samples to help achieve the overall process target.

In this chapter, a brief overview was first presented on the current research status of the welding process. Effects of gravity on melt pool formation in non-flat operating positions and process performed in various gravitational levels were discussed. Some previous research on narrow gap welding and dissimilar welding process were presented. A scientific gap was found from the review that most of the modelling work focused on the analysis of temperature distribution history and thermal induced residual stresses. The dynamic flow during the melt pool formation in narrow gap process has always been neglected. Some phenomena brought by gravity effects have not been fully explained.

Therefore, a three-dimensional model was built for the first time for narrow gap welding process. Lack of fusion could be observed from the model and process parameters were optimised to eliminate this phenomenon. Effects of gravity on weld bead formation in various welding positions were presented. Influence of gravitational level and surface tension on weld pool formation were discussed, after which modelling of thick section multi-pass narrow gap welding process was presented. Finally, the laser narrow gap dissimilar welding process was modelled by including four phases and the influence of re-melting and gravity effect were also discussed.

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6.2 Previous work

6.2.1 Effect of gravity on welding process

Different gravitational levels were created and applied to the welding process to investigate the melt pool behaviour difference. Eisazadeh et al. [236] discussed the effect of gravity on microstructure, melt pool formation and mechanical properties during gas tungsten arc (GTA) welding 316L samples, where a multi-gravity research welding system was used to create an environment with 3.6 g. They concluded that the increased gravitational environment would contribute to a stronger convection flow induced by the buoyancy force, which in turn would result in a wider but shallower melt pool. Comparison of tensile test between 1 g and 3.6 g samples indicated that the increased gravitational condition would slightly lower the tensile ductility. The multi-gravity research welding system was also utilized by Gandhi and Aidun [237] to investigate the effect of gravity acceleration level on microstructure formation during dissimilar welding process. They reported that the size of the unmixed zone could be reduced by applying an increased gravity level, where the convection flow induced by the buoyancy force was enhanced. They also observed that the stronger flow inside the melt pool also increased the cooling rate around the welding region which would contribute to a decreased size of heat affect zone (HAZ). Aidun et al. [238] concluded from their experimental results that with an increased gravity level, the ratio between melt pool depth and width would decrease, where buoyancy force would become more important in determining the fluid flow than surface tension inside the melt pool. Wang and Tandon [239] carried laser welding process in a reduced gravitational environment during parabolic flights. A wider and deeper melt pool was achieved under the reduced gravity condition when compared with melt pool dimension obtained on the ground using the same process parameters. Defects including porosities, slag and some un-melted particles were found in the welding which they assumed would affect the actual mechanical properties in real space applications.

Apart from applying different gravity levels, the effect of gravity when the welding process was performed in various operating directions was reported by many previous studies. Sun et al. [240] conducted some experiments on investigating the laser welding thick section process on horizontal positions. A U-shaped groove was machined on the

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Chapter 6 CFD modelling of narrow gap welding based additive manufacturing process sample top surface in order to increase the welding penetration and reduce the weld width. Mechanical properties were found to be different along the penetration line with slightly high tensile strength close to the surface. Shen et al. [241] compared the processing window and porosity distribution during laser welding in flat and horizontal positions. They concluded that the processing window for horizontal welding was much wider, where undercut on the top and sagging on the bottom of the welds were less likely to happen with full penetration of the weld. Higher specific (1.3 times) energy was required for the horizontal welding position to achieve the same penetration depth with the flat position. Most of the porosities were found on the upper part of the welds in horizontal welding position, while for the flat position, the majority of the porosities were located around the centre region of the weld. Chang et al. [242] investigated the effect of gravity on final laser welding quality of Ti6Al4V material on flat and horizontal positions. They found that the undercut for both positions was larger on the top than on the bottom surface. They discovered that there were fewer porosities for the welds processed in flat position than horizontal position because there were more routes for the porosities to escape in the flat position. They concluded that in total flat position would have a better welding quality regarding the weld formation, number of porosities and mechanical properties. Chang et al. [243] compared the laser welding Ti6Al4V process in vertical down and vertical up positions. Greater undercuts were found on the vertical up position when applying the same process parameters for both cases. More porosities were found in vertical down position where the welding fractured through during tensile testing. They recommended that low energy input with high process velocity and high laser power should be applied for both vertical up and vertical down positions in order to reduce the number of porosities and improve the welding appearance.

Modelling can also be utilized to help understand the role gravity will play during the welding process under different gravity conditions. Kang et al. [244] developed a model to investigate the influence of gravitational direction on weld pool formation and free surface deformation during a gas tungsten arc welding process. They compared the melt pool shape in three operating conditions: flat, vertical up and vertical down. They reported that the deepest melt pool was found in the vertical up position, followed by the flat position and the vertical down position, where the depth of the melt pool in the vertical up position was 50% greater than that of the vertical down position. Bahrami et al. [245] built a model to simulate the melt pool formation during the welding process with

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Chapter 6 CFD modelling of narrow gap welding based additive manufacturing process increased gravity level. They discovered that both the width and depth of the melt pool would decrease when adding the operating gravity level which could be attributed to the enhanced convection flow induced by the buoyancy force. They also found that the maximal velocity on the melt pool surface would increase while the maximal velocity within the melt pool would decrease with a stronger gravitational level.

6.2.2 Narrow gap welding

Narrow gap welding is a process where two samples are joint by filling a small gap between them with heat input and filler material. It has been widely applied for thick section welding with the advantage of being economical and bringing less distortion. Elmesalamy et al. [246] investigated the effect of process parameters on residual stress distribution during narrow gap welding of 10 mm 316L sections. Neutron diffraction and contour method were used to measure the residual stress after the experiment. They reported that laser power had an important role in determining the residual stress peak value while the welding speed would significantly affect the sustained tensile residual stress area. Another study from Elmesalamy et al. [247] compared the residual stress distribution between narrow gap welding and gas-tungsten welding processes. They found that tensile stress was 30% lower in narrow gap welding than in the GTA process and the butterfly distortion was three times higher during the GTA process. Residual stress of NGLW would also vary with the welding method (double sided or single sided). Process parameters were found to have more influence on residual stress distribution rather than the sample thickness. Cai et al. [248] conducted narrow gap welding process using two wires in vertical down position. They reported that the molten material would flow downwards to the front region of the melt pool due to gravity effect which might cause lack of fusion problem especially when the welding speed was too low. Some melt pool asymmetric issue was found with more material deposited on the front wire side. They recommended that a small distance between two wires and a relatively high process speed should be applied in order to obtain a decent quality weld. An oscillation laser beam was employed by Yamazaki et al. [249] during the narrow gap welding process. Lack of fusion between side walls and the weld could be avoided because the penetration shape could be controlled by applying the oscillation laser beam. They found that the maximum hardness of the heat affected zone was higher in the lower layers than the top layers as a

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Chapter 6 CFD modelling of narrow gap welding based additive manufacturing process result of the reheating process. They pointed out that the layer thickness should be reduced between passes in order to achieve a reduced grain size and hardness.

Numerical methods have been implemented to simulate the narrow gap welding process by many groups. An FE model was developed by Liu et al. [250] to investigate the evolution of residual stress distribution during narrow gap welding of stainless steel pipes. They discovered that peak tensile stress in the axial and hoop direction would stay constant after the certain height (around 30-43%) of the gap was filled. The level set method was applied by Desmaison et al. [251] to model a multi-pass laser/GMA hybrid welding process. In order to improve the calculating efficiency, an enhanced material thermal conductivity was applied to model the main heat transfer process. Despite the fact that no actual fluid flow was modelled, the modelling results had a good agreement with the experiments. Krampit [252] developed a theoretical model for the understanding of welding formation during the narrow gap pulse welding process. They mentioned that the amount of filler material would affect the weld quality, where a lack of fusion on the side walls might occur if too much material was added to the gap. They also reported that the minimum gap width was dependent on the peak value of the welding current density. A moving heat source model coupled with element birth and death method was applied by Phannaim et al. [253] to simulate the laser narrow gap welding with a hot wire. Temperature distribution, as well as thermal strain, was calculated to predict the formation of solidification cracking. They reported that the appearance of cracking could be reduced by using hot wire as filler material during the process.

Since gravity will have a significant influence during the welding process when performed in not-flat operating positions, the effect of gravity in the form of buoyancy force and liquid mass should not be neglected. Because FEM model is difficult to handle the fluid flow inside the melt pool which is quite essential, FVM model will be more appropriate to help understand the melt pool formation related to the fluid flow development.

However, apart from FEM simulation of the temperature history and stress distribution, not too much work has been found on modelling the actual melt pool formation taking into consideration the heat-mass transfer and fluid flow condition. Hu and Tsai [254] built a three-dimensional model to simulate the single pass gas metal arc welding (GMAW) process with a V-shaped groove. They pointed out that with the existence of V groove,

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Chapter 6 CFD modelling of narrow gap welding based additive manufacturing process fluid flow inside the melt pool would be smoothed which would result in a simpler flow pattern. The groove functioned as a confined gap which would stop fully mixing of the newly added material and melt pool.

6.2.3 Dissimilar welding

Some modelling work is found to simulate the dissimilar welding process. The VOF method was utilised by Shaibu et al. [255] to simulate the dissimilar laser keyhole welding process combining copper and 304 stainless steel. The melt pool was observed to be asymmetric along the centerline despite the fact that a symmetrical Gaussian beam was used. Melt pool was shifted towards the stainless-steel side because copper had a higher thermal conductivity than stainless steel. Fluid flow inside the melt pool was mainly driven by the Marangoni stress starting from the steel side. However, only two metal phases were included in their model where the effect of gas phases was neglected. The top surface was fixed to be flat which could not present the realistic deformable free surface situation. A two-dimensional model using level set method was developed by Tomashchuk et al. [256] to simulate electron dissimilar welding of TA6V and 316L stainless steel with copper as interlayer material. Proper keyhole offset position and welding speed were determined from the model to get a good continuity of interlayer quality. Ranjbarnodeh et al. [257] employed a 3D FEM model to investigate the effect of process parameters on residual stress distribution in dissimilar welding of stainless steel and carbon steel. They reported that tensile residual stresses would decrease when increasing the welding heat input, while a relatively low magnitude of residual stress could be achieved when a longer sample was used. They also compared three different welding methods, among which symmetric welding starting from the sample middle had the lowest residual stress. Phanikumar et al. [258] built a model on dissimilar welding of Cu and Nickel which took into account the effect of buoyancy force, convection flow induced by Marangoni stress and melting/solidification process. Stationary melt pool formation and moving heat source were both investigated in conduction mode. They reported that the melt pool appeared to be asymmetric (larger on the Nickel side) despite the fact that the heat source was symmetrically distributed along the centreline. Convection flow induced by Marangoni stress was found to be the driving force during development of the melt pool. Nickel was first melted, after which Cu was melted with

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Chapter 6 CFD modelling of narrow gap welding based additive manufacturing process the heat transferred from liquid Nickel. Insufficient mixing was observed on the copper side due to shorter time the cooper spent in the liquid state.

Considering the important role species transport played during the dissimilar welding process, many numerical models were developed to help understand the transportation of elemental composition inside the melt pool. Hu et al. [259] presented a model including the development of Fe element concentration during dissimilar welding of stainless steel and nickel. They found that at the beginning of process mass transfer rate was high with Fe element unevenly distributed in the stainless and nickel sides. The mixing rate decreased with the process going on until the elemental concentration became uniform inside the melt pool. They compared two cases with and without convection flow and they concluded that the temperature field far away from the melt pool region was not significantly affected by the convection flow. Esfahani et al. [260] developed a model to investigate the impact of process parameters on microstructure and mechanical properties in dissimilar laser welding low carbon and stainless steel. They reported that the composition distribution within the melt pool after solidification could be optimised by changing the specific energy. A homogeneous elemental mixing could be obtained by applying a relatively high specific energy. They also mentioned that the laser beam could be shifted towards the stainless-steel side, compensating the thermal conductivity difference in order to achieve a good mixture of the alloy. Bahrami et al. [261] demonstrated a dissimilar GTA welding model of Nickel 200 and Alloy 1018 taking into 휕푐 consideration the transportation of species concentration: 푖 + 풖 ∙ 훁푐 = 훁 ∙ (퐷 훁푐 ), 휕푡 푖 푖 푖 where 푐푖 and 퐷푖 are the concentration and the diffusion factor of the tracked element inside the melt pool. The transport of the major elements including nickel, iron and sulphur were calculated. The modelling results indicated that the mixture of alloy distribution experienced several stages. In the beginning, melt pool profile was asymmetric due to the thermal property difference and dissimilar material could not be fully mixed by convection flow induced by Marangoni stress on the surface. Molten material started to flow along the melt pool edge and towards the bottom of the solid- liquid interface. From their model, a particle was tracked to trace the travelling path before solidification. The path of the flowing particle seemed to cover most regions within the melt pool which would result in a homogeneous elemental distribution of the final microstructure. A dissimilar metal laser welding model was developed by Vázquez et al. [262] which included the calculation of temperature distribution and species mixing. The

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Chapter 6 CFD modelling of narrow gap welding based additive manufacturing process formation of interlayer phase mixing zone was achieved by considering the temperature and the concentration of the different reacting species in each mesh element to allow their recombination according to the stoichiometric relation. An energy source term was applied to take into account the energy released during the reaction.

6.3 Numerical modelling of narrow gap welding

6.3.1 Modelling strategy

Schematic sketch of narrow gap laser welding process is presented in Figure 6.1, where laser was used as the heat source to melt the filler wire fed into the gap and base material to form the joint.

Figure 6.1 Schematic of narrow gap welding process The dimensions of the whole calculation domain were 7 mm × 30 mm × 4.5 mm and the mesh was generated using GAMBIT which contained 483840 (56×2 40×36) hexahedral cells with uniform grid spacing 125 µm shown in Figure 6.2. The calculation time step 푢∆푡 was 3×10-4 s and tested by the Courant number which is defined as 퐶 = , where 푢 is ∆푥 the magnitude of the velocity, ∆푡 is the time step and ∆푥 is the grid size. Fluid cannot travel more than one cell in one time step meanning that the courant number cannot bigger than one. It took approximately 30 hours to simulate a single deposition line which would run 2.2 s in real time. With time step bigger than 3×10-4 s, more iterations would be required to reach the residual criteria during every calculation time step and hence the total modelling time would be longer. Grid and time-step independence test were carried out by reducing the grid spacing to 83 µm with 1632960 (84×360×54) hexahedral cells

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Chapter 6 CFD modelling of narrow gap welding based additive manufacturing process and decreasing the time step to 1.5×10-4 s. Little difference in simulation result was obtained with the finer grid and/or smaller time step, while it would take much longer time to complete simulating the same deposition length. Therefore, current grid system was refined enough to achieve accurate and stable simulation results.

7mm

4.5mm

Figure 6.2 Mesh for the calculation domain A “U” shaped groove of 3 mm width and 1.5 mm depth was initially patched with stainless 316L material as base material, on top of which the rest calculation domain was patched with air phase. Schematic sketch of the model after initialisation is shown in Figure 6.3, where air phase is in blue and base material phase is in grey.

outflow (side and top surfaces of upper domain)

wall (side and bottom surfaces of the lower domain) Figure 6.3 Boundary conditions and initial phase patch of the calculation domain

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PISO algorithm was used as the pressure-velocity coupling solution. The velocity and temperature fields were discretized with a second order upwind scheme, and the pressure field was discretized with a PRESTO! scheme. The convergence criteria for residuals of continuity and momentum equation was 10-3 and 10-6 for energy equation. Considering that the whole deposition process needed to be modelled, a transient time solution with pressure-based solver was used. VOF model combined with Geo-Reconstruct method was applied to track the interface between base material and gas phase, where user- defined mass, momentum and energy source terms were added.

Figure 6.4 Illustration of boundary conditions on cross-section plan The top and side surfaces of the upper domain were gas (air) phase boundaries treated as outflow shown in Figures 6.3 and 6.4. The bottom and side surfaces of the lower domain were base material phase boundaries defined as wall conditions. The calculation domain was chosen as part of the whole workpiece, and hence heat conduction and heat radiation were applied on the wall boundaries. Thermal transfer including heat conduction towards the base material, heat radiation towards the surrounding environment, and heat convection with the shielding gas were taken into consideration on the free interface. The free interface between steel and air was tracked with the VOF method by tracing the fraction of steel phase with 0.05

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푃 ∗휂 3∗푟2 퐻푆 = 푙푎푠푒푟 ∗ exp (− 푛) 푛 푛 3∗푟2 푟2 (6.1) 푉 ∗∫ exp (− 푛) 푐푒푙푙 1 푟2

3 where 푉푐푒푙푙 (m ) is the cell volume, 푃푙푎푠푒푟 (W) is the laser power used for the laser welding process and 휂 (%) is the laser power efficiency, 푟 is laser beam radius, 푟푛 is a radial distance from the laser beam centre. 푟푛 can be expressed as:

2 2 푟푛 = √[푦푛 − (푣푙푎푠푒푟 ∗ 푡 + 푦0)] + (푥푛 − 푥0) (6.2) where 푣푙푎푠푒푟 (m/s) is welding speed, 푡 is the calculation flowing time, 푥0 and 푦0 are the locations of starting point inside the groove on the base material. Filler wire was 316L which was fed into the melt pool travelling together with the laser beam. Wire feeding was applied on the free interface as mass source and added to the continuity equations as:

휌∗휋푟2 ∗푓 푀푆 = 푤푖푟푒 푤푖푟푒 (6.3) 푛∗푉푐푒푙푙 where wire feed rate 푓푤푖푟푒 is 1.2 m/min and the wire radius 푟푤푖푟푒 is 0.6 mm. Normal surface tension force and tangential surface tension stress induced by temperature gradient were applied as volumetric forces added to the momentum equation:

휌푎푣푔κ푖∇퐹푖 푆푛표푟푚푎푙 = 훾푖푗 1 (6.4) (휌 +휌 ) 2 푖 푗 where 휌푎푣푔 is the average density of the interfacial cell related to the phase volume fraction, 휌푖 and 휌푗 are the density of phase i and j respectively. Marangoni stress on tangential direction of the surface induced by the temperature dependent surface tension coefficient can be presented as:

휕훾 휕푇 휏 = (6.5) 푀푎푟푎푛푔표푛푖 휕푇 휕풏

The momentum brought by the wire feeding velocity was also taken into account. Solidification and melting process was taken into consideration with the enthalpy- porosity method by applying a damping force to the mushy material to decrease the fluid velocity. The gravity force was added to the cells with temperature higher than the melting point. The Boussinesq approximation was used to present the buoyancy force to improve the calculation speed.

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6.3.2 Lack of fusion

Lack of fusion during narrow gap welding happens when heat input is not high enough to fully melt the newly added filler material, substrate (or previously deposited layer) and side walls. Lack of fusion may lead to many problems including the appearance of long cracking and low tensile strength mechanical properties. A concave shaped weld bead is normally preferred to prevent lack of fusion from happening. Lack of fusion between weld bead and side walls was successfully simulated using this current model. The resultant weld morphology and temperature distribution are shown below in Figure 6.5.

a b

c d

e f

Figure 6.5 First layer narrow gap welding, low laser power (a,b), higher laser power(c,d), after re-melting (e)(f)

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Figure 6.5 (a) shows the welding morphology when 2 kW laser input was used and Figure 6.5 (b) presents the temperature distribution at the end of the welding process. Figure 6.5 (c) and (d) illustrate the welding results after increasing the laser power to 3 kW. Figure 6.5 (e) and (f) show the welding results after applying a re-melting pass based on the bead using 2 kW laser power.

Liquid fraction

a b

Liquid fraction

c d

Liquid fraction

e f

Figure 6.6 Middle plane cross-section and melt pool velocity field: low laser power (a)(b), higher laser power (c) (d), after re-melting (e)(f)

Cross-section of the weld shape and melt pool-velocity field for the three cases are presented in Figure 6.6. Some surface irregularity can be observed in Figure 6.5 (a) when the heat input was not high enough to fully melt the added material and the side wall,

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Chapter 6 CFD modelling of narrow gap welding based additive manufacturing process which can also be verified in Figure 6.6 (a) where the bead shape appeared to be convex and the side walls were not completely wetted. By increasing the laser power, better surface quality can be achieved in Figure 6.5 (c) and the weld shape turned into concave, although some part of the side wall was still not completely melted. By applying a re- melting pass on the deposited layer, weld surface was improved greatly shown in Figure 6.5 (e) and a smooth concave shaped weld could be obtained in Figure 6.6 (e). Melt pool was more elongated in Figure 6.6 (f) during the re-melting process as a result of the temperature increase inside the groove. It can also be noticed that in Figure 6.6 (f) longer side walls were melted in comparison with Figure 6.6 (b) and Figure 6.6 (d). During the re-melting process, more heat was transferred to the side walls causing enough mixing between the weld and the side walls, which would explain the good surface finish and smooth concave shaped welding result. However, with more heat input during re-melting or increasing the laser power, the final sample temperature would increase shown in Figure 6.6 (b)(d)(f). Excessive heat input may cause unwanted residual stress and deformation.

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6.3.3 Effect of gravity direction on narrow gap welding process

g x g x a v b v

g g

c v d v

g ⊙ g ⊙ e v f v

g g

g v h v

g g

i v j v

Figure 6.7 second and third layer for different directions : flat (a)(b), horizontal (c)(d), overhead (e)(f), vertical up (g)(h), vertical down (i)(j)

No previous work has been found on modelling the multilayer weld formation during the narrow gap welding process, especially when the process is operated in various directions. Figure 6.7 presents the simulated results of second and third layer of the narrow gap welding process in different operating directions. The gap was not completely filled for all the cases in different positions after the second layer was deposited. The difference in morphology of the unfilled region could be related to the melt pool formation shown in Figure 6.8.

After the melt pool entered steady state during the second pass, a slope would occur on the deposition surface which could be observed from all the cases. Due to the gravity effect, the slope shape inside the groove appeared to be different in different operating directions. The melt pool in horizontal position was asymmetric with a bigger area in the lower part since gravity force would drag the molten material towards the gravitational direction. In comparison with the flat position, the unfilled groove region appeared to be

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Chapter 6 CFD modelling of narrow gap welding based additive manufacturing process shallower in the overhead position because of the gravity effect. A large portion of the side wall was not fully melted in the vertical positions after two layers, where a bulge existed at the beginning of the weld in the vertical up position while a relatively shallower melt pool slope could be observed in the vertical down position.

g x g x a v b v

g g

c v d v

g ⊙ g ⊙ e v f v

g g

g v h v

g g

i v j v

Figure 6.8 Second layer melt pool and third layer temperature: flat (a)(b), horizontal (c)(d), overhead (e)(f), vertical up (g)(h), vertical down (i)(j)

In order to fill the groove, a third pass was deposited into the gap and the final results are shown in Figure 6.7 (b)(d)(f)(h)(j). Without the groove, the bead shape changed from concave to convex as a result of surface tension. Not much difference could be noticed

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Chapter 6 CFD modelling of narrow gap welding based additive manufacturing process between flat and overhead positions, apart from that the slope angle in the solidification front was slightly bigger for overhead position. However, more irregularity or dripping phenomenon was believed to appear when higher power and more material were added into the groove, where normal surface tension force could no longer hold the molten material. The weld was shifted towards the gravitational direction and the collapse of deposition could be observed on the sample surface in horizontal position. The bulge at the beginning of the weld became bigger in vertical up position while in vertical down position more material was accumulated in the solidification front causing surface irregularity on the final welding result.

Figure 6.9 Melt pool development history of first two layers Melt pool development history can be used to help understand the effect of gravity on bead formation shown in Figure 6.9 and Figure 6.10. Considering that the melt pool would behave differently with the influence of side walls inside the groove and without the restriction of gap on the surface, Figure 6.9 presents the melt pool development of the first two passes when the melt pool was still lower than the top gap surface while Figure 6.10 shows the whole welding process of the three passes.

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Figure 6.10 Melt pool development history of three layers No distinct difference among positions could be observed in the melt pool development during the first layer. With the heat accumulation effect during the second layer, melt pool area and the total melt pool volume appeared to be the highest in vertical down position and the lowest in vertical up position. This could be explained by the fact that in the vertical down position the newly melted material would have the potential to flow along the gravity direction which was the same direction as the welding speed and hence the liquid would accumulate at the melt pool font contributing to a high total melt pool volume. On the contrary, in vertical up position the just melted material would flow along the gravity direction against the welding direction to a relative cool region resulting in a bulge at the tailing area of the advancing melt pool, which in turn would lead to a lower melt pool area and total melt pool volume. In horizontal position, with the expanding of the melt pool, surface tension could no longer hold the gravitational force of the molten material and hence the collapse happened. This brought the highest melt pool volume among all the cases. However, with a bigger calculation domain the molten material would keep on falling in the air phase instead of lying on the side wall which could result in different melt pool development history. In the future work, a bigger calculation domain will be created to let the simulation run in a bigger scale and more phenomena can be observed.

6.3.4 Effect of gravitational level on narrow gap welding process

Some of the previous research has reported that an increased gravitational level could enhance the convection flow inside the melt pool induced by the buoyancy force. No

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Chapter 6 CFD modelling of narrow gap welding based additive manufacturing process research has been found on discussing the effect of gravitational level on welding bead formation when filler material is added to the process. With the introduction of additional material, under various gravitational levels the melt pool might behave differently compared to the conditions where only base material is melted. Therefore, narrow gap welding process in vertical up position with increased gravitational level 5 g and decreased gravitational level 0 g were investigated and the results were compared with normal gravity 1 g. The results are presented in Figure 6.11 below.

g g

a v b v

g g

c v d v

g g

e v f v f Figure 6.11 Second and third pass for vertical up with: 0 g (a)(b), 1 g (c)(d), first and second pass for vertical up 5 g (e)(f)

The main purpose of this model was to understand the melt pool and weld bead formation inside the gap, and hence only limited air phase region on top of the base material was included when patching the phases during initialization. With the effect of increased gravity force on molten material, collapsing happened during the second layer under 5 g with some of the molten material falling out of the domain from the air phase. Therefore, only two passes were modelled for the case with 5 g. Not too much wetting on the side walls could be observed in Figure 6.11 (e) and a convex shaped bead was formed as a result of gravity. The enhanced gravity level brought more irregularity during the second pass shown in Figure 6.11 (f). It can be concluded from Figure 6.11 that with the increase of gravity level, the weld bead formed in the vertical up position tended to show more convex on the surface and more lack of fusion with the side walls.

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g g

a v b v

g g

c v d v

g g

e v f v

Figure 6.12 Melt pool-velocity field and temperature distribution at end of second pass: 0 g (a)(b), 1 g (c)(d), 5 g (e)(f) Figure 6.12 presents the melt pool-velocity field and temperature distribution of the three cases. With the increase of gravity level, the slope angle of the melt pool front became smaller. Apart from this, it could be easily noticed that the melt pool was much longer under 5 g condition in Figure 6.12 (e) when compared with the other two cases.

Figure 6.13 Melt pool development history of first two passes The melt pool development history could be used to help understand the bead formation under different gravity levels. Two stages could be observed for the case when enhanced

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Chapter 6 CFD modelling of narrow gap welding based additive manufacturing process gravity level was applied. Due to the gravity effect, melt pool would get elongated with the liquid material flowing towards the rear region of the melt pool which would increase the melt pool surface. However, since the molten material was flowing towards a relatively cool area opposite to the laser moving direction, the liquid metal would experience a shorter time to be completely solidified. This could explain the smallest total melt pool volume under 5 g in the first stage. With more liquid material flowing and accumulating towards the gravity direction, some waviness would occur and cause some distinct humps on the weld surface. Heat dissipation rate was much lower when melt pool existed on these humps since there was less contact between these humps and the side walls. As a result, heat could not be conducted quickly which contributed to the second stage where a dramatic rise on total melt pool volume occurred.

Figure 6.14 Melt pool development history of three passes This rapid increase could also be noticed for the rest two cases with lower gravity level during the third welding pass shown in Figure 6.14. The deposited bead was no longer inside the groove during the third pass and as a result the heat could not be conducted through the side walls. The number of heat conduction directions decreased from three to one (substrate only), which resulted in a dramatic rise in total melt pool volume.

6.3.5 Effect of surface tension coefficient on narrow gap welding

Surface tension plays an important role in determining the melt pool formation and the effect of surface tension has been reported by many people under various welding conditions. However, not too much work has been found on investigating the influence

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Chapter 6 CFD modelling of narrow gap welding based additive manufacturing process of surface tension coefficient during the narrow gap welding process with additional material fed into the groove. Surface tension might function differently with the existence of the side walls. Cases with positive and mixed surface tension coefficient were simulated in flat welding position and the results were compared with negative surface tension coefficient situation. The weld bead morphology after second and third pass of each case is presented in Figure 6.15.

g x g x a v b v

g x g x c v d v

g x g x e v f v

Figure 6.15 Second and third pass of the weld with different surface tension coefficient: negative ∂γ/∂T (a)(b), positive ∂γ/∂T (c)(d), mixed ∂γ/∂T (e)(f)

g x g x a v b v

g x g x c v d v

g x g x e v f v

Figure 6.16 Second pass melt pool-velocity field and third pass temperature distribution: negative ∂γ/∂T (a)(b), positive ∂γ/∂T (c)(d), mixed ∂γ/∂T (e)(f)

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Positive surface tension coefficient would promote an inward convection flow from edge to the centre of melt pool while negative surface tension coefficient would create an outward convection flow on the melt pool surface. Melt pool tended to appear wider and shallower in the negative case while narrower but deeper in positive case. As a result, during the second layer the weld bead would appear to be less concave inside the groove 휕훾 with positive as shown in Figure 6.15 (c). Wetting condition with the side walls was 휕푇 휕훾 not as good as the case under negative . During the third pass without restriction from 휕푇 휕훾 the side walls, the melt pool with negative spread out and the final weld bead was much 휕푇 wider than the groove. Figure 6.16 illustrates the comparison of melt pool-velocity field 휕훾 of the second pass and temperature distribution of third pass. Negative had the biggest 휕푇 휕훾 휕훾 fully molten area (red region), followed by mixed while positive had the smallest 휕푇 휕푇 melt pool surface. However, the maximum temperature inside the melt pool was the 휕훾 highest in positive case since the inward convection would slow down the heat transfer 휕푇 휕훾 towards the surrounding environment. For mixed situation, a transitional temperature 휕푇 existed for the surface tension coefficient to change from positive to negative. Liquid metal tended to flow both from melt pool centre and edge to this transitional point which brought a multi-directional and more complicated flow condition inside the melt pool.

Figure 6.17 Melt pool development history of first two passes Melt pool development history could also be used to help understand the effect of surface tension coefficient on bead formation. Figure 6.17 presents the melt pool formation 휕훾 during the first two passes. Negative brought about biggest melt pool surface but the 휕푇

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휕훾 total melt pool volume was the lowest. Positive had the smallest melt pool surface with 휕푇 the highest total melt pool volume due to the fact that the inward conviction flow was against the temperature gradient which would slow down the heat dissipation rate.

Figure 6.18 Melt pool development history of three passes During the third pass, heat could no longer be transferred by conduction through the side walls which would contribute to a dramatic increase in total melt pool volume and the difference of melt pool volume among cases was more significant shown in Figure 6.18.

6.3.6 Narrow gap welding of thick sections

Multi-pass narrow gap welding with 7.5 mm groove depth was successfully modelled. Figure 6.19 shows development of the weld bead formation for the first six passes, where 1 s time interval was applied between passes.

Figure 6.19 Weld bead development

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Figure 6.20 presents the development of cross-section of first six passes. It could be noticed that during the first two passes, the shape of the deposited weld was not completely concave and only limited wetting with the side walls could be achieved. This was quite common during the narrow gap welding process due to the fact that the sample temperature was relatively low at the beginning of the process when the input laser could not provide enough energy to melt all the added material together with the side walls to produce a smooth concave shaped weld. In the third pass, some lack of fusion could be observed between passes and interestingly it disappeared during the re-melting process from the fourth pass. After the fourth layer, because of the heat accumulation effect, weld bead turned to be smooth concave shape and a good wetting between the weld and side wall was achieved.

Figure 6.20 Cross-section development (middle y plane)

Figure 6.21 Cross-section development (middle x plane)

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Figure 6.21 presents the cross-section development of the welds in middle x plane. Some difference could be noticed when compared with the deposition on flat surface without the existence of side walls. Slope instead of bulge occurred at the beginning of the weld which could be explained that the side wall was able to transfer most of the heat by thermal conduction and hence the surface tension variation at the beginning was not too distinct. Slope still existed at end of the weld bead because of the tilted melt pool front.

Figure 6.22 Melt pool and temperature distribution of first four passes

Figure 6.23 Melt pool development history

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Figure 6.22 shows the melt pool-velocity field and temperature distribution of the first four passes in the middle moment. With the increase of overall sample temperature, melt pool became more elongated which could also be verified by the melt pool development history presented in Figure 6.23. With the heat accumulation effect, melt pool kept growing with more passes being deposited. In real experiments, laser power is normally reduced for the upper passes to keep a constant interlayer temperature.

6.4 Numerical modelling of narrow gap dissimilar welding

6.4.1 Modelling strategy

Based on the narrow gap welding model, two more phases were added to realize laser narrow gap dissimilar welding process. Two base materials were 316L and SA508 with filler material Alloy 52. The calculation domain after initialization is shown in Figure 6.24. The main chemical content of SA508 and Alloy 52 is shown in Table 6.1.

Figure 6.24 Schematic sketch of the calculation domain Table 6.1 Chemical composition of SA 508 Gr.3 Cl.2 steel and Alloy 52 (wt.%) [263]

Materials C Ni Fe Cr Si Mn Mo S Cu SA508 0.2 0.8 Bal. 0.2 0.25 1.4 0.50 0.004 0.04 52 0.03 59.3 9.8 29.2 0.13 0.24 0.03 <0.001 0.02

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The thermal conductivity of SA508 is approximately 40 W·m-1·K-1 [263] and for Alloy 52 is 13 W·m-1·K-1 [264]. Therefore, SA508 has a higher thermal diffusivity compared with 316L. As a result, the temperature will rise more quickly on the 316L side when laser starts to impact on the sample surface. Since these two materials have a similar melting range around 1700 K, 316L material will melt first and the resultant melt pool will have the potential to shift towards 316L side which will contribute to an asymmetric melt pool shape along the centreline inside the groove.

6.4.2 Modelling of bead asymmetry

In this model, 3 kW laser power was first applied on the sample surface in flat welding position and the result is shown in Figure 6.25(a). Lack of fusion on the SA508 side could be easily noticed, where not too much melting and wetting happened on the SA508 side wall. In order to eliminate this phenomenon, welding in the horizontal position was conducted where SA508 was placed on the bottom side and the welding result is shown in Figure 6.25(b). Gravity force would drive the liquid metal flow towards the SA508 side and hence better melting on the SA508 side wall could be achieved. Another method was to increase the laser input in order to deliver more total energy to the sample. Laser power was increased from 3 kW to 4 kW and the welding result is presented in Figure 6.25(c). As a result, more material could be melted and a larger melt pool would be created, which contributed to better wetting on both side walls and smooth concave shaped weld bead could be achieved inside the groove. Because of higher total heat input, the melt pool would become more unstable which would bring about stronger mixing of the three materials.

Figure 6.26 presents the melt pool status and temperature distribution of these three cases. SA508 side wall could be better melted when changing from flat to horizontal position with 3 kW laser power. Melt pool shape and temperature distribution became more symmetric along the groove centreline when enough heat input was provided, where the phenomenon of lack of fusion on SA508 side wall was eliminated.

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a SA508

Alloy 52 v g x 316L

b SA508

Alloy 52 v g 316L

c SA508

Alloy 52 v g x 316L

Figure 6.25 (a) 3 kW flat position (b) 3 kW horizontal position (c) 4 kW flat position

SA508 SA508 Alloy 52 Alloy 52 v v a g x 316L b g x 316L

SA508 SA508 Alloy 52 Alloy 52 v c g 316L d v g 316L

SA508 SA508 Alloy 52 Alloy 52 v e g x 316L f v g x 316L

Figure 6.26 (a,b)3 kW flat position (c,d) 3 kW horizontal position (e,f) 4 kW flat position

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6.4.3 Modelling of re-melting process

a b c

d e f

Figure 6.27 (a,b,c) before re-melting, (d,e,f) after re-melting

v a g x b v g x

Figure 6.28 (a) before re-melting, (b)after re-melting Figure 6.27 illustrates the comparison of welding result before and after the re-melting process. Figure 6.28 presents another view of the overall welding bead comparison. Increased wire feedrate and laser power were applied to fill the gap with one pass. The unevenness of the weld surface and lack of fusion with the side walls can be observed in Figure 6.27(a) and Figure 6.28(a). Surface roughness was improved and lack of fusion was eliminated after re-melting shown in Figure 6.27(d) and Figure 6.28(b).

6.5 Conclusion

A three-dimensional model was built which covered various aspects for the narrow gap welding process. Lack of fusion phenomenon between weld bead and side walls was first simulated which could be attributed to the low heat input that could not provide enough energy to melt the filler material and side wall. After applying an increased laser power or conducting a re-melting pass, better surface quality could be achieved and lack of

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Chapter 6 CFD modelling of narrow gap welding based additive manufacturing process fusion with the side walls could be eliminated. The shape of the weld bead would change from convex to concave and a good wetting with the side walls could be realised.

Effects of gravity on narrow gap welding in various positions were investigated. Slope shape of the melt pool front would appear differently under the influence of gravity. On the other hand, the effect of the gravity was decreased with the existence of side walls which would help conduct most of the heat, when compared to the deposition process on a substrate without the groove. During the last pass, gravity would cause some serious dripping and collapsing problem in some of the non-flat positions. This was attributed to the formation of a bigger melt pool on the surface and the balance between gravity and surface tension force was disrupted.

Different gravitational level was applied during the narrow gap welding process in vertical up position. More lack of fusion between the weld bead and side wall could be observed with the increase of gravitational level. Increased gravity force would drive more liquid material to the rear region of melt pool which would contribute to a more serious dripping and more convex shaped welding bead. With the help of melt pool development history, it could be noticed that when the weld bead was no longer in the groove or when there was less direct contact with the side walls, melt pool volume would increase dramatically because less heat could be dissipated through conduction.

Effect of surface tension coefficient on narrow gap welding process was investigated. When positive surface tension coefficient was applied, the weld bead would appear to be less concave inside the groove and wetting condition with the side walls was not as good as the case under negative ∂γ/∂T condition. Without the existence of side walls, the weld bead would show more difference between negative and positive coefficient cases during the last pass on the surface. Surface tension coefficient played a vital role in determining the fusion condition with the side walls inside groove.

Weld bead development with multiple passes during narrow gap welding of thick sections was investigated. It was found that the bead shape would become more concave with more passes being deposited due to the heat accumulation. Lack of fusion between passes sometimes could be eliminated by the re-melting process from the next pass.

Process of narrow gap welding with dissimilar materials was successfully simulated. The asymmetric formation of melt pool and weld bead along the centreline could be attributed

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Chapter 6 CFD modelling of narrow gap welding based additive manufacturing process to the difference in material thermal properties, which as a result would cause lack of fusion with the side walls. The bead asymmetry problem could be avoided by performing the welding process in horizontal position or increasing the total heat input to help melt the side wall of higher thermal conductivity. Re-melting process was found to be useful when applying on the welded surface to improve the surface quality and eliminate the lack of fusion problem with the base material.

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Chapter 7 Conclusions and future recommendations

Chapter 7 Conclusion and Future work recommendations

7.1 Conclusions

This research investigated the heat and mass flow characteristics in a number of laser additive manufacturing processes, with a particular focus on the effect of gravity and deposition orientation. The main findings and conclusions from this study are presented below.

The formulation of a three-dimensional computational fluid dynamic model for laser additive manufacturing process was presented which took consideration of compound process factors including material addition, surface tension, melting and solidification process, buoyancy force, temperature dependent material properties and moving Gaussian laser beam heat input. The volume of fraction (VOF) method was applied to track the free surface between gas and material phase. Self-adaptive mass and energy source terms were developed to realise the conservation of total material and heat input with the evolvement of the free surface.

The relationship between various process parameters and single-layer deposition dimensions were investigated experimentally and numerically, where a good agreement was reached. The influential factor level of each process parameter was also discussed and powder federate was found to be the most influential factor determining the deposition dimensions, under the condition that enough laser power was provided to fully melt the powder material.

Multiple layer deposition with adjacent passes was also successfully performed through experimental and numerical study. Process parameters including process patterns were found to be important in determining the final deposition surface quality.

A new three-dimensional model was built to simulate the deposition process with various complex features. The development of deposition process could be understood by analysing the melt pool surface area and total melt pool volume. The influence of existing layer on the new developing melt pool could be observed and explained by the model.

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Control of spot size and inter-pass temperature were found to be essential in real deposition environment with different geometric features required.

Modelling of re-melting process on improving the deposition surface roughness was presented and selection of the optimised re-melting parameters was found important to achieve an improved surface quality.

Modelling of full three-dimensional deposition with overhanging structures were demonstrated for the first time and the effect of gravity on deposition process was observed.

The effect of gravity on final deposition result in various process directions was further investigated experimentally and numerically. It was found that the gravity had a distinct effect during the process when performed in non-flat directions, especially in vertical up and overhead directions. The dripping phenomenon would occur as a result of gravity effect during the formation of melt pool, where gravitational force tended to enlarge the rear region of the melt pool and the molten material would accumulate in this region until the laser beam moved far away. The melt pool would get elongated before the channel between the rear and front region of the melt pool broke, after which another cycle would begin.

The effect of process parameters on dripping phenomenon was also discussed and a conclusion could be reached that when applying the same material feeding rate, the process performed with bigger spot size would have less bead shape irregularity comparing with smaller spot size. This might attribute to the fact that a bigger ratio between molten material height and width would result in a more prominent bead angle where gravity would have more effect on the molten material and the balance between surface tension and gravity was more likely to be disrupted.

Apart from the dripping phenomenon, other surface unevenness including bulge at the beginning and collapsing in the end were investigated. Irregularities caused by thermal behaviour and system instability were explained separately. It was found that due to the surface tension variation when starting, a bulge would appear at the beginning of deposition. After the melt pool development reached steady state, a slope would occur at the melting front causing the collapsing of deposition at the end. System instability including robot acceleration and deceleration, synchronizing delay between laser and

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Chapter 7 Conclusions and future recommendations wire feeder would also add to the formation of surface irregularity. More distinct difference in deposition height was noticed which increased from 0.76 mm to 1.16 mm when the effect of acceleration and deceleration was considered. Bulge width was also found to increase from 2.8 mm to 2.97 mm due to the smaller average process velocity during the acceleration. The developed model was proved to be capable of simulating all of the irregularities caused by both thermal behaviour and system instability.

The gravity effect was also found to have an impact on the bulge formation. In vertical down position the gravity would decrease the bulge width from 2.80 mm to 2.75 mm when compared with flat position, while the bulge width would be enlarged in vertical up and overhead positions to 3.07 mm and 2.88 mm respectively.

The effect of surface tension coefficient on melt pool formation in the non-flat deposition was investigated. It was found that with a bigger melt pool, normal surface tension force rather than tangential surface tension stress would have more impact on the development of overall melt pool formation.

For potential space applications, gravity was gradually decreased from 1 g to zero g including the moon gravity 0.16 g to investigate the effect of reduced gravity on final deposition. An extra case with increased gravity 2 g was also considered for the possible overweight situations. It was found that during the first layer there was no clear variation in deposition morphology among cases with different gravity values. During the deposition of the second layer, surface unevenness became increasingly prominent with the reduction of gravity value. When decreasing the gravity from 2 g to zero g, contact angle would increase from 93.0º to 99.3º while aspect ratio would decrease from 1.369 to 0.974. In flat deposition position, the component force of gravity would help surface tension better hold the liquid metal. Without the effect of gravity, the liquid metal would have the potential to form a sphere with increased contact angle under the impact of low aspect ratio and high melt pool volume & surface area, and hence the irregularity of the deposition would become more distinct. Process parameters were optimised to reduce the contact angle under zero gravity situation in order to improve the surface result. The contact angle was decreased from 99.3º to 84.4º and 79.5º when the scanning velocity was increased from 0.01 m/s to 0.015 m/s and 0.02 m/s, keeping the other process parameters constant. Reduced contact angle from 99.3º to 79.4º was achieved by halving the wire feed rate from 0.02 m/s to 0.01 m/s. The contact angle was reduced by 20.1º from

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Chapter 7 Conclusions and future recommendations

99.3º to 79.2º when increasing the laser power from 1.3 kW to 1.6 kW. However, increased heat input can lead to an enlarged melt pool size with higher liquid metal volume, which in turn may cause a reduction in layer height and deposition collapsing.

A three-dimensional model was successfully developed to simulate the narrow gap welding process. With the existence of side walls, melt pool and deposition formation inside the groove would behave differently compared with laser metal deposition process.

Lack of fusion phenomenon between the weld bead and side walls was simulated, for the first time, which was found to be attributed to the overall low heat input. After applying an increased laser power or conducting a re-melting pass, better surface quality could be achieved and lack of fusion with the side walls could be eliminated, where the shape of the weld would change from convex to concave and a good wetting with side walls could be realised.

Slope shape of the melt pool front would appear differently under the influence of gravity in different welding positions. Gravity could cause some serious dripping and collapsing problems in some of the non-flat positions during the last pass. This could be explained by the formation of a bigger melt pool on the surface where the balance between gravity and surface tension force was disrupted.

More lack of fusion between the weld bead and side wall could be observed with the increase of gravitational level when the welding process was performed in vertical up position. Increased gravity force would tend to drive more liquid material to the rear region of the melt pool which would contribute to a more serious dripping and more convex shaped welding bead.

When positive surface tension coefficient was applied, the weld bead would appear to be less concave inside the groove and wetting condition with the side walls was not as good as the case under negative ∂γ/∂T condition.

During narrow gap welding of thick sections, it was found that the bead shape would become more concave with more passes deposited due to the heat accumulation. Lack of fusion between passes could be eliminated by the re-melting process from the next pass.

The process of laser narrow gap welding with dissimilar materials was successfully simulated. The asymmetric formation of melt pool and weld bead along the centreline

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Chapter 7 Conclusions and future recommendations could be attributed to the difference in material thermal properties, which as a result would cause lack of fusion with the side walls. The bead asymmetry problem could be avoided by performing the welding process in the horizontal position or increasing the total heat input to help melt the side wall with higher thermal conductivity. The re-melting process was found to be useful applying on the welded surface to improve the surface quality and eliminate the lack of fusion problem with the base material.

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7.2 Future work recommendations

• The current model uses 316L as deposition material during the calculation. Some other widely used materials like Ti-6Al-4V can be easily added to the simulation process. Due to the difference of material properties, the deposition process and bead formation can behave differently. • Overhang structures can be further investigated by considering different depositing methods based on different required geometries and process parameters can be optimised to minimise the effect of gravity. • This numerical model can be further developed into a software package to simulate the whole deposition process with all complex features and validate the feasibility of certain deposition process before conducting the real experiments, especially for the deposition with overhanging features. An improved manufacturing efficiency can be obtained by avoiding material wastage and saving the time spent on parameter optimization. • For narrow gap welding process, different groove shape should be taken into consideration. Thermal stress distribution related to the temperature history can be modelled by coupling the current model with FEM. Residual stress induced deformation will be included in the future narrow gap welding simulation. • The current model does not include species diffusion when simulating narrow gap welding process with dissimilar materials. Species diffusion can be added to the model in the future to achieve a better result of material mixing when multiple materials are included. • Metallurgical model for phase transformation can be coupled with current model to get a better understanding of the microstructure development based on the thermal behaviour history of the process. • Based on the current model, deposition process in vacuum environment will be investigated together with non-gravity condition for potential space application. • Some numerical diffusion problem still exists during the calculation. The combined VOF-Level set method can be applied in the future to achieve both mass conservation and sharper free surface tracking.

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