LASER NARROW GAP WELDING OF THICK SECTION ...

LASER NARROW GAP WELDING OF

THICK SECTION DISSIMILAR

METALS

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

Doctor of Philosophy in the Faculty of Science and Engineering

2019

Timo Tapio Väistö

School of Mechanical, Aerospace and Civil Engineering

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Table of Contents

List of Figures ........................................................................................ 11

List of Tables ......................................................................................... 25

List of Symbols ...................................................................................... 29

List of Abbreviations............................................................................. 30

Abstract .................................................................................................. 33

Declaration ............................................................................................. 34

Copyright Statement ............................................................................. 35

Acknowledgements ................................................................................ 36

1 Introduction ............................................................................... 37

1.1 Background .................................................................................................. 37

1.2 Primary Cooling Circuit Pressure Vessel Nozzles....................................... 38

1.3 Narrow Gap Laser Welding ......................................................................... 39

1.4 Research Motivation and Research Questions ............................................. 40

1.5 Aim and Objectives...................................................................................... 41

1.6 Structure of This Thesis ............................................................................... 43

2 Literature Review ..................................................................... 45

2.1 Pressurised Water Reactor (PWR) ............................................................... 45

2.1.1 Primary Cooling Circuit Environment (PCC).............................................. 46

2.1.1.1 Heat and Pressure ......................................................................................... 46

2.1.1.2 Corrosion...................................................................................................... 47

2.1.1.3 Other Considerations ................................................................................... 48

2.2 Narrow Gap Welding ................................................................................... 49

2.2.1 Filling strategies ........................................................................................... 52

2.3 Narrow Gap Laser Welding ......................................................................... 53

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2.3.1 Challenges in NGLW ................................................................................... 58

2.3.2 Welding efficiency ........................................................................................ 58

2.3.3 Welding Parameters for Narrow Gap Laser Welding ................................... 58

2.3.4 Hot Wire Laser Welding (HWLW) .............................................................. 63

2.4 Present-day Solutions in Ferritic-Austenitic Dissimilar PWR PCC

Welding ......................................................................................................... 65

2.4.1 Dissimilar Welding Processes and Geometries Used in PWR Primary

Cooling Circuit Manufacturing ..................................................................... 67

2.4.2 Metals Used in PWR Primary Cooling Circuit Construction ....................... 72

2.4.2.1 Base Materials .............................................................................................. 73

2.4.2.2 Filler Metals .................................................................................................. 76

2.5 Considerations in Austenitic-Ferritic Dissimilar Metal Welding in PWR

Cooling Circuit ............................................................................................. 80

2.5.1 Corrosion ...................................................................................................... 81

2.5.2 Sensitisation of Austenitic Stainless Steel .................................................... 82

2.5.3 Residual Stresses .......................................................................................... 84

2.5.3.1 Types of Residual Stresses ........................................................................... 85

2.5.4 Environmentally Assisted Cracking and Stress Corrosion Cracking ........... 87

2.5.5 DMW Creep Failure in the HAZ of Low Alloy Steel .................................. 89

2.5.6 Radiation Damage ........................................................................................ 90

2.6 Dissimilar Metal Welding (DMW) ............................................................... 91

2.6.1 Phase Prediction of Austenitic Weld Metal .................................................. 92

2.6.2 Solidification Mechanisms in Dissimilar Metal Welding ............................ 93

2.6.3 The Heat Affected Zone ............................................................................... 95

2.6.3.1 Carbon Equivalent ........................................................................................ 98

2.7 Weldability of Alloy 52 Filler Metal ............................................................ 99

2.7.1 Solidification Cracking ............................................................................... 101

2.7.2 Ductility Dip Cracking ............................................................................... 105

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2.7.3 Liquation Cracking .................................................................................... 107

2.8 Laser Welding of Dissimilar Materials ...................................................... 109

2.8.1 HAZ in Laser Welding............................................................................... 112

2.8.2 Laser Welding Parameters ......................................................................... 115

2.8.3 Laser Welding With Filler Material ........................................................... 116

2.9 Testing of Weldments ................................................................................ 119

2.9.1 Destructive Testing .................................................................................... 119

2.9.2 Non-Destructive Testing ............................................................................ 122

2.9.2.1 Visual Testing ............................................................................................ 122

2.9.2.2 Radiographic Testing ................................................................................. 123

2.9.3 Measurement of Residual Stresses and the Contour Method .................... 124

2.10 Standards for evaluation of welds in nuclear applications......................... 129

2.11 Summary of the Literature Review and Rationale for the Current Work .. 136

2.11.1 Knowledge gaps ......................................................................................... 139

3 Materials, Methods and Equipment ..................................... 141

3.1 Welding Programme Stages ....................................................................... 141

3.2 Materials .................................................................................................... 142

3.2.1 Stage I – Similar Metal Welding in Stainless Steel ................................... 142

3.2.2 Stage II – Dissimilar Metal Welding with Mild Steel ............................... 143

3.2.3 Stage III - Dissimilar Metal Welding with Pedigree Steel ........................ 144

3.3 Sample Geometry....................................................................................... 145

3.4 Welding Groove Design ............................................................................ 146

3.5 Welding Parameters ................................................................................... 149

3.6 Welding Setup ............................................................................................ 151

3.6.1 Laser and Robot Used ................................................................................ 151

3.7 Development of the Equipment and Procedures ........................................ 152

3.7.1 Welding Restraint ...................................................................................... 153

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3.7.2 Wire Feed Arrangement ............................................................................. 157

3.7.3 Shielding Gas Setup .................................................................................... 164

3.7.4 Welding Procedure Development ............................................................... 169

3.7.5 Alignment of the Weld and Equipment ...................................................... 170

3.8 Monitoring of the Welding Process ............................................................ 171

3.8.1 Interpass Temperature Measurement .......................................................... 172

3.8.2 Measurement of Welding Distortions ......................................................... 172

3.8.3 Laser Illumination Imaging ........................................................................ 175

3.9 Measurement Uncertainties ........................................................................ 180

3.9.1 Laser Power ................................................................................................ 180

3.9.2 Spot Size Control ........................................................................................ 182

3.9.3 Accuracy of Welding Consumable Feed .................................................... 182

3.9.4 Welding Robot Accuracy ........................................................................... 183

3.9.5 Accuracy of the Alignment of the Welds ................................................... 183

3.9.6 Welding Groove Dimensional Accuracy .................................................... 183

3.9.7 Accuracy of the Distortion Measurements ................................................. 183

3.9.8 Summary ..................................................................................................... 184

3.10 Analysis Methods for the Weldments ......................................................... 184

3.10.1 X-ray Radiography and Acceptance ........................................................... 185

3.10.2 Microstructural Analysis and Triple Etching ............................................. 185

3.10.3 Hardness Mapping and Hardness Line Scanning ....................................... 187

3.10.4 Contour Method Residual Stress Measurement Procedure ........................ 189

3.10.5 Tensile Testing ............................................................................................ 197

3.10.6 Digital Image Correlation ........................................................................... 198

3.10.7 Charpy-V Impact Toughness Analysis ....................................................... 199

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4 Thick Section Narrow Gap Laser Welding of 316L

Stainless Steel .......................................................................... 201

4.1 Introduction ................................................................................................ 201

4.1.1 Expectations ............................................................................................... 203

4.2 Materials .................................................................................................... 204

4.3 Experimental Work and Results ................................................................ 204

4.3.1 Experiments 1-3 ......................................................................................... 205

4.3.2 Experiment 4 .............................................................................................. 207

4.3.3 Experiment 5 .............................................................................................. 209

4.3.4 Experiment 6 .............................................................................................. 211

4.4 Discussion .................................................................................................. 216

4.4.1 Welding Restraint, Welding Distortions and Groove Design .................... 216

4.4.2 Gas Shielding ............................................................................................. 218

4.4.3 Wire Feeding .............................................................................................. 219

4.4.4 Laser Illumination Imaging........................................................................ 221

4.4.5 Pre-Heating and Interpass Temperature Control ....................................... 222

4.5 Conclusions ................................................................................................ 223

5 Thick Section Narrow Gap Laser Welding of Dissimilar

Metals S275 and AISI 316L ................................................... 225

5.1 Introduction ................................................................................................ 225

5.1.1 Expectations ............................................................................................... 227

5.2 Materials and Experimental Methods ........................................................ 228

5.2.1 Materials .................................................................................................... 229

5.2.2 Weld Setup and Parameters ....................................................................... 229

5.3 Experimental Work .................................................................................... 231

5.3.1 Experiments 1-3 ......................................................................................... 231

5.3.2 Experiments 4 and 5................................................................................... 234

5.4 Analysis of Process Characteristics ........................................................... 240

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5.4.1 Weld Pool Behaviour Analysis ................................................................... 240

5.4.2 Symmetry of the Weld Bead ....................................................................... 244

5.4.3 Welding Distortion Analysis ...................................................................... 245

5.5 Analysis of Weldments ............................................................................... 249

5.5.1 Radiographical Analysis of Welding Flaws and ASME IX Acceptance ... 249

5.5.2 Hardness Mapping ...................................................................................... 250

5.5.3 Hardness Line Scan Measurements ............................................................ 253

5.5.4 Microstructural Analysis ............................................................................ 257

5.5.5 EDX Dilution Analysis ............................................................................... 260

5.6 Discussion ................................................................................................... 263

5.7 Conclusions................................................................................................. 267

6 Thick Section Narrow Gap Laser Welding of Dissimilar

Metals SA508 Gr3 Cl2 and 316L ........................................... 271

6.1 Introduction ................................................................................................. 271

6.1.1 Expectations ................................................................................................ 272

6.2 Materials and Experimental Methods ......................................................... 274

6.2.1 Materials ..................................................................................................... 274

6.2.2 Experimental Setup and Welding Parameters ............................................ 274

6.3 Experimental Work and Results ................................................................. 277

6.3.1 Experiment 1 ............................................................................................... 278

6.3.2 Experiment 2 ............................................................................................... 279

6.3.3 Experiment 3 ............................................................................................... 282

6.3.4 Experiment 4 ............................................................................................... 285

6.3.5 Experiment 5 ............................................................................................... 287

6.3.6 Experiment 6 ............................................................................................... 290

6.3.7 Experiment 7 ............................................................................................... 291

6.3.8 Experiment 8 ............................................................................................... 293

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6.4 Analysis of Process Characteristics ........................................................... 295

6.4.1 Welding Process Monitoring ..................................................................... 295

6.4.2 Mechanical Distortion Measurements ....................................................... 297

6.4.3 Effect of Remelting .................................................................................... 299

6.4.4 Temperature Distribution ........................................................................... 300

6.5 Analysis of Weldments .............................................................................. 301

6.5.1 Radiographical Analysis and ASME IX Acceptance ................................ 302

6.5.2 Macrographical Analysis ........................................................................... 303

6.5.2.1 SA2 ............................................................................................................ 304

6.5.2.2 SA5 ............................................................................................................ 310

6.5.3 Hardness Evaluation .................................................................................. 312

6.5.4 Microstructural Analysis ............................................................................ 317

6.5.5 Residual Stress Analysis ............................................................................ 335

6.5.6 Longitudinal Tensile Strength Analysis..................................................... 339

6.5.7 Composite Tensile Analysis ...................................................................... 346

6.5.8 Digital Image Correlation Analysis ........................................................... 350

6.5.9 Impact Toughness Analysis of SA508 Base Material ............................... 352

6.5.10 Impact Toughness Analysis of SA508 Heat Affected Zone ...................... 359

6.6 Discussion .................................................................................................. 367

6.7 Conclusions ................................................................................................ 372

7 General Discussion on Dissimilar Metal NGLW ................. 375

7.1 Introduction ................................................................................................ 375

7.2 Process Characteristics............................................................................... 375

7.3 Geometrical Tolerances ............................................................................. 376

7.4 Oxidation of the Weld Metal ..................................................................... 377

7.4.1 Remedies for Oxidation Issues .................................................................. 378

7.5 Undercutting of Stainless Steel .................................................................. 378

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7.6 Mechanisms Leading to Formation of Lack of Fusion ............................... 381

7.6.1 LoF Due to Oxidation ................................................................................. 381

7.6.2 LoF Due to Undercut .................................................................................. 384

7.6.3 Interpass LoF .............................................................................................. 386

7.6.4 Location of Lack of Fusion ......................................................................... 387

7.7 Laser Beam Reflections from the Weld Pool Surface ................................ 388

7.8 Alignment of the Filler Wire ...................................................................... 391

8 Conclusions .............................................................................. 395

8.1 Introduction ................................................................................................. 395

8.2 Conclusions................................................................................................. 395

8.3 Process Characteristics ............................................................................... 397

9 Future Work ............................................................................ 401

10 Publications ............................................................................. 403

11 References ................................................................................ 405

12 Appendices ............................................................................... 416

64820 words

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List of Figures

Fig. 1.1 Geometry of a full-size mock-up of an RPV safe-end nozzle weld

[12] .......................................................................................................... 38

Fig. 1.2 Main process steps for Laser Multi-Pass Narrow Gap welding [21] ...... 40

Fig. 2.1 Pressurised water reactor. Primary cooling circuit in red, secondary

cooling circuit in blue. U.S. NRC [28] .................................................... 46

Fig. 2.2 Boron and Lithium chemistry control regimes in PWR [32]. ................. 48

Fig. 2.3 Comparison of conventional V-groove to a typical NG J-groove

weld preparation ...................................................................................... 49

Fig. 2.4 Cross sections of ultra-NGLW and conventional GTA welds of

20 mm thick stainless steel [35] .............................................................. 50

Fig. 2.5 Comparison of traditional groove cross-sectional areas to narrow

gap welding in arc welding. Modified from [37] .................................... 50

Fig. 2.6 Three backing support strategies, one permanent and two

removable designs [36] ........................................................................... 52

Fig. 2.7 Principle of NGLW according to Katayama 2013 [41] .......................... 53

Fig. 2.8 Comparison of narrow gap welding geometries and typical groove

dimensions, root support not shown ........................................................ 54

Fig. 2.9 Defocused laser beam in an NGLW groove............................................ 56

Fig. 2.10 The basic groove geometry used by Jokinen and a photograph of an

acceptable root pass. Different groove angles and material

thicknesses were experimented. [17]. ..................................................... 56

Fig. 2.11 Groove geometry for narrow gap welding of 50 mm 316L [20] ............ 61

Fig. 2.12 Experimental setup used by Zhang et al. 2011 [20] ................................ 62

Fig. 2.13 Schematic diagram of NGLW of 17 mm LAS [18] (V1:α=6.74º,

b=4, c=6;V2:α=3.25º, b=2.8, c=5; V3:α=2.3º, b=2.4, c=5) .................... 63

Fig. 2.14 Photo of CRDM (a) and schematic diagram of CRDM assembly (b)

[57]. ......................................................................................................... 66

Fig. 2.15 Typical locations of dissimilar welds in a PWR [13].............................. 67

Fig. 2.16 Geometry and dimensions of Westinghouse AP1000 PWR RPV

safe-end dissimilar metal weld. A – SA508 ferritic steel, B – Alloy

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82 buttering, C – Alloy 182 weld, D- 316L austenitic SS. Note the

cladding of RPV with SS [14].................................................................. 69

Fig. 2.17 Steam generator dissimilar metal weld design [60] ................................. 71

Fig. 2.18 Different pressuriser nozzle dissimilar welds before and after a

repair [60] ................................................................................................. 71

Fig. 2.19 Pressuriser surge line nozzle design and dimensions [61] ....................... 72

Fig. 2.20 Outline of PWR components and materials [66]. .................................... 73

Fig. 2.21 Intergranular corrosion in HAZ of 304 SS containing 0.05 % C

[98]. .......................................................................................................... 82

Fig. 2.22 The effect of carbon content on carbide precipitation [91] ...................... 83

Fig. 2.23 Examples of different types of residual macro- and micro-residual

stress [99] ................................................................................................. 85

Fig. 2.24 Schematic representation of changes of temperature and

longitudinal thermal residual stresses during bead-on-plate welding

[99] ........................................................................................................... 86

Fig. 2.25 Comparison of the average life of dissimilar welds, Type 309

austenitic steel = 1. FM82 both high (H) and low (L) heat input

welds are included [75]. ........................................................................... 90

Fig. 2.26 Schaeffler diagram showing the reduction in the extent of the

ferrite/austenite region for high cooling rate processes [116] .................. 92

Fig. 2.27 WRC-1992 diagram for predicting ferrite content and solidification

mode [98] ................................................................................................. 93

Fig. 2.28 “Normal” grain growth with Type I grain boundaries and the grain

growth exhibiting Type II boundaries in a dissimilar metal weld

[98] ........................................................................................................... 94

Fig. 2.29 Dissimilar metal fusion line and the formation mechanism of Type

II boundaries [75] ..................................................................................... 95

Fig. 2.30 A diagram of the zones in the HAZ of a 0.15 wt.% C steel [119] ........... 96

Fig. 2.31 Another classification of the low alloy steel HAZ subzones [120]. ........ 96

Fig. 2.32 Boundaries developing in austenitic weld metals [75] .......................... 101

Fig. 2.33 Migrated grain boundary (MGB), Solidification grain boundary

(SGB) and Solidification subgrain boundaries (SSGB) in austenitic

Filler Metal 52 weld metal [75]. ............................................................ 102

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Fig. 2.34 Examples of weld solidification cracking in Filler Metal 52M

(ERNiCrFe -7A) dissimilar weld overlays on a) carbon steel A36,

and b) stainless steel Type 304L [75].................................................... 103

Fig. 2.35 Hot cracking propagated to the surface of narrow groove laser

hybrid welding of AISI 316L-IG (ITER-grade) using Thermanit

19/15 filler [130] ................................................................................... 104

Fig. 2.36 Ductility as a function of temperature and ductility dip [75] ................ 106

Fig. 2.37 DDC in multi-pass weld using Filler Metal 52. Arrows showing

severe DDC and recrystallization along migrated grain boundaries

[75] ........................................................................................................ 107

Fig. 2.38 The principle of liquation cracking [98] ............................................... 108

Fig. 2.39 Weld metal contraction tearing PMZ of 7075 aluminium welded

with filler 1100 [134] ............................................................................ 108

Fig. 2.40 Sketch of critical locations in the ferritic/austenitic laser welds

[148] ...................................................................................................... 110

Fig. 2.41 HAZ of GTA weld of AISI 1018 steel [98] ......................................... 113

Fig. 2.42 HAZ microstructure of 1018 steel by a high-power CO2 laser.

Magnification of (A)–(D) 415x and of (E) 65x. B, High carbon

martensite [98]....................................................................................... 114

Fig. 2.43 The parameters of laser welding with filler wire [154]. ........................ 117

Fig. 2.44 Engineering stress-strain diagram [157] ............................................... 120

Fig. 2.45 Ductile to brittle transition region. Energy shelves and fracture

descriptions [158]. ................................................................................. 121

Fig. 2.46 Schematic indicative of the approximate current capabilities of

various techniques. Destructive techniques shaded grey [163]............. 127

Fig. 2.47 The intact sample with original stresses [166] ...................................... 128

Fig. 2.48 The cut sample with deformation induced by the residual stresses

[166] ...................................................................................................... 128

Fig. 2.49 The analytically flattened surface with residual stress map [166] ........ 128

Fig. 2.50 Appendix I of ASME IX defining the acceptable appearance of

rounded radiographic indications [176] ................................................ 131

Fig. 2.51 Example of BS EN ISO 13919-1. The maximum height of undercut... 132

Fig. 2.52 Charpy V-notch impact toughness sample according to BS EN ISO

148 ......................................................................................................... 132

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Fig. 2.53 Example of an impact test curve ............................................................ 133

Fig. 2.54 Principle of the Vickers hardness test .................................................... 134

Fig. 2.55 ASTM E8 rectangular tensile test sample [182] ........................................ 135

Fig. 2.56 Stress-strain diagram [182] .................................................................... 136

Fig. 3.1 Dimensions and design of NGLW samples. .......................................... 146

Fig. 3.2 5 mm parallel groove contraction, TV4 ................................................. 147

Fig. 3.3 The 4° V-groove design for countering groove contraction .................. 148

Fig. 3.4 The laser welding setup during wire feeder nozzle prototype trials ...... 152

Fig. 3.5 Sample TV1 restrained by one pair of Carver buttress clamps.............. 154

Fig. 3.6 Strongback welding arrangement for TV4. Note copper shielding

gas nozzles on the table and separate brackets to contain the gas

atmosphere for the last passes. ............................................................... 155

Fig. 3.7 Strongback welding restraint arrangement for TV4 viewed from

below. Note the machined slot to accommodate the backing plate

and the extensive manual GTAW-welding required.............................. 156

Fig. 3.8 42 kN Lenskes welding clamps restraining a dissimilar metal weld

sample DS4 on an 80 mm thick Lenskes T-groove steel welding

table. Note the location of the clamping points near the edges of the

sample, the pre-heating blankets and run-in and run-out blocks to

guide the shielding gas. .......................................................................... 157

Fig. 3.9 TecArc F4 wire feeder and the wire conduit (red). ................................ 158

Fig. 3.10 Jetline Engineering 9600 wire feeder unit mounted on top of the

robot ....................................................................................................... 159

Fig. 3.11 Jetline engineering 9600 wire feeder control unit .................................. 159

Fig. 3.12 Two examples of different experimental wire feeder versions. A) A

simple feeder pipe attached to the welding head (arrow) and B) the

MTRL NGLW nozzle which was used in welding the TV-series

stainless steel tests .................................................................................. 161

Fig. 3.13 Final wire feeder nozzle bracket (silver) and 300 mm stage (black)

used ........................................................................................................ 162

Fig. 3.14 The improved 5 mm diameter wire feed nozzle .................................... 164

Fig. 3.15 Different gas shoe designs. A) A version constructed for first trials.

B) A proposed design to alleviate dust accumulation issues. ................ 165

Fig. 3.16 Blade shielding gas and wire feed design .............................................. 166

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Fig. 3.17 Shielding gas setup used during welding of TV4 set up with

stationary nozzles aided by steel blocks to guide the gas flow. The

nozzles were raised according to the progression of the weld. ............. 167

Fig. 3.18 Final gas shielding and wire feed setup ................................................ 169

Fig. 3.19 Cross-hairs aligned to the centre of the laser spot seen during a trial

weld 1) Top of base materials, 2) Welding groove sidewalls, 3)

Backing plate. ........................................................................................ 171

Fig. 3.20 Measurement of the pre-heat and interpass temperature ....................... 172

Fig. 3.21 Punch mark locations for indirect distortion measurements ................. 173

Fig. 3.22 Conducting the indirect gap measurement between punch marks. ....... 174

Fig. 3.23 200 measurement line pairs for butterfly analysis on a 3D scanned

model of sample TV4. Note the omitted central area to avoid the

effect of local deformation near the weld.............................................. 175

Fig. 3.24 Laser illumination imaging setup on the welding head.

Experimental LED light source pictured, Cavilux laser aperture

unit similar in size ................................................................................. 177

Fig. 3.25 A frame from Laser illumination imaging video. 3rd

pass of test

DS1. 1) Top of the stainless steel plate, 2) Top of the ferritic steel

plate, 3) Weld bead, 34 mm below the previous, 4) Wire feed

nozzle and 5) welding groove side wall (dark area). ............................. 178

Fig. 3.26 A frame from the coaxial camera system video. 1st pass of test

SA5. 1) Welding groove sidewalls, 2) Filler wire and 3) wetting of

the sidewall ............................................................................................ 179

Fig. 3.27Laser power calibration chart .................................................................... 181

Fig. 3.28 Typical sample for microstructural and hardness analysis, thickness

varying from 2 mm to 10 mm ............................................................... 186

Fig. 3.29 High level overview of contour method analysis .................................. 190

Fig. 3.30 Contour cutting jig lower half. Top beams removed to reveal the

individually adjustable pads which ensure restrainment avoiding

inducing external stresses. The cut was performed between the

beams along the line drawn in red. ........................................................ 192

Fig. 3.31 Contour cut sample, cross-sectional view. Contour cut made first,

with ligaments in place to support the sample. ..................................... 193

Fig. 3.32 Sample of SA5 mounted to a jig undergoing the contour scanning ...... 193

Fig. 3.33 Profilometry raw data example, (µm) ................................................... 194

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Fig. 3.34 Averaged data with some residual noise, not in scale ............................ 194

Fig. 3.35 Profile outline silhouette spline with seeds (circles) for mesh

generation ............................................................................................... 195

Fig. 3.36 Finite element mesh (left) and the 3D extrusion (single material

sample shown)........................................................................................ 196

Fig. 3.37 Contour method residual stress map plotted using dense knot

spacing ................................................................................................... 196

Fig. 3.38 Contour method residual stress map plotted using less dense knot

spacing ................................................................................................... 196

Fig. 3.39 Tensile test sample ASTM E8/E8M. Longitudinal and composite

similar..................................................................................................... 197

Fig. 3.40 DIC setup for tensile testing. 1) cameras, 2) flashlights and 3)

tensile test sample. ................................................................................. 198

Fig. 3.41 Tensile test samples used for DIC by spraying a speckle pattern .......... 199

Fig. 3.42 Charpy-V impact toughness test coupon as per ASTM E23 and ISO

148-1 ...................................................................................................... 199

Fig. 4.1 Cross-section of TV1 showing a successful weld with no fusion

issues and good penetration to the backing plate ................................... 207

Fig. 4.2 Cross-section of TV2 showing asymmetry and poor capping due to

misalignment of the sample. .................................................................. 207

Fig. 4.3 Average groove width per pass, measured directly between the top

corners of the welding groove, TV4. ..................................................... 209

Fig. 4.4 Macrographical cross-section showing no issues with fusion, cracks

or porosity TV5. ..................................................................................... 210

Fig. 4.5 A LII video frame, showing wire, nozzle and the welding pool.

Note the stable welding process. Stainless steel similar metal weld

TV5 ........................................................................................................ 210

Fig. 4.6 The generation of butterfly distortion, 3D scan data analysis, test

TV6 ........................................................................................................ 211

Fig. 4.7 Groove top contraction development of a 5° V-groove 40 mm thick

316L similar metal weld. See the minimum staying over 5 mm.

Direct measurement was not applicable for last 2 passes due to

melting of the groove top edges, test TV6. ............................................ 213

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Fig. 4.8 Laser Illumination Image of pass 1 in a similar metal NGLW. 1)

Uneven wetting of side walls due to too small spot size, oxidation

of weld bead, 2) Wire nozzle melting due to reflected laser beam,

3) Wire contamination by particles of nozzle and vibration (in

video material), 4) Smoke generation due to excess laser power ......... 221

Fig. 4.9 12th

pass of a similar metal NGLW. 1) Good smooth fusion, 2) Gap

contracting excessively, laser beam heats the corners of the groove

prematurely, 3) Wire alignment off-centre and 4) Accumulated

oxides floating on top of the melt pool ................................................. 221

Fig. 5.1 Initial 4° welding groove design for S275 dissimilar metal welds ....... 230

Fig. 5.2 Experimental setup for a DMW aligned for the first pass. 1) Laser

head, 2) coaxial wire feeding and shielding gas nozzle, 3)

stationary shielding gas nozzle, 4) welding clamps and 5) high-

speed camera lens. Laser illumination laser is just outside the

picture on right. ..................................................................................... 231

Fig. 5.3 DS1 dual etched with Nital and oxalic acid. The hardness mapping

indentations also visible ........................................................................ 232

Fig. 5.4 The nozzle with added trailing nozzles as used for DS4....................... 235

Fig. 5.5 19th

pass of DS4, note undercutting at 130-170 mm. ............................ 237

Fig. 5.6 Close-up of undercut at 19th

pass of DS4, note lack of fusion to the

stainless steel (arrows) .......................................................................... 237

Fig. 5.7 Finished weld DS4. Note the mainly good fusion due to two

parallel capping passes. ......................................................................... 237

Fig. 5.8 First pass of DS5. Note smooth, symmetrical wetting, oxidation and no

irregularities .......................................................................................... 239

Fig. 5.9 Close-up of the 8th

pass of DS5. Note unacceptable irregular oxidation

and wetting. ........................................................................................... 239

Fig. 5.10 The finished weld DS5 prior to wire brushing. ..................................... 240

Fig. 5.11 Laser illumination imaging image, note the absence of glare. DS2,

second pass. ........................................................................................... 241

Fig. 5.12 Formation of lack of fusion / inadequate wetting (arrow) and

oxidation of the weld bead viewed by laser illumination imaging ....... 242

Fig. 5.13 Coaxial camera ambient light image. DS4, first pass. .......................... 243

Fig. 5.14 DS4 second pass pictured from the end of the welding groove. Note

the level surface of the bead with no inclination to either metal and

the sharp corners instead of smooth wetting. ........................................ 245

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Fig. 5.15 Welding groove contraction measured directly and calculated from

indirect measurements. Test DS2. Error bars represent min and

max values measured. ............................................................................ 246

Fig. 5.16 Welding groove top width of DS series of welds using 4° V-

groove. Note the groove contracting to below 5 mm. Average of

three tests, DS1-DS3. Error bars represent min and max values

measured. ............................................................................................... 247

Fig. 5.17 Welding groove top width of DS series of welds using 5° V-

groove. Note the groove width staying above 5 mm until the end.

Average of tests DS4 and DS5. Error bars represent min and max

values measured. .................................................................................... 248

Fig. 5.18 Hardness map of DS1. 69 (HV1) S275 left, FM52 middle, 316L

right. Fusion lines in red. A main observation is the strain

hardening of the austenitic materials. .................................................... 251

Fig. 5.19 Hardness map of DS1 69 (HV0.3. Approximate fusion lines in red.

Note the narrow area of high hardness at the S275 FL. ......................... 252

Fig. 5.20 Hardness line plot across the weld, second to top pass. Note the

narrow area of high hardness in S275. Fusion lines marked on the

graph. Sample DS1 69. .......................................................................... 254

Fig. 5.21 Hardness test grid and locations. Note top bead being considerably

wider than previous passes to ensure smooth capping. Sample DS2

116. ......................................................................................................... 255

Fig. 5.22 Hardness of S275 HAZ at different states of multipass tempering........ 256

Fig. 5.23 Hardness of the S275 HAZ. Close up near the fusion line. ................... 257

Fig. 5.24 S275 HAZ of a tempered filling pass, BM-tempered-ICHAZ-

FGHAZ-CGHAZ-WM........................................................................... 257

Fig. 5.25 Unaffected S275 base material .............................................................. 258

Fig. 5.26 Grain structure at a tempered filling pass .............................................. 258

Fig. 5.27 Untempered CGHAZ at the top pass ..................................................... 259

Fig. 5.28 DS1 weld metal microstructures ............................................................ 259

Fig. 5.29 S275 HAZ banding due to multi-pass tempering................................... 260

Fig. 5.30 EDX line plot of the weld metal and fusion lines. Note the rapid

transition from ferritic BM to WM composition ................................... 261

Fig. 5.31 EDX line plot for top bead at the S275 fusion line. The transition in

the composition is prompt. ..................................................................... 262

19

Fig. 5.32 EDX line plot for top bead at the 316L fusion line. .............................. 263

Fig. 6.1 Experimental setup for SA1 pedigree metal weld. Thermocoupling,

stationary gas shielding arrangement, pre-heating blankets and

Lenskes clamps. .................................................................................... 275

Fig. 6.2 First application of the commercial wire nozzle holder, as used for

SA1. ....................................................................................................... 276

Fig. 6.3 Wire feed setup with a linear stage for height adjustment and tri-

blade gas nozzle arrangement. As used from SA2 onwards. ................ 277

Fig. 6.4 The welding setup for SA1. SA508 LAS nearer the camera. ............... 278

Fig. 6.5 The gas and wire nozzle used for SA2 .................................................. 280

Fig. 6.6 SA2 first pass. The smooth wetting, symmetrical concave bead

shape and good gas shielding were characteristic for the first passes

of all Stage III welds. A and B are separate blocks to guide the

shielding gas. ......................................................................................... 282

Fig. 6.7 The welding setup for SA3 with welded run in and –out brackets.

Welding direction down. Stainless steel on the left. ............................. 283

Fig. 6.8 3rd

pass of SA3. Note the oxide-free weld extending to the run on

bracket at start. ...................................................................................... 283

Fig. 6.9 Smooth capping of SA4 with just some minor lack of fusion type

defects.................................................................................................... 287

Fig. 6.10 Symmetrical capping of SA4. ............................................................... 287

Fig. 6.11 First pass of SA5. Smooth wetting and bead surface with low

oxidation. ............................................................................................... 289

Fig. 6.12 12th

pass of SA5. Weld quality very similar to pass 1. Slight

colouration of the SA508 surface due to heating starting to appear. .... 289

Fig. 6.13 Finished weld SA5 uncleaned, as welded. Note lack of oxidation. .......... 289

Fig. 6.14 SA6, 11th

pass, low oxidation, smooth bead and good wetting of the

sidewalls ................................................................................................ 291

Fig. 6.15 Finished weld SA6. Minor underfilling (arrow) ................................... 291

Fig. 6.16 Collapse of the sidewall (arrow) SA7, 19th

pass. .................................. 293

Fig. 6.17 SA7 finished, with fusion flaws. ........................................................... 293

Fig. 6.18 A depression in the 14th

pass, SA8. ....................................................... 295

Fig. 6.19 SA1, first pass. LII snapshot. Good wetting and alignment.

Excessive oxidation. .............................................................................. 296

20

Fig. 6.20 SA1 13th

pass. LII snapshot. Groove contracting excessively, laser

beam melting the stainless steel prematurely, very severe oxidation. ... 296

Fig. 6.21 Welding groove top width of SA1 using 4° V-groove. Note gap

contracting below 5 mm. Error bars represent min and max values

measured. ............................................................................................... 298

Fig. 6.22 Welding groove top width of SA series using 5° V-groove. Note

gap remaining over 5 mm until the end of the weld. Average of 6

tests. Error bars represent min and max values measured. .................... 299

Fig. 6.23 2nd

pass SA7, poor wetting and fusion of the first 75 mm, see

arrow. ..................................................................................................... 300

Fig. 6.24 3rd

pass SA7, a re-melt applied to the previous pass has reduced the

length of the poorly fused area by 50 %, see arrow. .............................. 300

Fig. 6.25 Thermocoupling of SA1, top side. Bottom similar. ............................... 301

Fig. 6.26 Photomontage of Nital (for SA508) and ammonium persulfate

(FM52) etched weld cross-section. Sample SA5 161. ........................... 304

Fig. 6.27 Cross-section of an ASME IX rejected dissimilar metal weld

showing various fusion defects. Nital etch. 316L fusion line

superimposed as dotted line. Sample SA2 98 ........................................ 307

Fig. 6.28 Cross-section of an ASME IX accepted dissimilar metal weld

showing near-perfect fusion. Nital etch. Fusion line superimposed

in red. Sample SA5 161 ......................................................................... 311

Fig. 6.29 Hardness map of SA5 157. SA508 on the left, FM52 middle and

316L right. Approximate fusion line in red. (HV 0.3) ........................... 313

Fig. 6.30 Detail of the high hardness HAZ in SA508. Test point grid 200 x

200 µm. Sample SA5 157 ...................................................................... 315

Fig. 6.31 Hardness map with an alternate scale to emphasize strain

hardening. Sample SA5 157................................................................... 316

Fig. 6.32 Outline of top of the weld SA5. ............................................................. 317

Fig. 6.33 SA508 Base material, tempered martensite (light). Carbon

precipitated as cementite particles (grey) austenite grain boundaries

(dark), SA5 159 Nital etch. .................................................................... 318

Fig. 6.34 Martensitic (dark) HAZ with bainite (light) adjacent to FL of the

top pass, SA5 159 Nital etch. ................................................................. 319

Fig. 6.35 Martensitic/bainitic structure at the area of highest hardness,

200 µm perpendicular) from FL, top pass, SA5 159 Nital etch. ............ 319

21

Fig. 6.36 Transition from CGHAZ (right) to FGHAZ (left), approx. 1 mm

from FL, top pass, SA5 159 Nital etch. ................................................. 320

Fig. 6.37 FGHAZ, approx. 1.5 mm from FL top pass, SA5 159 Nital etch. ........ 321

Fig. 6.38 ICHAZ, approx. 2 mm from FL, top pass, SA5 159 Nital etch. ........... 321

Fig. 6.39 Vertical stitched image of SA508 HAZ. SA5 159 Nital etch. .............. 322

Fig. 6.40 Multi-pass tempered CGHAZ, 7th

to last pass, 15 mm from top,

SA5 161 Nital etch. ............................................................................... 323

Fig. 6.41 Austenitic weld metal at 316L FL. 1) SSGB and 2) SGB. SA5 161

Ammonium persulfate etch. .................................................................. 324

Fig. 6.42 Austenitic weld metal at 316L. 1) SSGB, 2) SGB and 3) MGB.

SA5 161 Ammonium persulfate etch. ................................................... 325

Fig. 6.43 FM52 weld metal at SA508 FL. 1) SSGB, 2) MGB and 3)

martensitic region. SA5 161 Ammonium persulfate etch. .................... 325

Fig. 6.44 Suspected DDC (arrow) at MGB near 316L fusion line. SA5 161

Ammonium persulfate etch. .................................................................. 327

Fig. 6.45 Suspected DDC in weld metal near SA508 fusion line with possible

recrystallization (arrow) SA5 161 Ammonium persulfate etch. ........... 328

Fig. 6.46 SA508 sidewall LoF at 2nd

to last pass. SA5 159 Nital etch. ................ 329

Fig. 6.47 Interpass LoF in weld metal at pass no 18, SA5 161 Ammonium

persulfate etch. ...................................................................................... 329

Fig. 6.48 Pore near fusion boundary of SA508. Note over-etched SA508.

SA5 161 Ammonium persulfate etch. ................................................... 330

Fig. 6.49 316L base material. Austenitic grain structure with δ-ferrite

stringers in rolling direction. SA5 159. ................................................. 331

Fig. 6.50 Local cluster of large grain in 316L base material. SA5 159 Oxalic

acid etch................................................................................................. 332

Fig. 6.51 Hot crack in 316L HAZ near the top SA5. Note heavily over-

etched weld metal SA5 159 Oxalic acid etch. ....................................... 333

Fig. 6.52 Intergranular ferrite formation at fusion line. SA5 159 Oxalic acid

etch. ....................................................................................................... 334

Fig. 6.53 Skeletal δ-ferrite formation at fusion boundary. SA5 159 Oxalic

acid etch................................................................................................. 334

Fig. 6.54 Sample SA5 showing the contour cut ................................................... 335

Fig. 6.55 SA2 residual stress map ........................................................................ 337

22

Fig. 6.56 SA5 residual stress map ......................................................................... 337

Fig. 6.57 Locations of longitudinal tensile test coupons numbered CPL1 to

CPL15 .................................................................................................... 340

Fig. 6.58 Longitudinal tensile test coupons; CPL1-3 316L BM, CPL4-6 316L

HAZ, CPL7-9 FM52 WM, CPL10-12 SA508 HAZ, CPL13-15

SA508BM .............................................................................................. 341

Fig. 6.59 Longitudinal tensile test coupons after testing ....................................... 342

Fig. 6.60 316L BM stress-strain curves, test coupons CPL1 to 3 ......................... 343

Fig. 6.61 316L HAZ stress-strain curves, test coupons CPL4 to 6 ....................... 343

Fig. 6.62 Filler Metal 52 weld metal stress-strain curves, test coupons CPL7

to 9.......................................................................................................... 344

Fig. 6.63 SA 508 Gr3 Cl2 HAZ stress-strain curves, test coupons CPL10 to

12 ............................................................................................................ 345

Fig. 6.64 SA 508 Gr3 Cl2 BM stress-strain curves, test coupons CPL13 to 15 ... 345

Fig. 6.65 Charpy-V coupon and composite tensile test coupons CPT1-18

locations in sample SA5 ......................................................................... 346

Fig. 6.66 Composite (CPT1-9) and transverse 508 base material (CPT10-18)

test coupons, 316L steel top ................................................................... 347

Fig. 6.67 Composite test stress-strain curves, test coupons CPT1 to 3. ................ 348



Fig. 6.70 Transverse 508 BM test stress-strain curves, test coupons CPT10 to

18. ........................................................................................................... 350

Fig. 6.71 Strain localisation images for composite tensile test SA5, coupon 5.

Stainless steel at bottom. ........................................................................ 351


Stainless steel at bottom. ........................................................................ 352

Fig. 6.73 Deformation of the fusion zone in SA5, coupon 8, arrow ..................... 352

Fig. 6.74 SA508 base material Charpy-V impact energy, J .................................. 354

Fig. 6.75 SA508 Base material fracture surfaces. A – D, G and H brittle

fracture, E transition and I ductile facture .............................................. 355

Fig. 6.76 Ductile fracture area in SA508 base material, 400x, 1000x and

5000x magnification, sample F .............................................................. 357

23

Fig. 6.77 Brittle fracture area in SA508 base material, 400x, 1000x and

5000x magnification, sample F ............................................................. 358

Fig. 6.78 SA508 HAZ Charpy-V impact energy, J .............................................. 359

Fig. 6.79 SA508 HAZ fracture surfaces. A and B brittle, C to I ductile. ............. 360

Fig. 6.80 SEM images of ductile fracture. 400x, 1000x and 5000x

magnification, Sample G ....................................................................... 362

Fig. 6.81 SEM image of Sample C, low magnification ........................................ 363

Fig. 6.82 SEM images of Sample C. 400x, 1000x and 5000x magnification,

ductile area ............................................................................................ 364

Fig. 6.83 SEM images of Sample C. 400x, 1000x and 5000x magnification,

brittle area .............................................................................................. 365

Fig. 6.84 SEM images of HAZ Sample A, three different locations A, B and

C. 400x, 1000x and 5000x magnification. ............................................ 366

Fig. 7.1 Near symmetrical undercut (arrow) due to re-melting of a pass in

an incomplete stainless steel weld. Note also overheated top

corners of the welding groove. Oxalic acid etch. .................................. 379

Fig. 7.2 Local widening of the welding groove caused by irregular collapse

of asymmetric undercut in the stainless steel, sample DS4................... 380

Fig. 7.3 LII image of DS2, pass 11. The weld bead generated is flat. Note

heavy oxidation. .................................................................................... 382

Fig. 7.4 Desirable and undesirable bead shapes ................................................. 382

Fig. 7.5 Optimal penetration between passes and reliable fusion ...................... 383

Fig. 7.6 Consequences of inadequate wetting increasing the risk of lack of

fusion ..................................................................................................... 384

Fig. 7.7 Undercut induced lack of fusion principle and two possible bead

shapes .................................................................................................... 385

Fig. 7.8 The effect of undercut induced lack of fusion to the following

welding pass .......................................................................................... 386

Fig. 7.9 Reflection of the laser beam towards the direction of welding ............. 388

Fig. 7.10 Overheating of wire nozzle (arrow) due to reflections. DS2 pass 2,

38 s from the start. ................................................................................. 389

Fig. 7.11 Copper debris (two right arrows) from the damaged nozzle (left

arrow) being carried to the melt pool by the filler wire. DS2 pass 2,

250 ms after Fig. 7.10............................................................................ 390

24

Fig. 7.12 The detrimental effects of asymmetric reflections ................................. 391

Fig. 7.13 Droplet formation due to filler wire inadvertently hitting the laser

beam above the melt pool. Oxidation of the droplet visible in video

material. Maximum droplet size shown, just before gravity pulls

the droplet down to the meltpool. DS2, pass 1. ..................................... 392

Fig. 7.14 Droplet at the moment of being absorbed by the meltpool, 40 ms

after Fig. 7.13. DS2, pass 1. ................................................................... 393

25

List of Tables

Table 1.1 Power reactors operating in the United Kingdom in 2017 [2] ................ 37

Table 2.1 Parameters to achieve 20 mm NGLW of stainless steel [17] .................. 59

Table 2.2 Per pass parameters for 20 mm NGLW of AISI 304LN [17] ................. 59

Table 2.3 Parameters for 35 mm NGLW of AISI 304LN [17] ............................... 60

Table 2.4 Per pass parameters for 35 mm NGLW of AISI 304LN [17] ................. 60

Table 2.5 Per pass parameters for Q235 LAS used by Yu et al. 2013 [18] ............ 63

Table 2.6 Chemical composition of A508 Grade 3 Class 2 steels used in a

study [67] wt.-% ...................................................................................... 74

Table 2.7 Mechanical properties of SA508 Gr3 Cl2 steel[67]. ............................... 74

Table 2.8 Standard chemical composition of AISI 316 and AISI 316L wt.-%

[71, 72] .................................................................................................... 75

Table 2.9 Mechanical properties of AISI 316L annealed [71]. ............................... 75

Table 2.10 Chemical composition 316L wt.-% [74] ................................................. 75

Table 2.11 Chemical compositions of Ni filler alloys wt.-% [58] ............................ 77

Table 2.12 Mechanical properties of Alloy 52 and Alloy 82 fillers [58] .................. 77

Table 2.13 Chemical composition of Filler Metal 82 [78] ........................................ 78

Table 2.14 Chemical composition of Filler Metal 52 wt.-% [80] ............................. 78

Table 2.15 Chemical composition of Filler Metal 52M wt.-% [87].......................... 79

Table 2.16 Chemical composition of Filler Metal 52MSS wt.-% [89] ..................... 79

Table 2.17 Basic types of EAC [33].......................................................................... 88

Table 2.18 Cracking types in Alloy 52 and 52M weldments [128] ........................ 100

Table 2.19 Approximate limit of diluting elements in Fe – Ni – Cr welds.

Adapted from [131] ............................................................................... 105

Table 2.20 Laser weldability of binary metal combinations. E = excellent, G =

Good, F = Fair, P = Poor, * = no data available [149] ......................... 111

Table 2.21 Critical parameters in laser welding [150] ............................................ 115

Table 2.22 Visual examinations after welding [161] .............................................. 123

26

Table 2.23 Comparison of three strain mapping methods [163] .............................. 126

Table 2.24 Standards used for evaluation of the weldments .................................... 130

Table 3.1 Welding programme stages and corresponding materials used ............. 142

Table 3.2 EDS analysis of the 316L stainless steel used throughout the study,

wt.% [35] ................................................................................................ 143

Table 3.3 Nominal chemical composition of ER 316L, wt.% [184]...................... 143

Table 3.4 Mechanical properties of materials used [185] [80] [186] .................... 143

Table 3.5 Chemical properties of the base materials used [185-187] .................... 143

Table 3.6 Chemical properties of the filler material used [80] .............................. 144

Table 3.7 Mechanical properties of materials used in the SA series of

experiments [185], [80], [188] ............................................................... 144

Table 3.8 EDS analysis of the 316L stainless steel used, wt.% [35] ..................... 144

Table 3.9 Analysis of SA508 Gr3 Cl2 material used, wt.%, Appendix IV ........... 145

Table 3.10 Analysis of Inconel Alloy FM52 material used, wt.%, Appendix V .... 145

Table 3.11 Base parameters for NGLW ................................................................... 150

Table 3.12 Typical values of different parameters and measurements with

errors ...................................................................................................... 184

Table 3.13 Etchants, compositions and applications ............................................... 187

Table 3.14 Minimum distances to prevent work hardening artefacts in hardness

testing. d = indent diagonal [180, 181] .................................................. 188

Table 3.15 Contour method workflow ..................................................................... 191

Table 4.1 Stage I of thick section similar metal narrow gap laser welds in

316L ....................................................................................................... 202

Table 4.2 Welding parameters TV1 to TV3 ........................................................... 205

Table 4.3 Groove contraction measurements TV1 to TV3 .................................... 206

Table 4.4 Welding parameters TV4, all passes. ..................................................... 208

Table 4.5 Groove contraction measurements, test TV6 ......................................... 213

Table 4.6 The preheating and welding parameters for similar metal welding,

test TV6. ................................................................................................. 215

27

Table 4.7 Temperature and thickness development of TV6 similar metal

316L weld .............................................................................................. 216

Table 5.1 The Stage II of dissimilar thick section narrow gap laser welds

using S275, Alloy 52 and 316L ............................................................. 227

Table 5.2 Welding parameters DS1. ..................................................................... 232

Table 5.3 Welding parameters DS2. ..................................................................... 233

Table 5.4 Visual examination of welds DS1 to DS3 ............................................. 234

Table 5.5 Welding parameters and notes DS4 ...................................................... 236

Table 5.6 Welding parameters and notes DS5 ...................................................... 238

Table 5.7 Visual examination of welds DS4 and DS5 .......................................... 240

Table 5.8 Summary of radiographical acceptance reports of DS series of

welds...................................................................................................... 250

Table 5.9 Dilution of weld DS1 ............................................................................ 261

Table 6.1 Stage III of dissimilar metal narrow gap laser welds, SA508 to

316L using Alloy 52 .............................................................................. 272

Table 6.2 Welding parameters and notes SA1. ..................................................... 279








Table 6.10 Thermocouple numbering SA1 ............................................................. 301

Table 6.11 Stage III weld quality visual inspection ................................................ 302

Table 6.12 Results and interpretation of radiographical analysis of Stage III

welds...................................................................................................... 302

Table 6.13 Record of basic welding parameters, fusion quality and undercut in

weld SA2 ............................................................................................... 306

Table 6.14 Record of basic parameters and fusion quality in test SA5................... 310

28

Table 6.15 Poisson’s ratios and coefficients of elasticity for PCC materials .......... 336

Table 6.16 Longitudinal tensile test coupon locations in sample SA6 by

location ................................................................................................... 340

Table 6.17 Longitudinal tensile properties. Sample SA6 ........................................ 342

Table 6.18 Transverse tensile test locations ............................................................. 346

Table 6.19 UTS and elongation of transverse tensile tests, SA5, average of 9

coupons .................................................................................................. 347

Table 6.20 Variation of transverse tensile properties according to location of

the sample, SA5 ..................................................................................... 348

Table 6.21 Impact energies and proportion of ductile fracture in SA508

Charpy-V tests........................................................................................ 356

Table 6.22 Impact energies and proportion of ductile fracture in SA508 HAZ

Charpy-V tests........................................................................................ 361

29

List of Symbols

A, γ Austenite

Creq Chromium equivalent

CN Maximum allowed force on welding clamp

D Diagonal of the hardness measurement indent

d0 Original gap measurement pre-welding

dn Measurement at pass n

E Modulus of elasticity

e Strain

ef Strain to fracture

F, α Ferrite

FZ Filler wire height relative to sample surface

G0 Original gap pre welding

Gn Gap at pass n

HV, Hv Vickers Number hardness

M Martensite

NC Number of pairs of clamps

Nieq Nickel equivalent

Pcm Carbon equivalent JWES

PH Plate height

PL Plate length

PW Plate width

S Stress

UTS, σTS (Ultimate) tensile strength

Wy, Wx, WZ Filler wire position coordinates relative to laser beam

αW Filler wire feed angle from horizontal

ν Poisson’s ratio

σy Yield strength

30

List of Abbreviations

AISI American Iron and Steel Institute

ASCII American Standard Code for Information Interchange

ASME The American Society of Mechanical Engineers

ASTM American Society for Testing and Materials

BCC Body Centric Cubic (crystal lattice)

BM Base Material

BWR Boiling Water Reactor

CE Carbon Equivalent

CF Corrosion Fatigue

CMM Coordinate Measuring Machine

CO2 Carbon Dioxide

Cr Chromium

CRDM Control Rod Drive Mechanism

CTE Coefficient of Thermal Expansion

DBA Design Basis Accident

DBTT Ductile Brittle Transformation Temperature

DDC Ductility Dip Cracking

DM Dissimilar Metal

DMW Dissimilar Metal Welding

EAC Environmentally Assisted Cracking

EDM Electrical Discharge Machining

EDX, EDS Energy-dispersive X-ray Spectroscopy

EPRI Electric Power Research Institute

FAC Flow Assisted Corrosion

FCC Face Centric Cubic (crystal lattice)

31

Fe Iron

GTAW Gas Tungsten Arc Welding

HAZ Heat Affected Zone

HSLA High Strength Low Alloyed (steel)

ISO International Organization for Standardization

JWES Japanese Welding Engineering Society

LBZ Local Brittle Zone

LAS Low Alloyed Steel

LC Liquation Cracking

LCF Low Cycle Corrosion Fatigue

LII Laser Illumination Imaging

LoB Lack of Bonding

LOCA Loss of Coolant Accident

LoF Lack of Fusion

MGB Migrated Grain Boundary

MTRL Manufacturing Technology Research Laboratory

Nd:YAG Neodymium Yttrium Aluminium Garnet

NDT Non-Destructive Testing

NGLW Narrow Gap Laser Welding

Ni Nickel

NNUMAN New Nuclear Manufacturing Programme

NPP Nuclear Power Plant

PCC Primary Coolant Circuit

PMZ Partially Melted Zone

PVH Pressure Vessel Head

PWHT Post Weld Heat Treatment

PWR Pressurised Water Reactor

32

PWSCC Primary Water Stress Corrosion Cracking

RPV Reactor Pressure Vessel

SCC Stress Corrosion Cracking

SEM Surface Electron Microscopy

SICC Strain-Induced Corrosion Cracking

SGB Solidification Grain Boundary

SS Stainless Steel

SSGB Solidification Subgrain Boundary

USE Upper Shelf Energy

UTS Ultimate Tensile Strength

WM Weld Metal

WPS Welding Procedure Specification

33

Abstract

Dissimilar metal welding (DMW) between thick section low alloyed and stainless

steels is essential in pressurised water reactor (PWR) construction. This study

explored the potential of narrow gap laser welding (NGLW) to improve the quality

and manufacturability of these welds. It concentrates on DMW joints located in the

primary cooling circuit, where austenitic stainless steel pipes are connected to forged

low alloy pressure vessel nozzles.

Current welding processes for this application are manual metal arc (MMA) and

narrow gap gas tungsten arc (NG-GTA) welding, which are slow and inefficient.

They require large amounts of filler material to be deposited and generate

considerable residual stresses. The residual stresses contribute to stress corrosion

cracking (SCC), which has been found a major issue to the longevity and reliability

of PWR’s. NGLW has the potential to reduce the amount of filler material required

and has been shown to reduce the detrimental residual stresses.

In this study NGLW was applied for welding SA508 Gr3 Cl2 low alloyed steel with

AISI 316L austenitic stainless steel using Inconel Alloy 52 filler metal up to 40 mm

thickness. This thesis is the first time that dissimilar metal NGLW has been reported.

The process characteristics are discussed. The resulting welds were subjected to

industry-standard radiographical approval according to ASME IX. Hardness

mapping and microstructural analysis were carried out. Tensile and impact toughness

tests were executed. Residual stresses were mapped using the contour method.

Welding equipment was developed. The unusually narrow welding groove required

special shielding gas and wire feed nozzles. Real-time weld monitoring systems

using two different approaches were developed. An appropriate restraint system for

the high distortion forces caused by the thick section welding was designed.

The microstructural and hardness properties were found to be sound and impact

toughness requirements were fulfilled in the as welded condition. This was caused

by an effective multi-pass tempering. Lack of fusion (LoF) defects were found to

limit the repeatability of the welds. Two main mechanisms were recognised:

oxidation induced LoF and undercut related LoF. The filler material used was found

to be prone to oxidation. Solutions to overcome the issues found are suggested.

34

Declaration

I hereby declare that no portion of the work referred to in the thesis has been

submitted in support of an application for another degree or qualification of this or

any other university or other institute of learning.

Tapio Väistö

1.2.2019, Manchester, United Kingdom

35

Copyright Statement

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Intellectual Property and Reproductions cannot and must not be made available for use

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36

Acknowledgements

During the course of this work I have been helped, supported and encouraged by many

people and I want to thank you all, no matter how big or small your contributions seemingly

are. It all adds up!

Firstly I would like to express my gratitude to my supervisors, Professor Lin Li and Dr. John

Francis. Thank you for all your guidance and advice over the hard working years. I also

would like to thank Dr. Neil Irvine and Ms Jacqui Grant of the NNUMAN programme for

their invaluable support during this challenging time.

I’m especially thankful to Dr. Jean Dhers from Areva for his advice and for providing the

filler material used in the welding. I’m also grateful to the EPSRC equipment pool and Mr

Adrian Walker in particular for his effort on donating the laser illumination system used for

monitoring the welds. I also want to thank the School of MACE for the scholarship I

received for three years and NNUMAN for funding my submission pending time.

A big thankyou goes to everybody involved in the substantial amount of experimental work

conducted, especially Dr. Wei Guo for guidance and Mr Damian Crosby for the robot

operation and assistance in data gathering and processing. For collaboration during the

analytical phase of the work I want to thank Dr. Anastasia Vasileiou for all her help and

expertise in contour method and hardness mapping, Dr. Dinesh Rathod for metallurgical

discussions and to Dr. Jiecai Feng for transferring NGLW knowledge and also to all the

colleagues at MTRL and LPRC.

Thanks also go to Heng and Isidro for being my best friends during my studies in

Manchester. Your support was valuable in so many different ways and is much appreciated.

I would like to use this opportunity to thank my mother Pirkko and father Pentti back home

for support and encouragement over the course of my academic studies since I was a little

boy. Thanks also go to my sister Piritta and brother Jouni for their support and reassurance

during all the twists and turns in my life, PhD and everything.

Finally, the biggest thanks of all go to my dear partner Heli. Thank you for the nudges and

shoves on the uphills of my struggle and for cheering me up on the downhills. Without your

unwavering faith this would have not been possible!

37

1 Introduction

1.1 Background

It can be argued that energy is the biggest problem faced by modern civilisation.

Global energy consumption is at unprecedented levels and rising [1]. As Britain’s 15

ageing nuclear power reactors currently (2017) supply 21 % [2] of the electricity and

as their shutdown grows nearer, there is a tremendous demand for new power plants

in the near future, Table 1.1. To add to the pressure there has been a decline in the

availability of current power plants due to ageing-related problems [2]. The renewal

of the reactor fleet is inevitable.

Table 1.1 Power reactors operating in the United Kingdom in 2017 [2]

The House of Lords Select Committee on Science and Technology [3] has identified

insufficient research and development capability as a potential threat for the UK’s

ability to produce nuclear power. New Nuclear Manufacturing Programme

(NNUMAN) [4] addressed this issue by developing R&D capabilities to support the

nuclear power needs into the future [5].

38

One of the four key research areas in NNUMAN was ‘innovative joining methods’,

including welding. Laser welding was identified as one ‘candidate advanced welding

process’ in the programme [6]. As a fully independent branch of NNUMAN, this

exploratory study investigated dissimilar metal (DMW) narrow gap laser welding

(NGLW) in pressurised water reactor (PWR) primary cooling circuits (PCC).

1.2 Primary Cooling Circuit Pressure Vessel Nozzles

The target applications considered in this study are the different pressure vessel safe-

end DMW’s in a PWR reactor PCC. The largest application, the RPV safe end weld

is illustrated as Fig. 1.1. These welds are among the most critical in the construction

and often a source of concern, as primary water stress corrosion cracking (PWSCC)

is often observed in them after prolonged use [7-10]. NGLW has shown potential to

produce lower residual stresses than conventional welding, hence reducing the

susceptibility to PWSCC [11].

Fig. 1.1 Geometry of a full-size mock-up of an RPV safe-end nozzle weld [12]

39

Currently, these welds are welded with conventional or narrow gap arc welding [13-

15]. The procedure includes laying a buttering layer of filler material, see section A-

A in Fig. 1.1. This buttering layer, with the resulting fusion line and the ferritic steel

HAZ, is then subjected to a post weld heat treatment (PWHT) to restore toughness

and relieve residual stresses before the stainless steel prone to sensitisation is present

[16]. The buttering layer is then welded onto the stainless steel pipe.

There are certain factors in the current practice to take into account. The actual weld

does not get PWHT and the residual stresses do not get relieved. This leaves a high

level of residual stress in the weld which makes it prone to PWSCC. Also, the

buttering has to be relatively thick to avoid excessive heating of the heat treated

ferritic steel, which adds to cost.

1.3 Narrow Gap Laser Welding

NGLW is a novel welding process which has seen rising interest after the invention

of high beam quality laser sources such as fibre lasers. The principle of NGLW is

illustrated in Fig. 1.2. It has been attributed to have benefits over conventional arc

welding such as higher process speeds, lower filler material consumption and lower

residual stresses. Studies of NGLW have thus far concentrated entirely on similar

metal welding [11, 17-20]. No research on dissimilar metal NGLW has been

published.

40

Fig. 1.2 Main process steps for Laser Multi-Pass Narrow Gap welding [21]

To reduce the cost and improve the manufacturability of the DMW joint, the NGLW

is to be applied without using a buttering layer.

1.4 Research Motivation and Research Questions

The welding processes currently used in PCC welds are based on arc welding. These

processes are well known, researched and there’s a wealth of experience,

commercial applications and products available. However, the processes are slow

and therefore costly. NGLW is a novel process with lots of promise, but research on

it is scarce. This study helps to understand the proposed benefits of NGLW on the

manufacturability of the joints.

One key design feature of the next generation of nuclear power plants is the

prolonged lifetime compared to previous designs. Currently, the failing PCC welds

are often among the decisive factors for decommissioning decisions for the power

stations [7-10]. Especially this is the case with the dissimilar metal welds required at

41

the pressure vessel unions [8, 22, 23]. This study probes the potential to improve the

lifetime of these welds by using NGLW.

Considering the above the following research questions arise:

Is NGLW capable of producing weld quality in accordance with the industry-

standard weld quality and mechanical property requirements?

Does the lower heat input of NGLW lead to reductions in residual stresses to

allow better resistance to PWSCC?

Can the buttering layer be omitted when using NGLW? Is the ferritic steel

HAZ toughness high enough in as welded condition?

1.5 Aim and Objectives

The motivation for this research is to investigate the proposed benefits of laser

narrow gap welding for new nuclear manufacturing. A central dissimilar weld in a

PWR primary cooling circuit was chosen for study. The welding setup, procedure

and the parameters were developed. Plateform samples were welded and the

resulting welds were subjected to industry-standard radiographic analysis

supplemented by various mechanical analysis. By doing the investigation on residual

stresses the susceptibility to PWSCC is evaluated. The ultimate target is the

reduction of the cost of manufacturing, shortening of the processing time and

improvement of the longevity of the future PWR reactors.

The aim of this study was:

To investigate the feasibility and the proposed benefits in material savings,

residual stresses and lower heat input of NGLW for SA508 Gr3 Cl2 to 316L

dissimilar metal welds in PWR reactor primary cooling circuits using nickel

Alloy 52 filler material using industry-standard acceptance criteria.

42

The objectives were:

To construct a welding setup and to develop process parameters capable of

welding thick section NGLW up to a relevant thickness to the applications

intended.

To understand the development of distortions and welding flaws during the

welding process using mechanical measurements, 3D scanning, laser

illumination imaging (LII) and X-ray analysis.

To assess the quality of the welds subjecting them to nuclear industry-

standard ASME IX radiographical acceptance and to understand the flaw

development mechanisms

To understand the formation of different microstructures and the multi-pass

tempering effect involved in the NGLW process.

To analyse the mechanical properties and understand the mechanisms leading

to them in the welds and heat affected zones by several different analysis

techniques including hardness, tensile strength and impact energy supported

by digital image correlation.

To measure the residual stresses and investigate their distribution using the

contour method and to understand their effect on susceptibility to PWSCC

To justify the omitting of the buttering and partial PWHT conventionally

involved in these types of welds by comparing the ferritic steel HAZ impact

toughness to U.S. NRC impact energy requirements.

43

1.6 Structure of This Thesis

Chapter 2 is a literature review. It describes the PWR primary cooling circuit and

the primary cooling water environment. Narrow gap welding is covered, as well as

weldability issues of high chromium filler materials. Current research of narrow gap

laser welding is reviewed. It discusses the current materials, methods and practices in

welding of the dissimilar metal weld investigated in this study. It considers the

demands posed to the materials and design by the application such as corrosion and

thermal issues. Industry-standard criteria for acceptance are described.

Chapter 3 describes the materials and methods. The welding programme structure is

described. The welding setup and parameters are presented. The chapter discusses

the extensive equipment development work conducted throughout the study. The

real-time monitoring of the welding is described. The analysis methods for the

weldments are presented in detail.

Chapter 4 presents the welding and development work conducted during the first

stage of the welding programme. The welding groove design process is described.

The observations during NGLW of 316L similar metal welds are discussed.

Chapter 5 describes the second stage of the welding programme, the first reported

dissimilar metal welds made with NGLW. The materials at this stage were S275

mild steel, AISI 316L austenitic stainless steel and nickel alloy filler metal Inconel

FM52. The welding process is described, welds are analysed and the results are

discussed.

Chapter 6 presents the third stage of the welding programme and the main objective

of this study, the dissimilar NGLW of SA508 Gr3 Cl2 low alloyed steel and 316L

stainless steel. The welding setup and the process are described. Detailed analysis of

the weldments is conducted and the results are discussed.

Chapter 7 summarises the welding programme and discusses the common findings

in the three previous chapters.

Chapter 8 concludes the key findings of this study and presents them with short

synopses.

Chapter 9 presents suggestions for future work.

44

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45

2 Literature Review

2.1 Pressurised Water Reactor (PWR)

Several different nuclear reactor systems have been developed to commercial scale

over the years [24-26]. Of these, the pressurised water reactor (PWR) type has

become the most common, with 289 reactors in operation as of 31.12.2016 [27].

The characteristics which have led to the success of the PWR design identified by

The Institute of Electrical Engineering [24] are the low construction cost resulting

from the design being amenable to fabrication in factory-built sub-assemblies and

the wealth of operating experience gained globally.

A PWR reactor, Fig. 2.1, is a light water moderated thermal nuclear fission reactor.

In a thermal nuclear reaction, the neutrons have to be slowed down (or moderated) to

maintain the chain reaction. In a PWR design, this is achieved using the cooling

water as moderant. The water also conveys the heat out from the reactor core. This

primary cooling water is kept under pressure to prevent steam formation. The low

pressure secondary circuit provides the steam driving the turbines. The steam is

generated in separate steam generators, which prevent the coolants from mixing. A

pressuriser is required to control the primary coolant pressure during thermal

fluctuations such as power changes [26].

46

Fig. 2.1 Pressurised water reactor. Primary cooling circuit in red, secondary

cooling circuit in blue. U.S. NRC [28]

2.1.1 Primary Cooling Circuit Environment (PCC)

To be able to understand the requirements for materials and design in the PCC of a

PWR reactor, the PCC environment must be understood. The following chapters

give an overview of the main considerations for design and material selection for the

components.

2.1.1.1 Heat and Pressure

Carnot’s theorem states that in order to achieve maximum efficiency any heat engine

has to operate at the largest temperature differential possible. In order to be able to

use water as an efficient means of transferring thermal energy it needs to be

pressurised to increase its boiling point. In a PWR design, the average core outlet

temperature is 320 - 325 °C [26]. This places significant thermal loading on the

structure of the pressure circuit in the thermal power plant.

47

The primary coolant pressure in a PWR is typically 150 – 155 bar and the design of a

PWR requires a large reactor pressure vessel (RPV) [26, 29]. This creates a design

challenge as exceptionally thick pressure vessel walls are required. In conventionally

fired thermal power plants higher pressures and temperatures can be used because of

the different design requiring only relatively small diameter tubing under high

pressure [26].

The primary cooling circuit piping of a PWR plant is also large-scale in comparison

to conventional power plants: 1000 mm diameters and 80 mm wall thicknesses are

typical [14, 30]. For the smallest piping in the PCC, the pressuriser surge line,

372 mm diameter and 33.45 mm wall thickness have been reported [31].

2.1.1.2 Corrosion

Hot water itself is a corrosive substance affecting the whole reactor primary cooling

circuit. In addition, the primary water is mixed with a small amount of boric acid, as

the neutron absorbing effect of boron is used to control the neutron flux and chain

reaction in the reactor core. Additionally, a minute amount of lithium is added to

stabilize the pH through the fuel cycle [32], see Fig. 2.2. The presence of boric acid

adds to the demands on the pressure vessel material as the water solution becomes

more corrosive [26]. Corrosion within PWR primary cooling circuits is a long-

established issue and has been reported by the United States Nuclear Regulatory

Commission [23]. A report by U.S. NRC specifically discusses such issues with

primary cooling circuit dissimilar welds found in 2000 and 2006 [8].

48

Fig. 2.2 Boron and Lithium chemistry control regimes in PWR [32].

2.1.1.3 Other Considerations

In addition to the static requirements, there are numerous other factors to consider.

The challenges include cyclic thermal stresses, water chemistry fluctuations,

dynamic loading, vibrations and neutron radiation embrittlement to mention a few.

These challenges have to be faced in daily operation and they must be taken into

account also in maintenance and emergency situations [26].

The unique demands set by the challenging environment of a PWR primary cooling

circuit have a considerable effect on the choice of materials. The requirements for

the materials result in using different metals for different components in the reactor.

To meet the demands of the primary cooling circuit environment in an economically

viable way, dissimilar metals are applied.

The use of dissimilar metals leads to the need for dissimilar metal joints using

welding. Dissimilar welding is common in power plant pressure vessels,

petrochemical assemblies and shipbuilding. Welds between ferritic and austenitic

49

steels are required for several reasons, mostly because the mechanical and corrosion

resistance requirements vary along the structure [33].

2.2 Narrow Gap Welding

In narrow gap (NG) welding a conventional wide V-groove is substituted with a

narrower, near-parallel welding groove. Narrow gap welding typically uses groove

angles between 2 and 20°, instead of the common V-angles in the region of 45-60°.

The aim is to reduce the consumable, time and welding passes required to weld thick

sections. NG welding is commonly applied for SAW, GMAW and GTAW processes

[34].

Fig. 2.3 Comparison of conventional V-groove to a typical NG J-groove weld

preparation

The narrow gap welding process uses a very narrow welding groove compared to

conventional processes [34-36]. This has direct implications to the weld geometry,

filler material volume and subsequent welding distortions. An extreme example is

given as Fig. 2.4, where 1.5 mm parallel groove ultra narrow gap laser welding is

compared to a conventional 60° V-groove GTA welding at 20 mm section thickness.

50

Fig. 2.4 Cross sections of ultra-NGLW and conventional GTA welds of 20 mm

thick stainless steel [35]

Comparison of commonly used weld preparation geometries for arc welding to

narrow gap arc welding is displayed as Fig. 2.5. The volumetric benefits of NG

welding compared to commonly used geometries are significant. In the case of a

single sided access, the narrow groove process can reduce the weld volume by 78%

[37]. An NGLW process is capable of reducing the gap to 4-6 mm [17], yielding a

reduction up to 93%.

Fig. 2.5 Comparison of traditional groove cross-sectional areas to narrow gap

welding in arc welding. Modified from [37]

51

NG welding generally requires specialised nozzles to access the bottom of the

narrow groove. However, in GTA welding a conventional welding torch can be

applicable up to 45 mm thickness by adjusting the electrode stick-out [36]. The wire

nozzle has to be narrow enough to fit the welding groove. In most cases, the gas

shielding nozzles have to be inserted deep in the groove to achieve efficient

shielding performance. The narrower the groove the more difficult it is to achieve

acceptable rigidity of the equipment [34-36].

Following general advantages have been listed for narrow gap welding: [34]

Lower costs due to less filler material required

Lower costs due to shorter welding time

Lower angular distortion as the weld is near-parallel

Disadvantages include: [34]

Weld is more prone to lack of fusion to the sidewalls and other defects

Reparability of defects is poor due to difficult access

More costly J-preparations required unless a root support is used

Different backing supports can be used instead of autogenous root pass [36].

Especially in the case of dissimilar metal welding, an autogenous root pass is

inapplicable as the whole thickness of the weld is required to a have similar

composition. To accomplish this, a backing support is used, see Fig. 2.6. Generally,

the backing support is removed after welding, but in certain applications, this is not

necessary.

52

Fig. 2.6 Three backing support strategies, one permanent and two removable

designs [36]

2.2.1 Filling strategies

Four main filling strategies for NG-GTAW can be listed.

Single bead per layer

Two (or more) beads per layer

Single oscillated electrode layer

Two (or more) oscillated electrode beads per layer[36]

The single bead per layer offers the best performance but can be difficult to

implement. This approach allows the use of the narrowest possible groove. The main

challenges are in machining precision and fit-up [36].

With two or three passes per layer welding energy may be reduced while

maintaining wetting and penetration, as both groove edges don’t have to be melted

simultaneously. It has benefits in residual stress which also can reduce hot cracking.

The groove width varies from 10 mm to 13 mm. Also, in dissimilar metal welding,

the properties of the base materials can be taken into account by selecting

appropriate parameters individually for each side [36].

Using a single oscillating pass with a sophisticated power source some of the

benefits of the dual pass approach can be reached. The sidewall penetration can be

optimised. Groove width, however, range up to 18 mm reducing the productivity

53

[36]. Only in cases where existing geometries cannot be changed several oscillating

passes per layer are used.

2.3 Narrow Gap Laser Welding

Narrow gap laser welding (NGLW) can be regarded as a relatively novel welding

process. For example, Steen 2003 does not mention it in the third edition of his

comprehensive handbook Laser Material Processing [38]. Also, Ion omits narrow

gap laser welding in his book Laser Processing of Engineering Materials in 2005

[39]. It also does not even appear in Steen’s 2010 revised and updated fourth edition

of Laser Material Processing [40]. However, in Handbook of Laser Welding

Technologies in 2013 Katayama [41] has dedicated a whole chapter to

“Developments in multi-pass laser welding technology with filler wire”. The

principle of dual sided NGLW is presented as Fig. 2.7

Fig. 2.7 Principle of NGLW according to Katayama 2013 [41]

54

NGLW is developed from autogenous keyhole laser welding. One major incentive

has been the possibility of welding thick section welds using much lower power than

conventionally required [42]. Other benefits have also been shown, such as lower

sensitivity to the stringent alignment tolerances [17] and smaller evaporation of

alloying elements [42] compared to keyhole laser welding.

Compared to conventional NG welding groove design NGLW uses narrower groove

design. This leads to further savings in filler material and smaller distortions [35]

with lower residual stress [41]. Katayama 2013 [41] states: ”The narrower welding

groove, lower heat input and fewer weld passes for laser welding lead to lower

residual stress.” The reduction of the weld cross-section is approximately 60% [17,

34, 37, 43]. Different NG geometries with typical dimensions are presented in Fig.

2.8.

Fig. 2.8 Comparison of narrow gap welding geometries and typical groove

dimensions, root support not shown

The ultra-narrow gap laser welding utilises a groove which is just wide enough to

accommodate the filler wire using no nozzle. The maximum applicable thickness for

ultra-narrow gap welding is limited by the rigidity of the filler wire and the laser

beam geometry. Reported thicknesses are below 20 mm [35].

55

According to Dittrich et al. 2013 [21], the so-called narrow gap laser welding has

been made feasible by the invention of new laser sources with extreme high beam

quality (beam parameter product BPP: 0.4 mm*mrad). “The excellent beam

propagation allows the laser beam to enter a narrow gap without interaction or

reflection with the accompanying material. The depth limitation into the narrow gap

is not set any more by the laser beam geometry such as in the recent past with

Nd:YAG-lasers”. Other experiments have also been conducted successfully with Nd-

YAG lasers, such as Jokinen and Kujanpää 2003 [44]. More modern fibre laser has

been used for example by Yu et al. 2013 [18], Elmesalamy et al. 2014 [11] and Feng

et al. 2016 [45].

Feng et al. [45, 46] and Guo et al. [47, 48] used a defocused laser beam in

conduction mode for the filling passes of their welds. The defocused beam was

designed to touch the side walls of the welding groove to provide heat to the base

materials, not solely relying on the heat conduction from the bottom of the groove.

The defocusing principle is described in Fig. 2.9.

56

Fig. 2.9 Defocused laser beam in an NGLW groove.

Jokinen 2004 studied laser welding stainless steel with filler wire in very narrow

grooves up to 20 mm thickness. This groove type was chosen to reduce the welding

time and distortions. In his experiments, the laser was a 3 kW Nd:YAG. The

thickness achieved was increased to 30 mm by introducing a laser – MAG hybrid

process. The groove type used in throughout the test is presented in Fig. 2.10. [17].

Fig. 2.10 The basic groove geometry used by Jokinen and a photograph of an

acceptable root pass. Different groove angles and material thicknesses

were experimented. [17].

Laser beam

Defocused beam

Base material

Focal spot

Unfocused beam

Welding groove

sidewall

57

In laser welding typical penetration depths are of the order of 1 – 2 mm / kW laser

power. With the multi-pass laser welding technique, which is based on the narrow

gap welding principle, thicker materials can be handled with relatively low laser

power.

Different methods for laser narrow gap welding have been investigated. Phaoniam et

al. 2013 have developed a hot-wire laser hybrid process for narrow gap welding. The

material was ASTM A304 austenitic SS. The filler wire was Inconel 600 nickel

alloy. [49] This resembles a technique called laser hybrid welding in which gas metal

arc welding is combined with laser welding. Hybrid welding has been studied by

several groups e.g. [50] who studied a T-joint with good results.

Yu et. al. 2013 demonstrated that after optimisation the groove size and the side

shielding gas, a sound weld without defects was achieved in welding 17 mm thick

plate by a combination of autogenous laser welding and laser welding with filler

wire technology. During the laser welding of thick plate with filler wire using a

narrow gap joint configuration and multi-pass technique, lack of fusion, inclusions

and porosity were the main weld defects [18].

In the study, they found that by using a carefully controlled shielding gas flow above

the weld pool, the concavity of the root side is reduced or can be avoided altogether

while welding the autogenous zero gap root pass. Recently Feng et al. 2017 [46]

evaluated NGLW as a candidate process for welding the ferritic nuclear reactor

pressure vessels. They found that NGLW can produce welds of a quality suitable for

critical nuclear components. The similar metal welds of SA508 Gr3 Cl1 steel were

analysed with several methods.

58

2.3.1 Challenges in NGLW

NGLW is known to add new challenges compared with traditional laser welding.

The process introduces several new parameters since it is necessary to introduce

filler wire in order to fill the gap [51]. The welding groove design is more complex.

Lack of fusion occurs easily in the welds [41, 46]. Yu et a.l. 2013 list lack of fusion,

inclusions and porosity as the main welding defects likely to be encountered in low

carbon steel welding. The narrow groove was found to be beneficial in avoiding lack

of fusion as more energy is conducted to help melt the sidewalls. The problem was

also countered by widening the weld pool, which was also found to reduce the

effects of wire feed inaccuracies [18]. Cracking in stainless steel has been studied by

Zhang et a.l. 2011 [20]. Feng et a.l. [45] reported cracking and porosity in ferritic

steel welding.

2.3.2 Welding efficiency

Conventionally the deposition rate has been used for comparing different welding

processes. In the case of NGLW, it is misleading. Using the deposition rate of

different welding processes is based on the assumption that the welding preparation

design does not change and it, therefore, is a powerful quick ratio to assess the

efficiency of welding. However, when the volume of the welding groove is reduced

dramatically, a seemingly low deposition rate may give a high rate of joining

capacity. Therefore in this study deposition rate is not used.

2.3.3 Welding Parameters for Narrow Gap Laser Welding

There are only few published articles about NGLW which directly discuss the

parameters. Jokinen 2004 experimented with several different parameters for

59

welding 20 mm stainless steel. According to his conclusions, the most reliable

parameters are presented in Table 2.1 and Table 2.2.

Table 2.1 Parameters to achieve 20 mm NGLW of stainless steel [17]

Parameter

focal length f = 200 mm

spot diameter d = 0.6 mm

laser power 3 kW

base materials AISI 316L and AISI 319LN

filler material AWS 5.9 ER 316 LSi and AWS 5.9 ER 308 Lsi

filler wire diameter 0.8 mm

wire feed angle 45 deg

groove partial V

groove angle 10 deg

air gap 0.85 mm

Table 2.2 Per pass parameters for 20 mm NGLW of AISI 304LN [17]

Pass

no:

Wire feed speed

[m/min]

Interaction point

[mm]

Focal point

[mm]

1 4 -2 -15

2 4 -2 -9

3 5.5 -2 -6

4 5.5 -2 -3

5 5.5 -2 -1

The filler wire feeding speed was selected carefully. Excess speed was found to

cause insufficient melting of the sidewalls and lack of fusion. For achieving a stable

process the interaction point of the laser beam and filler wire was found to be

important. Pointing the wire below the focal point gave the best results. As long as

the interaction point was not above the focal point, which was found to be

detrimental for the stability of the process, the accuracy was not found to be critical,

as interaction points from -2 mm to -4 mm were found to give similar good results

[17].

60

Jokinen 2004 [17] reported that the limiting factor for the weld thickness was found

to be the increasing air gap. The widening gap led to lack of fusion when the width

was 2.8 mm or more. In his experimental setup, the groove walls were inclined all

the way to the surface starting from the top of the parallel root surface. Even at the

best case, this led to an unsuitably wide groove at 21 to 24 mm thickness. Regardless

of these problems a sound weld at a thickness of 35 mm was achieved by 8 passes

with parameters which are presented in Table 2.3 and Table 2.4.

Table 2.3 Parameters for 35 mm NGLW of AISI 304LN [17]

Parameter

focal length f = 200 mm

spot diameter d = 0.6 mm

laser power 3 kW

base materials AISI 304LN

filler material ESAB OK 16.12 (AWS 5.9 ER

308LSi)

filler wire diameter 0.8 mm

feed angle 45 deg

groove partial V with 4 mm parallel root face

groove angle 8 deg

air gap 0.85 mm

Table 2.4 Per pass parameters for 35 mm NGLW of AISI 304LN [17]

Pass

no:

Wire feed speed

[m/min]

Interaction point

[mm]

Focal point

[mm]

1 3.5 -2 -30

2 4.5 -2 -25

3 5.5 -2 -20

4 6.0 -2 -16

5 6.0 -2 -12

6 4.5 -2 -8

7 4.5 -2 -3

8 4.5 -2 0

61

It is important to notice that the root pass was not made autogenously as in many

other cases. This does not apply to dissimilar metal NGLW as it does not allow

controlling of the composition and properties of the fusion zone at the first pass.

Jokinen 2004 [17] also observed porosity and fine oxidised dust. Porosity was

concluded to be related to rapid cooling and solidification of the melt pool not

allowing the bubbles to escape. Oxidation was found to be caused by an inadequate

shielding gas environment in the narrow groove. It was also found to have some non-

critical effect on porosity.

Narrow gap welding for thick section SS plates was also studied by Zhang et al.

2011. The study shows that a flawless weld of a 50 mm thick Type 316L SS can be

achieved using narrow gap laser welding. The groove geometries are described in

Fig. 2.11. [20].

Fig. 2.11 Groove geometry for narrow gap welding of 50 mm 316L [20]

The welding started with a root pass of an autogenous laser weld. It was succeeded

with four passes of build-up layers. The procedure of five passes was repeated on the

62

other side with similar parameters. The wire and shielding gas were fed from the

front of the laser beam, as illustrated in Fig. 2.12. The best results and flawless weld

were obtained at a laser power of 6 kW, welding speed 0.4 m/min, wire feed 5.5

m/min and focus spot size 0.69 mm [20].

Fig. 2.12 Experimental setup used by Zhang et al. 2011 [20]

Yu et al. 2013 tested NGLW of 17 mm thick Q235 LAS with filler wire H08Mn2Si.

The laser used was fibre laser YLR-4000 at 4 kW power. The wire feed angle was

70° and they used a special narrow nozzle to guide shielding gas to the narrow

welding groove to reduce the concavity of the root side. The experiments were made

with groove shapes described in Fig. 2.13 and parameters described in Table 2.5.

63

Fig. 2.13 Schematic diagram of NGLW of 17 mm LAS [18]

(V1:α=6.74º, b=4, c=6;V2:α=3.25º, b=2.8, c=5; V3:α=2.3º, b=2.4, c=5)

Table 2.5 Per pass parameters for Q235 LAS used by Yu et al. 2013 [18]

Pass

no:

Welding speed

[m/min]

Wire feed speed

[m/min]

Side blown

shielding gas

nozzle diam.

[mm]

Side blown

shielding gas

flow

[m3/h]

1 0.7 0 1.4 0.2

2 0.5 4.0 1.4 0.2

3 0.5 4.5 1.4 0.2

4 0.5 4.5 6.0 1.0

Certain uniformities can be found in the parameters for narrow gap laser welding

found in the literature. The laser beam is usually tilted a few degrees forward in

order to reduce back reflections to the laser source. As is the layout such that the

shielding gas and filler wire are introduced in front of the welding pool. [19, 20]

2.3.4 Hot Wire Laser Welding (HWLW)

A recent application of NGLW is the hot wire laser welding (HWLW). The gap

geometries used in hot wire laser welding are similar with very narrow grooves

64

offering the geometrical benefits [52, 53]. In HWLW the filler wire is attached to an

electric arc welding power source, but the voltage is kept low enough to prevent arc

formation. The concept is to heat the wire electrically prior to introducing it to the

melt pool. Westinghouse 2015 [52] lists the following benefits for HWLW:

Narrow grooves with up to 70 % volume reductions

High deposition rates, 5 times that of GTAW

Up to 90 % reduction of welding time in thick section welding

Lower distortion and residual stress

Reduced machining and non-destructive testing

Automated process

Improved welding performance on highly irradiated materials

Näsström et al. 2015 [53] reported that HWLW is sensitive to wire and laser

alignment, and a novel welding setup was designed to improve repeatability. They

report reaching 5 mm thick weld beads and using a wide laser spot improving the

quality of wetting, as it was found melting the sidewalls. They were using AR400

wear-resistant carbon steel.

Todo et al. 2015 [54] investigated HWLW of 9 % chromium steels P91 and KA-

SCMV28. They used a rectangular spot diode laser for welding up to 54 mm wall

thickness pipe. They reported low dilution and narrow HAZ. Fusion was improved if

a soft focus was used and the bead thickness was kept low.

In 2017 Näsström et al. [55] report that solidification cracking was observed on

HWLW of carbon steels. This phenomenon was more common in the filling passes

where the restraint was higher than during capping passes.

65

In 2016 Kaplan et al. [56] performed up to 14 mm thick welds in AR400 steel using

3 mm wide grooves. The process was found very flexible in terms of laser power and

welding speed used. They concluded that the electrical pre-heating of the filler wire

increased the performance and possibly improved the wetting.

It can be seen the characteristics of HWLW are similar to those of NGLW. The

major difference is in the deposition rate, which is higher in HWLW. Also, weld

pool behaviour such as wetting has been reported to improve when using a pre-

heated wire.

2.4 Present-day Solutions in Ferritic-Austenitic Dissimilar PWR

PCC Welding

The established practice in PWR PCC construction is to build the pressure vessels of

ferritic high strength low alloyed steel (HSLA or LAS) for its mechanical properties

and then clad the inner surface of the vessel with an austenitic stainless steel (SS)

layer to address the corrosion challenges [26]. The PCC piping is made of solid

stainless steel for simplicity.

The use of dissimilar metals in the construction of nuclear reactor components leads

to the need for dissimilar metal welding. Fang et al. [13] list typical locations for

dissimilar metal welds in PWR power plants. They can be found in several places in

the power plant, for example in the primary coolant circuit they are required in the

RPV nozzles, steam generator nozzles, the pressurizer and at the main circulating

pumps. An important detail in the pressure vessel is the pressure vessel head (PVH),

through which the control rod drive mechanism (CRDM) operates. The CRDM

assembly Fig. 2.14 is made of LAS and Alloy 690 [57].

66

Fig. 2.14 Photo of CRDM (a) and schematic diagram of CRDM assembly (b) [57].

The particular welds considered in this study are the RPV primary cooling circuit

pressure vessel nozzle dissimilar metal welds. One of the most critical welds in a

PWR NPP is the dissimilar metal weld joining the reactor vessel to the primary

cooling circuit nozzle. Its total failure would lead to a loss of coolant accident

(LOCA), which is the most serious condition a PWR is planned to withstand with no

leak of radiation, called design basis accident (DBA) [26]. The RPV cooling circuit

nozzle and its location in the reactor are pictured in Fig. 2.15.

67

Fig. 2.15 Typical locations of dissimilar welds in a PWR [13].

2.4.1 Dissimilar Welding Processes and Geometries Used in PWR

Primary Cooling Circuit Manufacturing

The applications considered in this study are the ferritic-austenitic dissimilar welds

in the primary cooling circuit of a PWR power plant. An example of a current

arrangement of welding the RPV nozzle to the primary coolant circuit pipe is shown

68

in Fig. 2.15. The actual dissimilar GTAW is preceded with thick arc-buttering of the

ferritic RPV nozzle using a Ni alloy. Then an austenitic SS safe-end is added with

GTAW to provide a nozzle. The dissimilar metal weld is performed at manufacturer

facilities, which allows a similar metal weld to be used for the assembly at the power

plant construction site [13]. Other welding processes such as manual metal arc

welding (MMAW) [14] and submerged arc welding (SAW) [15] are also currently

used.

Fig. 2.16 illustrates a dimensioned drawing of an RPV PCC nozzle of a current

design NPP by Westinghouse. Concluding from the filler materials used, the welding

methods are arc welding, most likely GTAW for the buttering and MMAW for the

actual weld [14]. This illustrates both the geometry and the current very conservative

practices applied by the industry.

69

Fig. 2.16 Geometry and dimensions of Westinghouse AP1000 PWR RPV safe-end

dissimilar metal weld. A – SA508 ferritic steel, B – Alloy 82 buttering, C –

Alloy 182 weld, D- 316L austenitic SS. Note the cladding of RPV with SS

[14].

The Shanghai Company of Nuclear Power Equipment produced a full-scale RPV

nozzle joint. The nozzle outer diameter was 1001.6 mm and wall thickness was 83.5

mm. The SA 508 was pre-heated to 125 °C. The buttering was applied with a total of

478 weld passes at a welding speed of 1.85 mm/s. The buttering was deposited with

70

GTAW creating a 20 mm thick layer. Then the buttered layer was heat treated

annealing it at 610 °C for 15 hours with furnace cooling to 300 °C to relieve residual

stresses. After that, a 100 % NDT ultrasound test was performed. The actual welding

took 439 GTAW passes at 1.75 mm / s to form a 19 mm wide weld [30]. Calculating

from the figures above, the buttering and welding took over 400 hours in total, not

including inspection and setup work.

The rationale behind PWHT is to reduce residual stresses and temper the

microstructure to restore toughness [16]. It has been found that the post weld heat

treatment affects the corrosion fatigue and SCC crack growth in simulated PWR

conditions. Huang et al. (2013) conclude that a long heat treatment at 621°C can

increase the crack growth rate, but a lower temperature heat treatment can improve

the corrosion resistance properties compared to in the as welded condition [58].

Primary cooling circuit dissimilar welds are needed also at steam generator nozzles.

An example of a design is presented as Fig. 2.17. One other application is the

pressuriser surge nozzle pictured in Fig. 2.18 and Fig. 2.19. In 2010 Rudland et al.

[59] report using a 35 mm wall thickness for their simulation of residual stresses at a

pressuriser surge line nozzle DMW.

71

Fig. 2.17 Steam generator dissimilar metal weld design [60]

Fig. 2.18 Different pressuriser nozzle dissimilar welds before and after a repair

[60]

72

Fig. 2.19 Pressuriser surge line nozzle design and dimensions [61]

2.4.2 Metals Used in PWR Primary Cooling Circuit Construction

Both low alloy steels and austenitic stainless steels are widely used in PWR nuclear

power plants. Typical examples are the various pressure vessels in the primary

cooling circuit made of LAS and clad with SS. The major branch lines of the primary

cooling circuit are made of SS to provide the required strength and corrosion

resistance [13]. Low alloyed ferritic ASME SA508 is widely used in RPVs The

cooling circuit piping is often made of austenitic stainless steel AISI 316L [13, 62].

73

Welding of the dissimilar metal connection between RPV and the cooling circuit

piping is currently done by arc welding with nickel-based filler materials. The trend

is to move to high chromium alloys, such as Inconel Alloy 52 [58, 63].

Conventionally Alloy 82 and Alloy 182 [64, 65] have been more popular and in

some earlier applications austenitic stainless steel [15] has also been used.

Fig. 2.20. illustrates the numerous different metals and materials used in a PWR with

several locations where dissimilar metal welding is applied.

Fig. 2.20 Outline of PWR components and materials [66].

2.4.2.1 Base Materials

The base materials chosen for this study from the outset were SA508 ferritic low

alloyed steel and ASME 316L austenitic stainless steel. ASME SA508 low alloyed

high strength steel is widely accepted and widely used in the construction of PWR

reactor vessels. There are several specifications for different heat treatments for SA

74

508 [67]. A variant called Grade 3 exhibits better mechanical properties than earlier

versions and is regarded as the material of choice for next-generation NPPs [68-70].

The chemical and mechanical properties of SA508 Grade3 Class2 are presented in

Table 2.6 and Table 2.7. SA508 Gr3 Cl2 is heat treated by austenising and

quenching followed by tempering to achieve the desired properties. In ASTM

A508/A508M – 17 [67] the austenising procedure is not specifically specified for

Gr3 Cl2. The tempering time is given as 30 min per inch of maximum section

thickness and the temperature of 620°C.

Table 2.6 Chemical composition of A508 Grade 3 Class 2 steels used in a study [67]

wt.-%

C Fe Mn Mo Ni Si Cr Cu

A508 Gr3 Cl2 0.25 Bal. 1.2-1.3 0.45-0.60 0.40-1.00 0.4 0.25 0.2

V P S Ca Ti Al Nb B

0.05 0.025 0.025 0.015 0.015 0.015 0.01 0.003

Table 2.7 Mechanical properties of SA508 Gr3 Cl2 steel[67].

Ultimate

tensile

strength

(MPa)

Yield

strength

(MPa)

Total

elongation

(%)

Reduction

of area (%)

Gr3 Cl2 620-795 450 16 35

AISI 316L stainless steel is a common austenitic stainless steel used for PWR PCC

piping because of its suitable properties, such as high resistance to stress corrosion

cracking [11]. The 316L is a derivative of the original AISI 316 stainless steel with

reduced carbon content to reduce the susceptibility to intergranular corrosion due to

the mechanism of sensitisation. The chemical compositions are presented in Table

75

2.8. Nominal mechanical properties of AISI 316L in an annealed condition are

presented in Table 2.9.

Table 2.8 Standard chemical composition of AISI 316 and AISI 316L wt.-% [71, 72]

C

max

Mn

max

P

max

S

max

Si

max

Cr Ni Mo N

AISI

316

0.08 2 0.045 0.03 1 16.00 –

18.00

10.00 –

14.00

2.00-

3.00

0.1

AISI

316L

0.03 2 0.045 0.03 1 16.00 –

18.00

10.00 –

14.00

2.00-

3.00 0.1

Table 2.9 Mechanical properties of AISI 316L annealed [71].

Tensile

Strength

[MPa]

Yield

Strength

[MPa]

Elongation

%

Reduction in

area

%

Hardness

[HB]

Hardness

[HV]*

550 200 45 55 140 142

* Converted from HB by [73]

In a real-life study [74] the chemical composition of a batch of AISI 316L was

investigated in detail. This is presented in Table 2.10. The actual carbon content was

found to be much lower than the standard maximum value. This will have

significance in lowering carbon migration and, hence, reducing susceptibility to

sensitisation.

Table 2.10 Chemical composition 316L wt.-% [74]

C Cr Ni Si Mn Mo Al Co

316L 0.016 17.68 12.6 0.663 1.53 2.38 0.018 0.121

Cu Nb Ti V W Fe P S

0.211 0 0.021 0.663 0.029 64.42 0.02 0.003

76

2.4.2.2 Filler Metals

Filler metals for welding of dissimilar ferritic-austenitic steels have changed through

the decades. The first successful industrial applications were in the 1940s. The early

welds were made with austenitic stainless steel filler metal. These were later found to

be prone to failure due to creep in prolonged use. This problem was largely solved

by introducing nickel-based filler materials to reduce the difference of thermal

expansion coefficients between LAS and SS. However nickel based materials have

also been found to fail due to stress corrosion cracking after decades of use [64, 75-

77].

INCONEL Filler Metal 82 (FM82) and INCONEL Welding Electrode 182 (WE182)

[78, 79], commonly abbreviated as Alloy 82 and Alloy 182 respectively, have been

widely used in dissimilar welds between LAS and SS in the nuclear industry. They

have a composition similar to the Alloy 600 base material. Inconel Filler Metal 52

and Inconel Welding Electrode 152 are typical high chromium welding consumables

[80, 81] for dissimilar welding and are more resistant to SCC. These have

compositions nearly matching the composition of Alloy 690.

Due to widely observed stress corrosion cracking (SCC) in welds made with alloy

182 welding rods in many PWRs currently in use [82, 83], the trend in recent

decades has been to use higher chromium content FM52 and WE152, which have

better resistance to SCC. The mechanism improving SCC resistance is segregation of

Cr to the solidification boundaries [84]. The ability to retard carbon migration to the

weld is also in favour of using alloy 52 in dissimilar welds [85]. Repairing SCC in

Alloy 82 based welds using Alloy 52 has been reported [8, 23]. They have also been

77

successfully used in repairing the control rod drive mechanism and thermocouple

penetration nozzles, pressuriser nozzles and hot leg nozzles, etc. [58].

One important factor in choosing a filler material is its susceptibility to hot cracking.

This can be evaluated using the Varestraint method. Tests have shown that Alloy 82

has greater hot cracking susceptibility than Alloy 52 [86]. Their experiment was

conducted in a similar metal weld of Alloy 690 base metal.

The properties of the nickel filler alloys currently commonly used are presented in

Table 2.11 and Table 2.12.

Table 2.11 Chemical compositions of Ni filler alloys wt.-% [58]

Ni Cr Fe C Si Mn Co

Alloy 82 71.88 20.46 1.41 0.03 0.16 2.94 N.A. Alloy 52 61.44 28.8 8.07 0.024 0.15 0.24 0.009

S Cu Al Ti Nb+ Ta P

Alloy 82 0.001 0.12 N.A. 0.37 2.47 0.004 Alloy 52 0.001 0.01 0.72 0.51 0.01 0.003

Table 2.12 Mechanical properties of Alloy 52 and Alloy 82 fillers [58]

Alloy 82

Alloy 52

Temperature 25 300 25 300 °C

UTS

734 627 714 553.6 MPa

Yield strength 524 367 540 430 MPa

Total elongation 25.3 29.23 26.5 19.58 %

Uniform elongation 19.5 22.78 17.5 11.1 %

The filler material versions of these alloys differ slightly from the wrought alloys

and are presented in the following chapters with descriptions of their characteristics

and typical applications.

78

Inconel Filler Metal 82 is used for gas tungsten arc (GTA), gas metal arc welding

(GMAW) and submerged arc (SMA) welding. It is described to have high strength

and good corrosion resistance at elevated temperatures. It is also used to join ferritic

steels to austenitic stainless steels, see Table 2.13 [78].

Table 2.13 Chemical composition of Filler Metal 82 [78]

Ni+

Co

Cr C Ti Mn Fe P S Nb+

Ta

Si Ti Cu Others

FM

82

67 18.0

–

22.0

0.10

max

0.75

max

2.5

–

3.5

3.0

max

0.030

max

0.015

max

2.0 –

3.0

0.50

max

0.75

max

0.30

max

0.50

max

Filler Metal 52 is designed for GTA and GMAW of Inconel Alloy 690. It is also

useful for dissimilar joints involving carbon, low alloy and stainless steels. The

chemical specification of Filler Metal 52 is presented in Table 2.14. [80]

Table 2.14 Chemical composition of Filler Metal 52 wt.-% [80]

Ni + Co Cr C Ti Mn Al Fe P

FM 52 rem 28.0 – 31.5 0.04 max 1.0 max 1.0 max 1.10 max 7.0 – 11.0 0.02 max

S Nb+ Ta Si Al + Ti Mo Cu Others

0.015 max 0.10 max 0.50 max 1.5 max 0.50 max 0.30 max 0.50 max

Filler Metal 52M, Table 2.15, is designed for GTA and GMW of Inconel Alloy 690

and overlaying of carbon and stainless steels. The high chromium level provides

excellent resistance to SCC in nuclear water environments. This alloy contains boron

and zirconium to minimise the tendency of ductility-dip cracking (DDC). It is

especially resistant to oxide “floaters” and inclusions [87].

79

Table 2.15 Chemical composition of Filler Metal 52M wt.-% [87]


FM 52M rem 28.0 – 31.5

0.04 max

1.0 max

1.0 max

1.10 max

7.0 – 11.0

0.02 max

S Nb Si Mo Co Zr B Cu Others

0.015 max

0.50 – 1.0

0.50 max

0.50 max

0.12 max

0.02 max

0.005 max

0.30 max

0.50 max

Inconel Filler Metal 52MSS, see Table 2.16, is a third-generation 30 % Cr Inconel

welding product designed to resist nuclear pure water intergranular SCC. The SCC

resistance has been found to be as good as FM52M [88]. Addition of Mo and Ni

increases resistance to DDC and cold cracking. Welds are to be free of inclusions,

oxides and porosity [89].

Table 2.16 Chemical composition of Filler Metal 52MSS wt.-% [89]


FM 52 MSS

54.0 – 62.0

28.0 – 31.5

0.03 max

0.50 max

1.0 max

0.50 max

bal 0.02 max

S Nb + Ta Si Al + Ti Mo Cu Others

0.015 max

1.5 – 3.5 0.50 max

1.5 max

3.0 – 5.0

0.30 max

0.50 max

Regarding the filler metals used in the nuclear industry and in the PWR construction,

one can draw the following conclusions. Stainless steel fillers became obsolete when

their tendency to fail after a decade of elevated temperature use was found: Alloy 82

and derivatives were used for most of the currently operating PWRs. This low

chromium filler has been found to be prone to stress corrosion cracking and has been

repaired and replaced by high chromium Alloy 52. The overall trend has been

moving away from Welding Electrode 182 towards Filler Metal 52M.

80

2.5 Considerations in Austenitic-Ferritic Dissimilar Metal Welding

in PWR Cooling Circuit

There are many phenomena taking place in the weld metal and the heat affected zone

adjacent to it. In the case of dissimilar welding, the phenomena are various and often

complex.

When welding carbon steels to austenitic steels in nuclear applications a filler

material with increased nickel and chromium is often used. The increased alloy

content helps form a similar composition in the weld when diluted with the carbon

steel base material. Nickel-based fillers are used particularly in high-temperature

applications because the thermal expansion coefficient is closer to carbon steel

compared to austenitic filler [90]. In nuclear dissimilar welding applications, nickel

alloys are used extensively [83].

The method currently used in nuclear dissimilar welding involves buttering of the

ferritic RPV steel with nickel alloy [12, 14]. This is done by weld depositing a layer

of nickel alloy on the ferritic steel. During the buttering process, the HAZ of the low

alloy steel develops a brittle martensitic structure, which has to be tempered with

post weld heat treatment [75]. The buttered ferritic steel can receive the heat

treatment required before the final welding. This eliminates the need to perform

PWHT to the final joint [90]. PWHT of the whole joint would cause sensitisation of

the stainless steel base material which increases the susceptibility to intergranular

corrosion [91, 92].

81

2.5.1 Corrosion

Corrosion is defined by ISO as “a physicochemical interaction leading to a

significant deterioration of the functional properties of either a material, or the

environment with which it has interacted, or both of these”. Over the chemical

aspects, an important aspect considering welding of metals, this definition includes

combinations of environmental deterioration where these are influenced with stresses

within the material [93]. The residual stresses caused by welding, for example, have

been found to be a significant source of SCC [94]. PWSCC (primary water stress

corrosion cracking) in PWR reactors has been reported extensively [7, 8, 23, 83] and

has been studied by several researchers [15, 85, 95].

Couvant (2013) describes the problem as “a large variety of the structural metals

present in primary and secondary circuits of pressurized water reactors suffer

corrosion. Uniform corrosion, flow-accelerated corrosion (FAC), pitting, stress

corrosion cracking, environmentally assisted fatigue and hydrogen embrittlement can

all affect the major components of PWRs, despite stringent selection of materials for

component manufacture” [7].

As the PWR primary coolant partially serves the purpose of a neutron moderator, its

pH is acidic. To control corrosion, the coolant chemistry has to be managed within

strict limits. The objectives of the coolant chemistry management are to minimise all

types of corrosion of materials, to assure fuel-cladding integrity, to assure pressure

boundary integrity and to optimise the radio-protection of workers by minimising the

corrosion and cation release and their activation in the reactor core [96].

Stainless steels are resistant to corrosion and therefore used in numerous applications

in PWR cooling circuits. The interior of the LAS pressure vessel in a PWR is clad

82

with SS to withstand the corrosive conditions [26]. Stainless steels have extremely

good general resistance to corrosion but are nevertheless susceptible to pitting

corrosion[97].

2.5.2 Sensitisation of Austenitic Stainless Steel

Weld decay, or sensitisation, of austenitic stainless steels, is caused by precipitation

of Cr carbides at grain boundaries, whose effect can be seen in Fig. 2.21. Within the

sensitisation temperature range, carbon atoms rapidly diffuse to grain boundaries

where they combine with Cr to form chromium carbides which causes intergranular

corrosion. [98] Steels containing less than 0.03 % C are however immune from weld

decay [91], which is the case when using AISI 316L with nominal Cmax of 0.03 %.

Fig. 2.21 Intergranular corrosion in HAZ of 304 SS containing 0.05 % C [98].

Lancaster [91] presents a graph, shown in Fig. 2.22., describing the time needed to

produce harmful amounts of chromium carbides. The graph suggests that when using

low carbon AISI 316L with measured 0.016 % C as Hasanbaoglu and Kacar 2007

[74] did, harmful amounts of sensitisation would occur only after hundreds of hours

83

of heat treatment at 600 °C. This indicates a possibility to heat treat the troublesome

ferritic-nickel boundary post welding and avoid the buttering altogether, as opposed

to current practice.

Fig. 2.22 The effect of carbon content on carbide precipitation [91]

Huang et al. studied the environmentally assisted cracking of DM welds of SA 508

using Alloy 52 and Alloy 82 in a simulated BWR environment. Their observation

was that a PWHT of 24 h at 621 °C increased the crack growth rates compared to the

as welded state. However, a PWHT of 8 h at 621 °C continued for 200 h at 400 °C

showed improvement of the corrosion resistance [58].

If the rationale for avoiding PWHT of the stainless steel safe-end is to avoid the

sensitisation of the stainless steel component, it can be argued, that according to this

literature review the sensitisation of real-life 316L stainless steels is not going to

occur during a typical PWHT cycle. Combining the chemical analysis of a real 316L

carbon content from Table 2.10 and the effect of carbon content on carbide

precipitation from Fig. 2.22 indicates, that a 0.016 % C steel would stand well over

100 h of heat treatment at any temperature.

84

2.5.3 Residual Stresses

The external loadings are important factors in designing any mechanical structures

and usually the external loading is the most easily considered factor. However, other

factors often play a determining role. The other contributing elements include

unfavourable microstructures, the likelihood of pre-existing defects and residual

stresses. There are many ways in which failure can occur including brittle fracture,

plastic collapse, fatigue, creep and stress corrosion cracking. Compared with the

attention received by the microstructure and defects, the residual stresses have

received relatively little attention. It is clearly important that the origins of residual

stresses are recognised [99].

Residual stresses are stresses that would exist in a body if all external loads were

removed. They are sometimes called internal stresses. Residual stresses that exist in

a body that has previously been subjected to non-uniform temperature changes, such

as those during welding, are often called thermal stresses [98]. Welding is one of the

most significant causes of residual stresses and typically produces large tensile

stresses. The development of residual stresses in dissimilar metal welding has been

widely studied [100-103].

There are several ways to estimate and compute the residual stresses created in

dissimilar welding and the necessary cladding thereafter. In 2006 Gilles & Nouet

[104] developed a numerical method for prediction of residual stresses in a PWR

PCC weld. Deng et.al. (2009) [105] suggested a simplified method considering

cladding, buttering, post weld heat treatment and multi-pass welding using a united

fashion. In 2011 they [101] presented a time-effective 3-D finite element model to

simulate welding residual stresses through a simplified moving heat source.

85

One of the biggest motives for investigating NGLW on thick dissimilar welding is

probably from the results published by Elmesalamy et al. in 2014 [11]. They

compared residual stresses of GTA welding and NGLW in 20 mm AISI 316L welds

and found the longitudinal tensile stresses to be 30 – 40 % lower in the case of

NGLW [11]. It should be noted that these tests were made using only 20 mm

thickness plate and the weld was not dissimilar.

2.5.3.1 Types of Residual Stresses

Different types of stress are characterised according to the characteristic length scale

over which they self-equilibrate. Residual stresses originate from misfits between

different regions. In many cases, these misfits span large distances, for example,

those caused by welding or heat treatment operations [106]. Fig. 2.23 illustrates

examples of residual stresses. In this study, the residual stresses resulting from

welding are to be evaluated.

Fig. 2.23 Examples of different types of residual macro- and micro-residual stress

[99]

86

There are at least four ways in which macro residual stresses can form: due to

misfitting parts in an assembly, and due to the generation of chemical, thermal, and

plastically induced misfits between different regions within one part [107]. Residual

stresses formed by a simple case of welding are presented in Fig. 2.24.

Fig. 2.24 Schematic representation of changes of temperature and longitudinal

thermal residual stresses during bead-on-plate welding [99]

The residual stresses caused by welding are of type I stresses because they span over

large distances. The residual stresses that vary over the grain scale are type II and in

atomic-scale type III. It is worth noting, that different scale residual stresses don’t

record in tests of other scales. For example, type II and III stresses don’t show in

measurements for type I [106].

In a single pass weld, the following observations have been reported. In the case of a

solid-state transforming material, the stress distribution has a distinctive M-shape

[108]. The residual stress distribution in a non-transforming material weld is more

like an inverted U shape with a single peak, see Fig. 2.24.

87

In a multi-pass weld, the overlaying passes heat treat the passes already made. This

will reduce the residual stresses considerably and the M or inverted U shapes are

only observed close to the top of the weld.

2.5.4 Environmentally Assisted Cracking and Stress Corrosion

Cracking

In 1991 Marston and Jones [9] state that in the last three decades environmentally-

assisted cracking (EAC) of structural materials in light water reactors (LWR) has

been one of the major causes of lost availability in both boiling water reactors

(BWR) and PWRs. They conclude that this has resulted in significant economic

losses.

An understanding of EAC, more accurately stress corrosion cracking is especially

important in the case of welded structures. One important challenge is the residual

stresses which are emphasised with differences in thermal coefficients in the

materials joined [11].

SCC is sometimes understood either in a narrow sense: EAC under purely static

mechanical loading or as part of the broader spectrum of EAC including the

transition to strain-induced corrosion cracking (SICC) and low-cycle corrosion

fatigue (LCF). When classified by the external loading the types can be related to

different operational states of an NPP. SICC and LCF are experienced during

transitions in operation, e.g. start-up or shutdown. SCC occurs during steady-state

operation due to static loading [109]. Different types of EAC are described in Table

2.17.

88

Table 2.17 Basic types of EAC [33].

Designation Type of mechanical loading

Stress corrosion cracking (SCC) Static

Strain-induced corrosion cracking

(SICC)

Dynamic — slow monotonically rising

or very low-cycle

Corrosion fatigue (CF) Cyclic — low-cycle, high-cycle

Several studies acknowledge SCC being observed in primary cooling circuit nozzles

near the dissimilar weld [7, 58, 95, 100, 110]. Ihara et al. 2011 [110] report SCC

being found near the weld zone of the core shroud and primary loop recirculation

pipes made of Type 304 austenitic stainless steel. The initiation and propagation of

SCC result from the superposition effect of sensitisation due to the temperature

history of welds, the corrosive atmosphere and the residual stress due to welding.

The problems have been addressed by changing the steel to low-carbon type 316L

austenitic stainless steel. Nevertheless, SCC has been found regardless of the

absence of sensitisation. The conclusion has been that residual stresses have a bigger

effect on SCC than the material or sensitisation. Ihara et al. 2011 [110] conclude that

the effect of machining is significant, especially in the early stages of crack growth,

and should be studied further. The SCC happening with or without sensitisation

gives an opportunity to consider PWHT of the whole weld avoiding buttering

altogether as the buttering is only made to avoid exposing the SS nozzle to the

PWHT temperature cycle.

As a result of full suppression of water radiolysis in the PWR primary coolant

circuit, it has been reported that the major water chemistry factor to determine

primary water stress corrosion cracking is not oxygen but hydrogen [111].

89

Chung [85] studied the microstructure and SCC behaviour in dissimilar welds

between A508 and Alloy52. They noticed in their study that post weld heat treatment

caused the near weld interface layer to transform to martensite, carbides and Type II

boundaries. The presence of Type II boundaries causes a reduction of resistance to

SCC and formed intergranular cracking under simulated PWR conditions [85]. The

welding process was conventional multipass GTAW and the microstructural

conditions may not be transferable to NGLW as such.

Karlsen (2010) [112] addressed the significance of localised deformations to the

development of stress corrosion cracking in stainless steels in PWR PCC

environment. The sources of deformations listed included the welding distortions.

2.5.5 DMW Creep Failure in the HAZ of Low Alloy Steel

The decisive factor for selecting nickel alloys for ferritic – austenitic steel dissimilar

metal welding is the similarity of the coefficient of thermal expansion (CTE) of the

ferritic steel and Ni alloys. Previously the difference in CTEs between the austenitic

stainless steel filler and the ferritic steel caused detrimental tensile thermal stresses in

the ferritic steel. This phenomenon was explained by Klueh and King in 1982 [113].

Creep failure in a low alloyed steel (LAS) heat affected zone after 10 to 15 years of

elevated temperature service had been reported when an austenitic stainless steel

filler was used [113]. The phenomenon was found to onset by a soft LAS HAZ

microstructure formed due to carbon migration to the Cr rich weld metal. The carbon

migration starts during welding and develops during PWHT and elevated

temperature use. The soft ferritic microstructure creep fails along the ferrite grain

boundaries under residual and thermal stresses due to the difference in CTE in HAZ

and WM [113]. To avoid the high thermal stresses the CTE mismatch is shifted to

90

the stainless steel fusion line by using nickel alloy as weld metal. The stainless steel

fusion line is much stronger and more crack-resistant [113].

Research made by W. Childs at EPRI (Electric Power Research Institute) shows a

significant increase in the service life of dissimilar welds using nickel alloys, see Fig.

2.25. The data is collected from various power plants around the world. The use of

Ni-alloys is recommended for service temperatures exceeding 425°C [75].

Fig. 2.25 Comparison of the average life of dissimilar welds, Type 309 austenitic

steel = 1. FM82 both high (H) and low (L) heat input welds are included

[75].

2.5.6 Radiation Damage

Electrons, gamma rays, neutrons and heavy charged particles can cause excitation,

ionisation and disassociation of materials. Organic materials are relatively weak but,

although stronger, metals are also affected. Metals in nuclear reactors exhibit

increases in tensile strength and hardness, but a reduction in ductility. The degree of

damage tends to be lower the higher the temperature because displaced atoms can

more easily return to their original places [25].

91

As continuous neutron irradiation changes the properties of materials in a nuclear

reactor, the effects on the ductile-brittle-transition-temperature (DBTT) are

monitored using prefabricated Charpy test pieces put into the RPV before bringing

the reactor to critical. These are then tested destructively at regular intervals. This

procedure poses a challenge in extending the plants’ operational life beyond the

originally planned service life. It is not possible to add these test pieces during the

operation of the plant and if they run out, a new non-destructive testing (NDT)

method of testing has to be implemented. This can be tedious since no NDT method

can give direct mechanical results as a Charpy test does. Therefore several tests have

to be conducted and correlation to DBTT statistically proved [114].

It has been found that the embrittlement process is very complicated and depends on

several factors such as chemical composition, irradiation temperature, irradiation

flux and fluence etc. Some semi-empirical laws have been established, based on the

macroscopic data, but unfortunately, these laws are never completely consistent with

all data and do not yield the wanted accuracy [115]. This, maybe one of the reasons

for the nuclear industry’s well known conservative attitude towards new

technologies, indicates that using tried and practice proven materials is beneficial

when introducing new methods to the industry.

2.6 Dissimilar Metal Welding (DMW)

Discontinuities in grain structure are characteristic for dissimilar metal welding. In

ferritic steel welds using nickel filler, instead of normal epitaxial grain growth,

different growth types occur and the resulting microstructure can be complex [75].

The focus of this study is at the ferritic steel to austenitic weld metal fusion line.

92

2.6.1 Phase Prediction of Austenitic Weld Metal

The solidification mode of the weld metal is a fundamental characteristic of the

resulting microstructure. A tool used for fusion zone microstructure prediction of

austenitic materials is typically the Schaeffler diagram, see Fig. 2.26.

Fig. 2.26 Schaeffler diagram showing the reduction in the extent of the

ferrite/austenite region for high cooling rate processes [116]

The filler metal FM52 used in this study has Nieq 54.125 and Creq 34.75. These place

the solidification mode well into the fully austenitic region.

For better compatibility with, for example, dissimilar metal welding, the WRC-1992

diagram, Fig. 2.27, is used.

93

Fig. 2.27 WRC-1992 diagram for predicting ferrite content and solidification mode

[98]

The equations for WRC-1992 give Nieq 53.9 and Creq 32.07 for FM52. These place

the predicted solidification mode again strongly in the fully austenitic area. This

mode will be expected even with a relatively high dilution of the ferritic or the

austenitic steels.

2.6.2 Solidification Mechanisms in Dissimilar Metal Welding

In similar metal welds, where the crystal structure does not differ between weld

metal and base material, epitaxial nucleation during solidification along the fusion

boundary gives rise to grain boundaries that are continuous from the base metal into

weld metal across the fusion boundary. These boundaries are roughly perpendicular

to the fusion boundary and have been referred to as Type I boundaries, see “Normal

Fusion Boundary” in Fig. 2.28 [75, 98]. This is the case, at the austenitic stainless

steel to austenitic filler metal fusion line in this study.

94

Fig. 2.28 “Normal” grain growth with Type I grain boundaries and the grain

growth exhibiting Type II boundaries in a dissimilar metal weld [98]

In the case of body centric cubic (BCC) ferritic LAS being welded with austenitic

face centric cubic (FCC) nickel filler metal the epitaxial growth is not possible, but

the new grains have to nucleate differently. In this case heterogeneous nucleation

starts at random misorientations at the fusion boundary [98].

In dissimilar welds, where an austenitic weld metal and ferritic base metal exist, a

second type of boundary that runs roughly parallel to the fusion boundary is often

observed; see Fig. 2.28. This has been referred to as a Type II boundary [117]. Type

II boundaries form at the prior austenite boundaries of the base material, a few

microns into the weld metal from the fusion boundary, Fig. 2.29 [75]. These

boundaries typically have no continuity across the fusion boundary to grain

boundaries in the base metal. Several investigators have reported that hydrogen-

induced disbonding typically follows Type II grain boundaries. The disbonding

95

phenomenon that occurs following fabrication and prior to service has also been

associated with these Type II boundaries [117].

Fig. 2.29 Dissimilar metal fusion line and the formation mechanism of Type II

boundaries [75]

The grain growth in weld metal continues, regardless of the initiation, as competitive

growth [98]. This makes the weld metal bulk near symmetric in general appearance.

2.6.3 The Heat Affected Zone

In welding, the part of BM which undergoes metallurgical changes but does not melt

is called the heat affected zone (HAZ). Properties of the HAZ are important in

achieving a proper weld. The yield and ultimate strength of HAZ are almost always

higher than the base material, causing no problems. The properties of main interest

are fracture toughness and hardness. In a weld of low carbon steel, there can be

defined three major regions. The least affected area is called the partial grain-

refining region, which is followed by a grain-refining region and often the most

important grain-coarsening region [118]. A diagram of the HAZ zones of a low

96

carbon steel by Easterling [119] is presented as Fig. 2.30. Another classification

often used is presented as Fig. 2.31

Fig. 2.30 A diagram of the zones in the HAZ of a 0.15 wt.% C steel [119]

Fig. 2.31 Another classification of the low alloy steel HAZ subzones [120].

97

Depending on the distance from the fusion zone the HAZ goes through different heat

cycles. The coarse-grained region is the part that has been heated above 1100 °C.

The fine-grained region is heated to between 900 °C and 1100 °C. Finally, the

tempered region is heated to below 700 °C. Of these regions, the coarse-grained

region is the most vulnerable to embrittlement and is often called as the local brittle

zone (LBZ) [121].

In a multipass weld, the average grain size of the HAZ is smaller than in the fusion

zone. Therefore when the fusion zone is replaced by the HAZs of the subsequent

weld passes, the fusion zone is “grain-refined” which is often desired [118].

In stainless steels, the metallurgical changes in HAZ are less prominent. The main

reactions involved are grain growth, ferrite formation, precipitation and grain

boundary liquation [122].

Grain growth is usually modest unless the heat input is exceptionally high. Ferrite

formation on grain boundaries in the HAZ can be beneficial as it restricts grain

growth and lessens susceptibility to liquation cracking. Ferrite formation is usually

sluggish and not very commonplace in welding, due to the fast heat cycles [122].

In the parts of HAZ which are heated near the solidus temperature, the precipitates

dissolve in the austenite. When they precipitate from the austenite again, they can

form at the grain boundaries and the ferrite-austenite interface. The most common

precipitates are carbides and nitrides. They may be challenging to detect using

metallography, but are very likely to be present in most austenitic alloys [98].

Extensive chromium precipitation can lead to degradation of corrosion resistance via

sensitisation induced intergranular corrosion [98].

98

Local melting, or liquation, of the austenite grain boundaries may result from the

segregation of impurities. In metals containing titanium or niobium, this may lead to

HAZ liquation cracking. Other impurities prone to segregation are sulphur and

phosphorus [122].

2.6.3.1 Carbon Equivalent

A commonly used way to estimate the properties of LAS HAZ is the Carbon

Equivalent (CE) which correlates to the achieved hardness. In a CE formula, the

hardening effect of alloying elements is compared with that of carbon and the

relevant alloy content is divided by a factor that gives the CE of that element.

Formulae for CE are displayed as ( 2.1 ) and ( 2.2 ) [91, 123].

𝐶𝐸 = 𝐶 +𝑀𝑛

6+

𝐶𝑢 + 𝑁𝑖

15+

𝐶𝑟 + 𝑀𝑜 + 𝑉

5

( 2.1 )

𝑃𝑐𝑚 = 𝐶 +𝑆𝑖

30+

𝑀𝑛 + 𝐶𝑢 + 𝐶𝑟

20+

𝑁𝑖

60+

𝑀𝑜

15+

𝑉

10+ 5𝐵

( 2.2 )

Equation ( 2.1 ) is applicable to plain carbon and carbon manganese steels. Equation

( 2.2 ), developed by the Japanese Welding Engineering Society [91], takes into

account more alloying elements.

The factors affecting fracture toughness of the HAZ are the thermal cycles, grain-

coarsening temperature, transformation characteristics, alloy content and non-

metallic content. The fracture toughness of HAZ can be better than the BM if the

fracture toughness of the BM is relatively low [91].

99

When considering SA508 it can be noted that Ni addition (0.96 %) improves the

toughness of the HAZ. Nickel lowers the ductile-brittle transition temperature.

Chromium (0.17 %) and molybdenum (0.47 %) affect toughness by modifying the

transformation characteristics. Carbon increases hardenability and decreases

toughness [91].

Kim & Yoon [124] state that traditionally, the HAZ adjacent to the weld fusion line

has been known to give lower toughness values than the other regions since the local

temperature peak from the welding process rises above 1100°C and produces the

coarse-grained microstructures in this region. However, they [124] conclude that

studies on this subject have shown that the toughness of the coarse-grained HAZ

region is not as bad as predicted but even better than that of the base metal. They

produced a weld of 220 mm thickness with a multipass narrow-gap SAW process.

Their conclusion was that in a case of multipass weld the minimum toughness values

were noted at the intercritically or subcritically reheated region of 4-5 mm apart from

the fusion line. Although the tests mentioned here were made with SA508 Class 1.3

this phenomenon will probably also occur in NGLW of SA508 Gr3 Cl2.

2.7 Weldability of Alloy 52 Filler Metal

High chromium up to 30 % wt. nickel alloy filler metals are desirable for improved

resistance to general corrosion, local corrosion and environmentally assisted

cracking. Higher chromium content Alloy 52 is more resistant to stress corrosion

cracking than the more common Alloy 82. It is, therefore, the preferred choice of

additive in nuclear power systems [125]. However, several weldability issues have

been observed in the use of high chromium nickel alloy filler metals [126]. The

welds are susceptible to ductility-dip cracking, (DDC) as the molten metal cools

100

down [84, 127]. Experience also shows that Alloy 52 is prone to hot cracking [82].

Weldability is compromised by a sluggish, viscous, weld pool [84] and sensitivity to

minor variations in alloying between different melts. This requires highly trained

welders. There is a tendency to lack of bond and lack of fusion. Ti and Al oxide

build-up may cause problems in multi-pass welding, by altering the shape of the melt

pool surface and preventing wetting of the sidewalls. The various types of cracking

the welds are susceptible to are listed in Table 2.18 [128].

Table 2.18 Cracking types in Alloy 52 and 52M weldments [128]

Cracking types in Alloy 52

weldments

Ductility dip cracking

Liquation cracking

Solidification cracking

The sluggish weld pool is difficult to manage. McCracken reported in 2012 that

magnetic stirring of the weld pool can help alleviate the problems and improve

joining, cladding and repair of nuclear power plant components [129].

Ti and Al oxides build up during multi-pass welding and can be trapped aligned to

cause rejectable ultrasonic testing indications. To alleviate this, a new type, Alloy

52MSS has been developed. This has a lower concentration of Al and Ti to reduce

the amount of their oxides floating on the melt pool [128].

McCracken 2010 [128] found also that lack of bonding LoB and LoF were most

common during the first layer of GTAW where the content of Cr was >24 % and

heat input was low. This was alleviated with increasing heat input among other

things.

101

2.7.1 Solidification Cracking

Solidification cracking is a form of intergranular cracking. It occurs at the final stage

of solidification when the tensile stresses developed exceed the tensile strength of the

almost completely solidified weld metal. The solidifying metal shrinks both due to

thermal contraction and solidification shrinkage. This contraction is prevented by the

surrounding solid base material, which is often restrained. Consequently, tensile

stresses develop. When solidification cracking occurs in nickel alloy weld metal it is

almost always along solidification grain boundaries, see Fig. 2.32 and Fig. 2.33 [98].

Fig. 2.32 Boundaries developing in austenitic weld metals [75]

102

Fig. 2.33 Migrated grain boundary (MGB), Solidification grain boundary (SGB)

and Solidification subgrain boundaries (SSGB) in austenitic Filler Metal

52 weld metal [75].

The various theories of solidification cracking agree that the solidifying grain

structure is separated by a continuous thin liquid film. This film is then ruptured by

the stresses induced, see Fig. 2.34 and Fig. 2.35. If a sufficient amount of liquid is

available, the cracks can get backfilled and thus “healed” [98].

103

Fig. 2.34 Examples of weld solidification cracking in Filler Metal 52M (ERNiCrFe -

7A) dissimilar weld overlays on a) carbon steel A36, and b) stainless steel

Type 304L [75]

104

Fig. 2.35 Hot cracking propagated to the surface of narrow groove laser hybrid

welding of AISI 316L-IG (ITER-grade) using Thermanit 19/15 filler [130]

In the case of dissimilar welds, weld metals that solidify as austenite and are fully

austenitic at room temperature are inherently susceptible to weld solidification

cracking. This is due to the segregation that occurs during solidification and the

tendency for liquid films to wet the austenite grain boundaries. With many Ni filler

materials dilution increases the susceptibility to cracking by the tendency of liquid

films to wet the grain boundaries. When dilution is low Alloy 52M has good

resistance for solidification cracking [75]. Avery has established limits for dilution to

avoid solidification cracking, see Table 2.19.

105

Table 2.19 Approximate limit of diluting elements in Fe – Ni – Cr welds. Adapted

from [131]

Weld metal Iron Nickel Chromium

Nickel 30% — 30%

Ni-Cr-Fe† 25% Unlimited 30%

The limit values should be treated only as guides. Absolute limits are

influenced by the welding process, weld restraint and small variations in

weld filler and base metal compositions

† Silicon should be less than 0.75% in the weld.

The impurity content of the weld can also influence the solidification cracking. If

steel base materials have high levels of S and P dilution can cause these impurities to

segregate during solidification increasing the susceptibility for solidification

cracking. Most Ni filler materials are extremely pure. As general rule the sulphur

should not exceed 0.010 wt% and phosphorus should be less than 0.020 wt%. At the

limit levels, the welding process requires careful control in heat input, dilution and

weld geometry. [75]

2.7.2 Ductility Dip Cracking

Ductility dip cracking (DDC) is a solid state phenomenon typically occurring in re-

heated metals between 400 and 900°C. A dip in tensile ductility, see Fig. 2.36, has

been reported in different alloys such as Ni-based and stainless steels. In Ni-Cr

alloys, the DDC typically occurs as intergranular cracks the length of one grain or

less [125]. The cracks develop at grain boundaries which have lower ductility than

the grain interiors [64]. Highly localised stresses have been suspected to trigger grain

boundary sliding as the cracking mechanism [132, 133]. This previously rare

phenomenon has become more common with the introduction of high Cr filler

materials [64, 98].

106

Fig. 2.36 Ductility as a function of temperature and ductility dip [75]

DDC is known to be a serious weldability issue in many nickel-based filler metals.

The issue is prominent at particularly highly restrained thick section multipass welds

such as is the case in the welds in this study. Fig. 2.37 displays a multi-pass weld in

FM52 having DDC. The cracking has occurred along the weld metal migrated grain

boundaries [75].

107

Fig. 2.37 DDC in multi-pass weld using Filler Metal 52. Arrows showing severe

DDC and recrystallization along migrated grain boundaries [75]

2.7.3 Liquation Cracking

Liquation cracking (LC) occurs in the partially melted zone. It is an intergranular

phenomenon, but can also occur along the fusion boundary. An LC crack forms

when the PMZ is weakened by grain boundary liquefaction and the solidifying weld

metal pulls it as it contracts during cooling. A wide solidification range, mushy zone,

108

increases the problem [98], as does the grain structure relative to the stress field.

The principle of liquation cracking is shown in Fig. 2.38 and an example is given as

Fig. 2.39.

Fig. 2.38 The principle of liquation cracking [98]

Fig. 2.39 Weld metal contraction tearing PMZ of 7075 aluminium welded with filler

1100 [134]

109

Kou [98] introduces four categories of sources for PMZ problems: filler metal, heat

source, degree of restraint, and base metal. Of these, in this study, the only variable

is the heat source. The smaller the welding energy the more narrow the PMZ is.

Multi-pass welding process in itself allows for smaller heat inputs to be used. Kou

also mentions arc oscillation during GTAW as a way to reduce the melt pool size

and to narrow the PMZ. As the base materials used in this study are not susceptible

to liquation cracking, liquation cracking may occur at previously welded Alloy 52

beads of the multi-pass weld, rather than in base materials and their fusion

boundaries.

2.8 Laser Welding of Dissimilar Materials

The potential of laser for dissimilar metal welds has been researched at least since

1968 [135]. Several studies have been made since [116, 136-138]. The subjects of

the studies have varied from mechanical properties of the weld, for example [137,

139], effects of shielding gas mixtures [140] to microstructural and fatigue properties

[141].

Simulation of different aspects of dissimilar laser welding has also taken place.

Mahmoud et al. 2012 have studied simulation of the temperature distribution in a

case of austenitic to a ferritic steel weld [142]. Ranjbarnodeh et al. 2011 used the

finite element method to model the effect of heat input on residual stresses in

dissimilar welds [143]. Hu et al. 2012 found that temperature fields close to the heat

source are very different in a case of laser spot welding of SS and nickel alloy when

compared to traditional spot welding [144].

Optimising the parameters of laser dissimilar welding with numerical analysis has

proven to be efficient by several studies [136, 145, 146]. Kraetzsch et al 2011

110

studied high-frequency beam oscillation to control the melting behaviour, seam

geometry, melt pool turbulence and solidification behaviour. They achieved crack-

free welds with good mechanical properties for Al-Cu dissimilar joints [147].

Sun 1996 [148] discusses the benefits and restrictions of laser welding. The low heat

input produces a small heat affected zone with limited residual stresses and

distortion. Subsequent rapid cooling can lead to beneficial fine microstructures

during solidification, but also non-equilibrium phases which can be detrimental. The

use of a laser will give precise control of the weld location and chemistry [148].

These comments are still valid today. Sun points out the classical critical points in

dissimilar ferritic/austenitic welding in Fig. 2.40. This description is made with

autogenous welding. When using filler material the composition and therefore

properties of the weld metal can be controlled, which can eliminate one possible

difficulty.

Fig. 2.40 Sketch of critical locations in the ferritic/austenitic laser welds [148]

Example of dissimilar laser welding is the laser weldability table of different metal

combinations cited from Belforte and Levitt, presented as Table 2.20.

111

Table 2.20 Laser weldability of binary metal combinations. E = excellent, G =

Good, F = Fair, P = Poor, * = no data available [149]

Certain generalisations have been made for real-life construction metals. Laser

weldability of high strength low alloy steels is regarded as being good. The

weldability for austenitic stainless steels is graded as being excellent with low

porosity and good corrosion resistance [150]. This applies only to similar metal

welds and can be challenged in the case of dissimilar welds. The welding methods

considered in this evaluation are conventional in terms of laser welding, which

means that the results are more likely to have been obtained by autogenous keyhole

welding than with a method which includes a filler material.

However, none of the studies mentioned above have considered NGLW. Most of

those were conducted without any filler material. The metallurgical, mechanical and

residual stress properties will differ significantly in a new type of process such as

NGLW. Certain phenomena such as the effects in HAZ can be similar but the

conditions in the fusion zone differ drastically when a filler wire of a third material is

introduced to the weld.

112

2.8.1 HAZ in Laser Welding

Martensite is normally not observed in the HAZ of low-carbon steel, as illustrated in

Fig. 2.41. High-carbon martensite can form when both the heating and the cooling

rates are very high, as in laser beam welding. The HAZ microstructure of AISI 1018

steel produced by a high power CO2 laser beam is shown in Fig. 2.42. At the least

affected part of the HAZ high carbon martensite (and, perhaps, a small amount of

retained austenite) is formed in the prior-pearlite colonies. The austenite formed in

these colonies during heating has not had enough time to allow carbon to diffuse,

and it transformed into hard and brittle high-carbon martensite during the rapid

cooling. Hard martensite embedded into softer ferrite can significantly degrade the

mechanical properties of the HAZ. In HAZ positions C and D, both the peak

temperature and the time available for diffusion increase. This allows the prior-

pearlite colonies to expand while transforming into austenite and form martensite

colonies of lower carbon contents during the subsequent cooling [98].

113

Fig. 2.41 HAZ of GTA weld of AISI 1018 steel [98]

114

Fig. 2.42 HAZ microstructure of 1018 steel by a high-power CO2 laser.

Magnification of (A)–(D) 415x and of (E) 65x. B, High carbon martensite

[98]

115

2.8.2 Laser Welding Parameters

The basic parameters in laser welding are welding power, welding speed and focal

point position. Also affecting the properties are the focal length, the laser wavelength

and the shielding gases, to mention a few. In the case of dissimilar welding, a set of

new parameters is introduced. There are the two different BMs and the possible filler

material. Also in several studies, the dilution and properties of the weld have been

controlled by diverting the laser beam to one of the materials [151-153].

Duley 1999 gives a checklist about parameters in laser welding. It is presented in

Table 2.21. The optimisation of these parameters can be done only in terms of

“good” and “bad”. In many cases, this optimisation involves sectioning the welds

and mechanical, metallurgical and microstructural analysis [150].

Table 2.21 Critical parameters in laser welding [150]

Critical parameters

Laser power

Amplitude

CW or pulse wave

CW plus pulse

Pulse shape and repetition rate

Focusing

Location of the focal spot on surface

Focus above or below surface

Depth of focus

Intensity distribution within spot on surface

Shielding gas

Gas type

Flow rate

Orientation relative to focus;

trailing, leading, coaxial

Root side

Nozzle flow pattern

116

One of the parameters in dissimilar welding is the beam deflection or offset. During

dissimilar welding of Ti-6Al-4V and AZ31B magnesium, it was found that deviating

the beam to one side of the butt joint had significant effects on the weld. Gao et al.

2012 state that the laser offset plays an important role in the process stability and

weld characterisation [152]. This may not affect NGLW as much as autogenous laser

welding, but it may have some effect.

Using electron beam welding (EBW), Lin et al. 2010 experimented dissimilar

autogenous welding of SUS-304L SS and nickel-based Alloy 690. They conclude

that deviating the beam 0.30 mm to Alloy 690 enhances the interdendtritic corrosion

resistance [153]. This can have a significant effect on laser welding also.

2.8.3 Laser Welding With Filler Material

In many cases, the stringent joint requirements are the major drawback of laser

keyhole welding. The use of filler material alleviates these challenges. In laser

welding, the filler material is usually introduced in the form of a filler wire.

However, the use of filler wire is often considered as too difficult for industrial

applications, having too many parameters and too stringent requirements for wire

positioning [154].

During laser keyhole welding of a low alloy steel, the requirements for filler wire

positioning were found to be relatively strict. Tolerance in the order of 0.2 [155] to

0.25 mm [154] in the positioning of the filler wire was found to be acceptable. The

inaccuracies can be compensated to a degree by increasing the heat input [156]. The

increased heat input widens the melt pool and helps the melting of the wire. The

studies on this subject have concentrated mostly on the metallurgical and mechanical

117

properties of the weld in a certain application [154]. These studies have been made

mostly with single pass keyhole welding.

The parameters of filler feed position relative to the laser beam (Wy, Wx, WZ),

workpiece surface (FZ) and joint during laser welding are depicted in Fig. 2.43. The

filler wire is usually introduced from the front of the weld pool. This gives more

even distribution of the filler material. The feed angle (αW) has been reported to be

between 30°-75° to the surface of the work piece. The geometrical features of the

setup often dictate the angle chosen [154]. In NGLW the feed angle cannot be at the

lower end of the scale due to the geometrical restrictions of a deep weld groove. In a

study of NGLW of SS, the angle was 45°[35].

Fig. 2.43 The parameters of laser welding with filler wire [154].

A relatively steep angle can be beneficial in some cases. When using leading feed a

tendency of uneven heat distribution has been found. This may lead to a beneficial

weld geometry that has a wider root than the top [156]. A greater portion of the heat

conducting deeper into a weld may be a benefit in the case of NGLW, but the uneven

heat distribution may cause issues.

118

A filler wire introduced into the keyhole process is not an optimal solution. The

interaction angle and power intensity are not optimal for the wire. The interaction

takes place outside of the focal point where power intensity is low. Also, the wire is

fed at such a speed that any pre-heating does not occur [154]. In practice in most

applications, authors don’t mention having significant problems with adding filler as

wire [35].

The angle of incidence of the laser beam is of great importance, as the absorption is

at its best at near 90° but decreases rapidly when deviating from perpendicular to the

surface when considering low alloy steels. The absorption increases with rising

temperature. Iron alloys vaporise when the beam intensity exceeds 106 W/cm

2. This

triggers the creation of the keyhole. When the keyhole is formed the reflectivity

closes to zero [154]. This and the possibility to reduce back-reflections to the optical

system allows for slight tilting of the laser head. An angle of 5° is mentioned in the

literature [20].

In a study, the highest welding speed was achieved with a trailing feed filler wire.

This is due to the wire being fed partly to the weld pool (instead of the focal point)

which has a high heat content contributing to the melting of the wire. In this case,

more laser power is introduced to melt the base material via the keyhole [156].

The wire diameter did not have a significant effect on weld quality. The use of thin

wire is recommended because its more compact size makes it possible to feed it to

the desired point [156].

119

2.9 Testing of Weldments

2.9.1 Destructive Testing

Tensile tests and impact toughness tests are the most important destructive tests

considering testing of weldments. A bend test, either in the longitudinal or in the

transverse direction is applicable in some cases. [123]

For tensile testing, a standardised flat specimen is cut out of the weld. This can be

done in a longitudinal or transverse direction. In transverse, composite tests, the

main observation is whether the breaking occurs in the weld zone, HAZ or base

material. The tensile strength measured correlates to the part which has failed.

Different variations of the test can be made, such as making several smaller test

pieces or cutting a piece longitudinally inside the fusion zone. [123]

The results are displayed as a stress-strain diagram. An example with key points is

presented as Fig. 2.44.

120

Fig. 2.44 Engineering stress-strain diagram [157]

Fracture toughness is the material’s ability to resist crack growth by plastic

deformation. Charpy V is the commonly applied impact fracture energy test for

weldments. Its benefits are simplicity and the ability to induce the crack in a

predetermined location allowing localised determination of properties. The test can

easily be conducted for pieces at different temperatures and the practice is to state

the toughness at graded temperatures. Standards commonly add flexibility by

allowing the use of reduced size samples. [158] [123] A diagram of a typical ductile

to brittle transformation curve is presented as Fig. 2.45.

121

Fig. 2.45 Ductile to brittle transition region. Energy shelves and fracture

descriptions [158].

In nuclear manufacturing, an important design parameter is the minimum fracture

toughness requirement for ferritic materials of pressure-retaining components of the

reactor coolant boundary of light water nuclear power reactors set by the United

States Nuclear Regulatory Commission in Appendix G to Part 50 of the NRC

Regulations (10 CFR) [159]. This applies to A) Low alloyed ferritic steel plates,

forgings, castings and pipes, B) Welds and heat affected zones of the mentioned

materials and C) Fasteners. The appendix states that the upper shelf energy (USE)

must be no less than 75 lb-ft (102 J).

Many standards, codes and regulations, such as ASTM E 185, RCC-M App. SI,

KTA 3203 and JEAC 4210 require that the Charpy-V notch should be located less

than 1 mm from the fusion line. However, studies have shown that the toughness of

the coarse grain HAZ may be even better than that of the base material. In 1998 Kim

& Yoon propose placing the notch at the intercritical or subcritical zone of the HAZ,

although they acknowledge that this does not correspond with the regulatory

standards or codes [124].

122

2.9.2 Non-Destructive Testing

By Hellier 2001 the definition of non-destructive testing (NDT) is an examination,

test or evaluation performed on any type of test object without changing or altering

that object in any way, in order to determine the absence or presence of conditions or

discontinuities that may have an effect on the usefulness or serviceability of that

object. [160]

NDT can be divided into different categories. Halmshaw 1996 lists visual,

radiographic, ultrasonic, magnetic, penetrant, electrical and other methods [161].

Hellier presents a slightly different list adding eddy current, acoustic and thermal

infrared testing. [160]

It is important for non-destructive testing (NDT) to understand that some flaws are

more significant than others. Certain flaws can be considered acceptable. Certain

methods can detect smaller flaws than others and acceptable flaw sizes are related to

the methods used [161]. In this study, the emphasis on NDT is on visual and

radiographic methods.

2.9.2.1 Visual Testing

Visual checks after welding are the first indicator of the success of the weld. Table

2.22 explains the details which should be checked after each welding.

123

Table 2.22 Visual examinations after welding [161]

Feature Check Notes for NGLW by author

Cleaning and

dressing

metal merges smoothly, is slag

removed

Penetration and

root

penetration over the full length,

root concavity, burn-through

and shrinkage grooves

fusion can be challenging in

NGLW

Contour regularity of the surface, height

of weld

dependent on the last pass in

NGLW

Width of the weld consistency over length Distortions affect the shape

Undercut can be measured accurately if

necessary

unacceptable

Overlap usually unacceptable

Stray arcing can cause local hard spots not applicable in LBW

Weld flaws cracks can usually be seen

visually

cracking is possible due to

high cooling rates

2.9.2.2 Radiographic Testing

Radiographic testing includes X-ray and gamma-ray testing. From the NDT

engineering viewpoint, they are very similar only the source is different. The main

practical difference is the penetration potential. The ability to penetrate material by

radiation is dependent on energy, material and the thickness of the material. Gamma

rays have more energy than X-rays and therefore can penetrate thicker specimens.

The denser the material, the less penetration is achieved. Metals as in this case steels

at greater thicknesses are challenging to test [160].

As the majority of industrial X-ray radiographic techniques use an energy of 100

keV to 400 keV the X-ray methods can practically penetrate only up to 3 inches

thick steel samples. Bigger thicknesses can be addressed with gamma-ray testing.

For example, Cobalt 60 decays emitting radiation at 1.17 MeV and 1.33 MeV which

in practice can penetrate samples up to 5 inches in thickness [160].

124

The discontinuities detectable by radiographic testing include cracks, lack of fusion,

incomplete penetration, inclusions and porosity. Also, certain geometric conditions

can be detected, such as concavity, convexity, undercut, underfill and

overreinforcement. This is noteworthy as the radiographic images are by far the most

important permanent records achieved by NDT [160].

2.9.3 Measurement of Residual Stresses and the Contour Method

Residual stresses can be measured with many methods. Methods include X-ray

diffraction [103], hole drilling, neutron diffraction and the contour method [106,

162]. NDT methods for determining residual stresses usually measure some

parameter related to the stress. Many methods are based on diffraction, which is used

to measure the elastic strain of specific atomic lattice planes. Conventional X-ray

diffraction is limited to a depth of 5 µm due to poor penetration capability of the

wavelength of the X-rays. Where larger penetration is needed neutron diffraction is

usually applied giving the ability to penetrate up to 50 mm steel. [163]

Advances in synchrotron X-ray sources have made very intense high-energy (80 –

300 keV) X-rays available. They are only slightly inferior to neutrons for penetration

in steel, but due to their very high intensity, they offer a better spatial resolution.

Synchrotron techniques can be considered more suitable for mapping experiments

for FE model validation while neutron diffraction is more suited to deep line stress

profiling [163].

Other NDT methods for evaluating residual stresses are based on the influence of

stress on the magnetic properties, conductivity, speed of sound, Raman excitations or

optical fluorescence. Based on lasers, the last two have a high spatial resolution,

whereas the former are more suitable for bulk testing with lesser resolution. Three of

125

the methods mentioned are worth noting as they are the ones widely available,

namely the magnetic, high energy synchrotron and contour methods. [163].

Withers 2008 [163] compared three different residual stress measurement techniques

in Table 2.23. The contour method is shown as a process which is still

developmental. Remaining challenges listed are concerns of the quality of the

contour cutting process and the effect of cutting induced plasticity to the results.

However, developmental work has been done on understanding these issues for

example by Sun 2017 [164]. Geometrical and resolution capabilities of different

methods are presented in Fig. 2.46. It shows that the contour method is well suited to

the purpose of this study.

126

Table 2.23 Comparison of three strain mapping methods [163]

Magnetic Synchrotron Contour

NDT YES YES NO

Scope Suitable for in-situ

plant

measurements and

stress state

monitoring

Test-pieces and mock-

ups can provide detailed

maps as a function of

life suited to

testing/validating FE

models

Suited to ex-situ

examination of

plant components

as well as test-

pieces

Capability Ferritics: near-

surface < 10 mm

Samples less than 35mm

thick. Best suited for in-

plane strain

measurement in thick /

thin plates.

Components must

fit within EDM and

CMM

Maturity-

repeatability

Mature but a

unified theory

relating magnetic

signals to basic

magnetic

parameters lacking.

At present signals

are equipment

supplier-specific

Very mature; Bragg’s

equation provides a

straightforward link to

lattice strains from

which stress can be

derived

Tools for repeatable

extraction of stress

fields from

measured profiles

under development

Remaining

challenges

Can be sensitive to

microstructure

effects especially

in welds; for

accurate result

requires extensive

calibration tests on

alloy using biaxial

loading

Can be difficult to apply

for coarse-grained

materials. Need

industrial standard for

repeatable measurement,

validated software

analysis procedures and

instruments dedicated to

engineering strain

measurement.

Sensitive to

material movement

during the cutting

process. This may

lower the maximum

stresses and distort

recovered stress

field if not

prevented. The

importance of

plasticity is yet to

be quantified

127

Fig. 2.46 Schematic indicative of the approximate current capabilities of various

techniques. Destructive techniques shaded grey [163].

The contour method residual stress mapping method published by Prime [162] in

2001 is a destructive method for generating a map of the perpendicular residual

stresses of a cross section. Johnson [165] describes the contour method as an

elegantly simple technique with modest instrumentation requirements but laborious

in the data processing stages.

The contour method consists of experimentally cutting the sample, Fig. 2.47, to

allow the residual stresses to relax as translations, Fig. 2.48. The resulting

deformation is measured scanning the resulting surface contour and then analytically

forcing the resulting surface back to flat, Fig. 2.49. The resulting forces are displayed

as a 2D map [166, 167].

128

Fig. 2.47 The intact sample with original stresses [166]

Fig. 2.48 The cut sample with deformation induced by the residual stresses [166]

Fig. 2.49 The analytically flattened surface with residual stress map [166]

The ideal cut would be straight, not remove any material from the already cut

surfaces and not induce any additional residual stresses. Wire EDM is proven to

meet these requirements [168] and is widely used [11, 46, 167, 169, 170]. The

methods for scanning the distortion include using a coordinate measuring machine

129

(CMM) [171, 172] and laser profilometry [46, 173]. The analytical post-processing

can be made for example using Matlab scripts presented by Johnson in 2008 [165].

The contour method has been shown to produce data with very good correlation to

other residual stress measurement methods [167, 169]. It has been found that

although the general profile of the stress distribution measured by contour method

gives reliable readings, the peak values at steep gradients of change in the stress are

strongly affected by the data smoothening. Therefore using the method requires

careful judgement between noise removal and actual data retention.

Plastic deformation during cutting has been found a major cause for systematic error

using the contour method [108, 164, 167, 174]. Plasticity errors have been shown to

increase when the stress magnitude approaches the yield strength of the material. It

may cause a shift in the map towards the cutting direction [174]. To counter the

plasticity effects, the cutting strategy has to be well thought out [175] and the

samples should be well restrained [167].

2.10 Standards for evaluation of welds in nuclear applications

Standards used for evaluating the welds in the current study are listed in Table 2.24.

130

Table 2.24 Standards used for evaluation of the weldments

Standard Description

ASME IX Boiler and Pressure Vessel Code

BS EN ISO 13919-1:1997 Beam welded joints

BS EN ISO 148-1:2016 Charpy impact toughness

U.S. NRC NRC10 appG Fracture toughness requirements

BS EN ISO 6507-1:2005 Vickers hardness testing

ASTM E384-17 Microhardness testing

ASTM E8 Tensile testing

The “Section IX of the ASME Boiler and Pressure Vessel Code relates to the

qualification of welders, welding operators, brazers, and brazing operators, and the

procedures that they employ in welding and brazing according to the ASME Boiler

and Pressure Vessel Code and the ASME B31 Code for Pressure Piping [176].”

The ASME IX code is a vast source of information. It presents many of the

commonly used welding practices, procedures and nomenclature. It defines the

welding positions. It presents tensile test and bending test procedures for each

thickness for groove welds as well as for fillet, spot, laser lap joint, flash and stud

welds. The code presents radiographic and Charpy V-notch impact toughness test

acceptance criteria [176].

To establish an industry-standard assessment for the integrity of the welds in this

study a radiographic acceptance according to ASME IX was applied. The code

defines two types of indications, linear and rounded. Linear indications are defined

longer than 3 times their width and include cracks, incomplete fusion, inadequate

131

penetration and slag. Rounded indications may be circular, elliptical or irregular in

shape and represent porosity and different inclusions [176].

The following indications will be unacceptable (thickness t = 40 mm) [176]:

A) Linear indications

1) Any type of crack or zone of incomplete fusion or penetration

2) Any elongated slag inclusion over 1/3 t

3) Any group of slag inclusions in line that have an aggregate length greater

than t in a length of 12t, except when the distance between the successive

imperfections exceeds 6L where L is the length of the longest imperfection in

the group

B) Rounded indications

1) The maximum permissible dimension shall be 1/5 T or 3 mm whichever is

smaller

2) For materials thicker than 3 mm the code presents charts for acceptable

types of rounded indications, see Fig. 2.50. However, indications smaller than

0.8 mm shall be omitted.

Fig. 2.50 Appendix I of ASME IX defining the acceptable appearance of rounded

radiographic indications [176]

132

To further evaluate the quality of the welds in addition to ASME IX also BS EN ISO

13919-1:1997 [177] was used. It describes permitted imperfections for laser and

electron beam welds in three categories; moderate, intermediate and stringent. The

standard considers welding imperfections such as cracks, porosity, lack of fusion,

undercut, filling, misalignment etc. see Fig. 2.51.

Fig. 2.51 Example of BS EN ISO 13919-1. The maximum height of undercut

The BS EN ISO 148 [178] presents a standard procedure for pendulum impact

toughness tests. It defines the standard sample geometry and test conditions. ASME

IX refers to the Charpy V-notch test procedure to be used for impact toughness

verifications [176]. ISO technical specification ISO/TS 7705 provides guidelines for

implementing the ISO 148 for steel [179].

The Charpy impact test uses a standardised pendulum to impact the test sample. The

standard V- notch sample is presented as Fig. 2.52.

Fig. 2.52 Charpy V-notch impact toughness sample according to BS EN ISO 148

133

In the case of mild and low alloyed steels, the Charpy impact test will result in an

impact test curve, Fig. 2.53. The curve is plotted following: X-test temperature, °C,

Y-absorbed energy, J. The results form three distinct zones, 1-lower shelf, 2-

transition zone and 3-upper shelf. A specified energy 4 can be presented at a

specified temperature 5. The scatter of results on the upper and lower shelves is

relatively small but the scatter in the transition part of the curve is relatively large

[179].

Fig. 2.53 Example of an impact test curve

Many materials, such as stainless steels are ductile even at very low temperatures

and instead of a curve, their results yield a straight rising line without a distinct steep

incline for the transition zone [179].

U.S. NRC 10 CFR Appendix G to part 50 [159] specifies fracture toughness

requirements for ferritic materials of pressure-retaining components of the reactor

134

coolant pressure boundary of light water nuclear power reactors. It states that the

requirements also apply to welds and heat affected zones of the materials.

The code specifies that reactor vessel materials and welds must have Charpy upper

shelf energy of no less than 102 J and must maintain a level of no less than 68 J

throughout the operation.

BS EN ISO 6507-1:2005 specifies the Vickers hardness test for metallic materials

[180]. The test is performed by indenting the material with a standardised diamond

indenter, see Fig. 2.54, with a specified force. The resulting indent is then measured.

The Vickers hardness is proportional to the quotient obtained by dividing the test

force by the sloping area of the indentation [180].

Fig. 2.54 Principle of the Vickers hardness test

Considerations for the Vickers test to ensure reliable results include a minimum

material thickness and a minimum distance between indentations and to the edge of

the test piece. ASTM E384-17 specifies the minimum distance to be 2.5 dV, where

dV is the Vickers indent diagonal [181].

135

ASTM E8/E8M − 16a specifies test methods for metallic materials at room

temperature. It specifies the nomenclature, procedure and test samples [182]. The

rectangular test piece is presented as Fig. 2.55.

Fig. 2.55 ASTM E8 rectangular tensile test sample [182]

The dimensions of the sample are very flexible and using subsize specimens is

acceptable. The key dimensions are G-gauge length, W-sample width and T-

thickness [182].

The tensile test results in a stress-strain diagram, see Fig. 2.56. The key figures are

the ultimate yield strength (UYS), the ultimate tensile strength (UTS) and the

elongation.

136

Fig. 2.56 Stress-strain diagram [182]

2.11 Summary of the Literature Review and Rationale for the

Current Work

A comparatively small amount of relevant publications considering NGLW were

found. During the literature review the following publications considering narrow

gap laser welding were found directly relevant to this study:

Dittrich, et al, 2013 [21] “Laser-multi-pass-narrow-gap-welding of hot crack

sensitive thick aluminium plates”

Elmesalamy, 2013 [35] “Narrow Gap Laser Welding of 316L Stainless Steel

for Potential Application in the Manufacture of Thick Section Nuclear

Components”

Elmesalamy, et al., 2013 [51] “Understanding the process parameter

interactions in multiple-pass ultra-narrow-gap laser welding of thick-section

stainless steels”

137

Elmesalamy, et al., 2014 [11] “A comparison of residual stresses in multi-

pass narrow gap laser welds and gas-tungsten arc welds in AISI 316L

stainless steel”

Feng, et al., 2016, [45] “Narrow gap laser welding for potential nuclear

pressure vessel manufacture”

Ficquet, X. et a.l., 2009, [102] “Residual Stress Measurement on a Narrow

Gap Dissimilar Metal Weld Pipe”

Jokinen et al. 2003, [44] “High power Nd:YAG laser welding in

manufacturing of vacuum vessel of fusion reactor”

Jokinen, 2004, [17] “Novel ways of using Nd:YAG laser for welding thick

section austenitic stainless steel”

Kong, et al., 2013 [19] “Feasibility study of laser welding assisted by filler

wire for narrow-gap butt-jointed plates of high-strength steel”

Yu, et al., 2013 [18] “Multi-pass laser welding of thick plate with filler wire

by using a narrow gap joint configuration”

Zhang, X. D. et a.l., 2011 [20] “Welding of thick stainless steel plates up to

50 mm with high brightness lasers”

By a very small welding groove volume, NGLW is proposed to offer geometrical

benefits over conventional welding. These include less filler material required,

smaller distortions and shorter process times. This study investigates these

experimentally.

NGLW offers a significant improvement over the conventional laser keyhole

welding, as the thickness of the weld can be increased without the need for

increasing the laser power. Conventionally laser welding of thick section steels has

required approximately 500 – 1000 W per 1 mm of penetration [51]. With NGLW

138

there is no such technical lower limit, as each welding pass melts only a relatively

thin bead [183]. Thicknesses up to 50 mm have been reported being welded using

NGLW [20].

The austenitic stainless steel to ferritic low alloyed steel dissimilar welds commonly

use nickel-based filler materials. A key feature of these welds is the ferritic to

austenitic microstructure fusion line, which has a very different microstructure from

conventional similar metal welding fusion line. This fusion line and the adjacent

HAZ can suffer from different issues like microfissures, carbon migration/depletion

and ductility dip cracking. The development of these possible flaws during NGLW is

investigated.

There are several nuclear codes and standards for welding. The resulting welds are

evaluated using the industry-standard codes for quality. These analyses consist of

optical, radiographical, and mechanical analysis.

The key findings of the literature review can be summarised as:

NGLW is expected to cause smaller welding distortions and residual stresses

leading to longer service life.

Elmesalamy et al. 2014 [11] reported reduced residual stresses in NGLW

compared to GTAW of stainless steels. Reduction of residual stresses is

beneficial to counter stress corrosion cracking. Primary water SCC of

dissimilar welds in PWR reactors has been reported to be a significant issue

for safety, longevity and costs of the power plants.

NGLW will reduce the number of welding passes compared to MMAW and

GTAW.

Welding time will be reduced as the number of passes reduces. This will lead

to savings in labour costs. The reduction of passes also leads to fewer

inspection cycles.

139

Gap volume reduction from NG-GTAW and especially MMAW will grant

significant material savings.

Welding without buttering is justifiable. Buttering is performed to enable

PWHT of the ferritic steel HAZ before adding the austenitic stainless safe-

end. The rationale is to prevent sensitisation of the austenitic steel component

during the PWHT. Real-life low carbon austenitic stainless steels were found

to have very low carbon content, and therefore be capable of withstanding the

required dwell times at elevated temperatures without sensitisation.

2.11.1 Knowledge gaps

A knowledge gap was established as no publications considering dissimilar metal

NGLW were found. The similar metal welding research published does not answer

many of the questions rising in DMW.

There is no data on the behaviour of the challenging materials, such as FM52, in

NGLW DMW’s. There is no published research on the resulting grain structures,

which will be characteristically different in DMW. The multi-pass tempering effect

in NGLW will cause variation of mechanical and microstructural properties along

with the depth of the weld and has not been addressed for this type of materials. The

mechanical and the residual stress properties of this type dissimilar metal NGLW

have not been studied.

The following questions arise as the main interests of this study:

Is NGLW able to produce acceptable weld integrity for nuclear

manufacturing?

Does the impact toughness meet the acceptance criteria?

What are the resulting microstructures, are they sound?

140

What is the effect of multi-pass tempering; is there variation in hardness,

grain structure, tensile strength or impact toughness between samples near

the root, middle or top of the weld?

Can an improvement on the residual stresses be achieved compared to

conventional methods?

The work described in the following chapters aims to provide answers to these

questions.

141

3 Materials, Methods and Equipment

This chapter describes the welding programme structure. It presents the choice of

materials and the methods for experimental work. The construction of the welding

setup with its various challenges is discussed. The monitoring and measurement

practices during the welding process are presented. The methods and equipment for

analysis of the welds is described.

3.1 Welding Programme Stages

The welding conducted in this study can be divided into three main stages, in which

different materials were used. The first stage was to use similar metal welding of

316L stainless steel to develop the welding procedures and equipment. The second

stage was a series of dissimilar metal welding trials using S275 mild steel as a

cheaper substitute to the SA508 steel. Last, when previous stages were completed

satisfactorily, the dissimilar metal welds were made using SA508 Gr3 Cl2.

The three welding programme stages are described, with the materials applied, in

Table 3.1. Detailed material information is presented in appropriate chapters.

Alongside the monitored welds, numerous simpler tests were made to evaluate

changes in the equipment and parameters. All monitored welds are presented

individually in Appendix I.

142

Table 3.1 Welding programme stages and corresponding materials used

Stage I II III

Identifier TV DS SA

Base

materials

316L +

316L

S275 +

316L

SA508 Gr3 Cl2 +

316L

Filler

material

ER 316L

Alloy 52 Alloy 52

Description Stainless steel similar

metal welds to

develop equipment

and procedures

Mild steel

dissimilar welds for

further process

development

Pedigree steel

dissimilar welds for

full analysis

3.2 Materials

The dissimilar metal welds investigated in this study consist of ferritic low alloyed

steel and austenitic stainless steel. These base materials are normally fusion welded

using nickel alloy filler material. Ferritic LAS is a typical material for pressure

vessels in a PWR. Stainless steel is used for the PCC piping. Nickel alloys are

commonly used filler materials for dissimilar welding due to their resistance to creep

failure in prolonged elevated temperature use. Narrow gap laser welding using these

materials has not been reported in the literature.

3.2.1 Stage I – Similar Metal Welding in Stainless Steel

In the similar metal welding trials, 316L austenitic stainless steel was welded using

ER 316L filler wire. The composition of the base material was analysed by energy

dispersive spectroscopy (EDS) by Elmesalamy 2013 [35]. The analysis is presented

in Table 3.2. The literature states the nominal composition of ER 316L, it is

presented in Table 3.3.

143

Table 3.2 EDS analysis of the 316L stainless steel used throughout the study, wt.%

[35]

C Mn Cr Ni Mo N S

316L 0.027 2.02 16.68 9.9 2.03 0.05 0.02

Table 3.3 Nominal chemical composition of ER 316L, wt.% [184]

C Mn Si Cr Ni Mo S P Cu

ER

316L

0.04 –

0.08

1.0 –

2.5

0.30 –

0.65

18 –

20

11 –

14

2 –

3

0.03

max

0.03

max

0.75

max

3.2.2 Stage II – Dissimilar Metal Welding with Mild Steel

To start the dissimilar metal welding trials a preliminary, cheaper substitute mild

steel S275 was used instead of the actual nuclear grade LAS. The filler material was

Alloy 52 Ni alloy. The mechanical properties of the materials used are described in

Table 3.4. Table 3.5 and Table 3.6 describe the chemical compositions of the

materials. The 316L stainless steel was analysed by EDS and its composition is

presented earlier as Table 3.2.

Table 3.4 Mechanical properties of materials used [185] [80] [186]

AISI 316L FM52 S275

Tensile strength (MPa) 530 – 680 536 410 - 560

Yield strength (MPa) 240 240 275 -

Elongation (%) 40 30 23

Equivalent standards 1.4404 (X2CrNiMo17-12-2) ERNiCrFe-7 1.0044

Table 3.5 Chemical properties of the base materials used [185-187]

C max Mn max P max S max Si max Cr Ni Fe

316L 0.03 2.0 0.045 0.03 1.0 16.00 – 10.00 – Rem.

144

18.00 14.00

S275 0.25 1.6 0.04 0.05 0.5 - - Rem.

Table 3.6 Chemical properties of the filler material used [80]

Ni + Co Cr C Ti Mn Al Fe

FM52 Rem 28.0 –

31.5

0.04

max

1.0

max

1.0

max

1.10

max

7.0 –

11.0

P S Nb+ Ta Si Al + Ti Mo Cu Others

0.02

max

0.015

max

0.10

max

0.50

max

1.5

max

0.50

max

0.30

max

0.50

max

3.2.3 Stage III - Dissimilar Metal Welding with Pedigree Steel

The final welding trials to produce the samples for full analysis programme was

welded using the target materials, AISI 316L, Alloy 52 and SA 508 Gr3 Cl2. Table

3.7 describes the mechanical properties of the materials used for welding the

stage III of welds. Table 3.8, Table 3.9 and Table 3.10 show the chemical properties.

Table 3.7 Mechanical properties of materials used in the SA series of experiments

[185], [80], [188]

Tensile

strength

(MPa)

Yield strength

(MPa)

Elongation

(%)

Equivalent standards

AISI 316L 530 – 680 240 40 1.4404 (X2CrNiMo17-12-2)

FM52 536 240 30 ERNiCrFe-7

SA508

Gr3 Cl2

701 576 24

Table 3.8 EDS analysis of the 316L stainless steel used, wt.% [35]

C Mn Cr Ni Mo N S

316L 0.027 2.02 16.68 9.9 2.03 0.05 0.02

145

Table 3.9 Analysis of SA508 Gr3 Cl2 material used, wt.%, Appendix IV

C Si Mn P S Cr Mo

SA508 0.18 0.24 1.32 0.019 0.022 0.24 0.53

Ni Cu V Nb Ca B Ti Al

0.47 0.052 0.003 0.004 0.0004 0.0001 0.001 0.003

Table 3.10 Analysis of Inconel Alloy FM52 material used, wt.%, Appendix V

C S P Si Mn Ni Cr

FM52 0.025 <0.001 0.003 0.14 0.24 60.46 28.87

Co Mo Cu Fe Ti Al N2

0.003 0.01 0.01 8.98 0.52 0.72 0.005

3.3 Sample Geometry

The samples used for the welding trials were 200 mm long and 40 mm thick, see Fig.

3.1. The width of the samples was 100 mm each, which was determined by using

calculations and the clamp manufacturer specifications. The choice of plate form

samples instead of a self-restraining tubular geometry was chosen as several authors

in literature were found to have used it for their studies [11, 45, 189] and it was used

as a standard approach in the NNUMAN programme [6]. It was acknowledged that

this approach is suitable only for preliminary studies as the restraint conditions vary

considerably when using different real-life geometries.

146

Fig. 3.1 Dimensions and design of NGLW samples.

3.4 Welding Groove Design

To fully exploit the benefits of NGLW the welding groove geometry was decided to

be kept as narrow as possible as this would increase efficiency; accelerate the

production speed and reduce the amount of filler material needed. The main

limitation was seen to be the nozzle design, sufficient rigidity of a thin nozzle being

difficult to achieve. Special nozzles for the wire feed and shielding gas delivery had

to be developed. After some study, it was decided to use a 5 mm groove width to be

able to construct gas and wire nozzles with sufficient rigidity.

In literature certain authors had used a straight wall welding groove [45, 56], many

used a straight V-groove [18, 20, 43].

First welding tests with 40 mm thick 316L stainless steel were conducted with 5 mm

wide parallel sidewall geometry to explore the groove contraction under the

147

experimental setup conditions. The resulting measurements after each welding pass

are presented as Fig. 3.2. The measurements were made at A 1/3 and B 2/3 along the

length of the weld. It can be seen that the groove contracts 1.3 mm by average. This

indicated a need for a slight V-groove to counter the effects of butterfly distortion

experienced during the process.

0 2 4 6 8 10 12 14 16 183

4

5

Ga

p W

idth

(m

m)

Pass no:

A [mm]

B [mm]

Fig. 3.2 5 mm parallel groove contraction, TV4

For the following trials a 4° V-groove, Fig. 3.3, was chosen as the basis of the

development. The aim was to keep the groove 5 mm wide for each actual pass by

anticipating the distortion caused by the welding.

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Fig. 3.3 The 4° V-groove design for countering groove contraction

Later in the programme, the restraint conditions changed somewhat. This was due to

the introduction of the Lenskes welding clamps and the change of materials. In the

end, a 5° V-groove was needed to counter the distortions in the dissimilar metal

welds. This was detected by methodically monitoring the distortions of each weld.

The choice of a backing strip design was justified by the fact that in all dissimilar

applications considered in this study the aim is to produce a layer of weld metal

between the base materials [9, 190, 191], hence preventing using an autogenous

welding pass at the beginning. Several researchers used sharp [11, 18, 20] instead of

filleted corner designs for the bottom of the NGLW groove instead of the more

conventional filleted root design for NG-GTAW [6, 36]. The economical and time-

saving benefits of the straightforward machining of the sharp corners were

considered to outweigh the risks as the possible imperfections would be machined

off in real applications. The welds were analysed without machining in this study,

and no issues were found at these corners.

149

3.5 Welding Parameters

The NGLW in this study were made using the laser in conduction mode. This means

that the power density at the focal spot would be kept below the threshold of

evaporation. Literature states a power density of 5 x 10^5 W/cm2 as the upper limit

for conduction limited welding mode [40]. This was however found to be just a

guideline, as factors such as heat conduction conditions, wire feed rate and welding

speed were found to affect the propensity to evaporation.

The welds conducted shared some basic specifications. The same sample geometry,

the effective width of the welding groove and the shielding gas BOC Pureshield

Argon (99.9998% purity) were used throughout the tests. The parameters which

were changed over the tests were laser power, travel speed, spot size, wire feed rate

and shielding gas flow rate.

The laser power was indicated by the laser system. The travel speed was set using

the robotic system. The wire feed rate was controlled by the closed-loop wire feeder.

Generic rotameter gas flow meters were used for gas flow control.

For the similar metal stainless steel welds, a set of parameters was developed, see

Table 3.11. The parameters were optimised for maximum bead thickness to reduce

the number of passes thus allowing for the highest possible deposition rate and fast

processing time.

In the dissimilar metal welds the parameters, Table 3.11, were developed based on

an unpublished technical report of NG-GTAW. The characteristics, such as the heat

cycle to which the materials were exposed were replicated. It was found that the

parameters as such did not yield satisfactory results but development was required.

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Table 3.11 Base parameters for NGLW

Similar Metal Welds Dissimilar Metal Welds

Laser power 8.14 2.8 to 3.5 [kW]

Welding speed 0.48 0.08 [m / min]

Welding wire feed rate 3.5 1.0 [m / min]

Argon flow rate 20 Up to 30 [l / min]

In both cases, the final few passes required tailored parameters. For example, the

capping pass was welded wider to ensure good smooth fusion. This required the use

of a larger spot size, which was compensated with higher laser power to maintain the

power density. The filler deposition rate was adjusted to avoid excess reinforcement,

but to fill the weld fully. The capping of the weld was considered of relatively low

importance as in the real world applications welds of this type would be finished by

machining.

In dissimilar metal welding, the parameters were varied according to the results of

each pass. The first pass was usually welded with higher power than the following

filling passes. This was to ensure the fusion of the square groove shape and in a

condition where the heat transfer from the separate backing plate to the base material

sidewalls was not optimal. As the bottom of the weld with the backing structure

would also be machined after completion of a production weld, the less than optimal

shape was considered acceptable. The main benefit of this solution was a much-

simplified groove design. In practice issues with fusion at the bottom of the welding

groove were rare.

Re-melt passes were sometimes used to remedy issues in the weld. These were

usually welded with the same parameters as the previous pass, but without

introducing filler wire. Issues which were treatable with re-melting were lack of

151

fusion and unevenness of the bead among others. One important observation of re-

melting was that a re-melt pass causes as much distortion as a filling pass. Therefore

a carefully optimised welding groove design may become too narrow if re-melting is

applied.

3.6 Welding Setup

The setup for thick section NGLW was developed based on literature, mainly on the

work by Elmesalamy [35] and Feng [45]. The narrow gap laser welding setup

consisted of a laser welding head, which was manipulated by a robotic arm, and a

sample which was restrained to a welding table. Auxiliary devices, such as wire

feeding nozzle were mainly mounted to the welding head. Larger apparatus like the

filler wire feeder or parts of the monitoring system were mounted externally. The

thick section weld required a substantial restraint system to counter the welding

distortions.

3.6.1 Laser and Robot Used

The laser used in this study was a 16 kW IPG YLS ytterbium fibre laser and the

welding head was Precitec with 300 mm focal length optical system. The laser was

conveyed to the welding head via means of an optical fibre. The fibre diameter was

300 µm and the collimation focal length 150 mm. The head was manoeuvred with a

KUKA KR 30 HA industrial robot. The welding setup was developed constantly

during the course of this study. An example of an early welding setup is pictured in

Fig. 3.4.

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Fig. 3.4 The laser welding setup during wire feeder nozzle prototype trials

3.7 Development of the Equipment and Procedures

Narrow gap laser welding is a novel welding process. The required equipment is not

commercially available. Therefore a considerable portion of the research had to be

dedicated to developing the welding equipment and setup.

Producing the welds posed unique demands for the welding equipment. The welding

groove was deep and unusually narrow, hence reaching to the bottom required long

focal length laser optics and special nozzles for the filler wire and shielding gas. The

thickness of the weld necessitated the use of a highly rigid welding restraint to

manage the welding distortion forces. Observing the welding distortions was done

using HandySCAN 700 3D scanning and mechanical measurements. Video

monitoring of the welds was developed to allow real-time viewing and recording of

Wire

feeder

unit

Wire

feed

conduit

Wire feed

nozzle

Laser welding

head

Optical

fibre

(yellow)

Rotary

table and

shielding

gas

nozzles

Restraint

clamps

153

the welding process. Originally the alignment of the robot for laser head movement

was programmed by using the low power pilot laser. Later the coaxial camera system

was applied to improve precision.

Numerous different shielding gas nozzle arrangements were experimented. Two wire

feeder units were tried as well as various wire feeder nozzle setups. Four different

restraint systems were used. During the tests, three different video camera systems

were developed and applied and several mechanical distortion measurement

strategies devised.

3.7.1 Welding Restraint

The most significant component of the welding setup was the restraint system. The

properties of the restraint had a direct influence on the distortions occurring during

welding. These, in turn, determined the welding groove design. It was established

that safety, rigidity, reliability and repeatability were the main characteristics of a

good welding restraint.

The welding restraint system was developed during the similar metal weld tests. The

buttress clamp setup used to restrain the samples was found satisfactory for up to

30 mm material thickness. However, the distortions observed were severe enough to

indicate the need for a more rigid setup for thicker section welds. The rigidity of the

table used was also suspected. The first 40 mm weld TV4, using similar metal

stainless steel, was restrained with a strongback arrangement and the butterfly

distortion was well constrained. Later a setup with 3 pairs of 42 kN welding clamps

was used on an 80 mm thick steel table.

154

The first tests and trials up to 10 mm stainless steel were conducted by using

machining clamps on the rotary table of the robotic system, as in Fig. 3.4. However,

they were found to be less than satisfactory in their capability to prevent distortion

and concerns arose about their safety in case of overloading.

Welds up to 30 mm thickness were still restrained with clamps on the rotary robot

system table. The clamps used were Carver buttress clamps, Fig. 3.5, with a

maximum loading of 62.5 kN. They were found to be much stiffer. The 60 mm thick

robot table was identified to be a major contributor to the distortion during thick

section welds.

Fig. 3.5 Sample TV1 restrained by one pair of Carver buttress clamps

The flexibility of the buttress clamp setup was found unsatisfactory. The mounting

threads on the robot table were too close to each other to allow clamping the samples

from the outer edges. To work around the less than optimal clamps available at the

time, a strongback arrangement, Fig. 3.6 and Fig. 3.7, was trialled. Two steel plates

were welded to the actual weld sample. As a backing strip was used, machining of

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the strongbacks was required. This process of machining and manual GTAW of the

40 mm thick strongbacks along with the backing strip took over 8 weeks. Although

the resulting restraint was good, the process was found to be too slow and dependent

on too many external factors that it was determined to be inapplicable in the long

run.

Fig. 3.6 Strongback welding arrangement for TV4. Note copper shielding gas

nozzles on the table and separate brackets to contain the gas atmosphere

for the last passes.

156

Fig. 3.7 Strongback welding restraint arrangement for TV4 viewed from below.

Note the machined slot to accommodate the backing plate and the

extensive manual GTAW-welding required.

A calculation for integer value for the number of pairs of clamps, ( 3.1 ), revealed a

need for a stronger setup to ensure the safety of the work. In the equation NC =

number of pairs of clamps, PL = plate length, PW = plate width, PH = plate height,

σy = yield strength and CN = maximum allowed force on the clamp. An 80 mm

thick steel table, Fig. 3.8, with T-grooves, was obtained with 3 pairs of 42-ton

welding clamps. The clamps were tightened to 100 Nm using a torque wrench. This

setup was then used throughout the dissimilar metal welding programme giving

reliable and repeatable restraint.

𝑁𝑐 = ⌈𝑃𝐿𝜎𝑦𝑃𝐻

2

4𝐶𝑁𝑃𝑤⌉

( 3.1 )

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Fig. 3.8 42 kN Lenskes welding clamps restraining a dissimilar metal weld sample

DS4 on an 80 mm thick Lenskes T-groove steel welding table. Note the

location of the clamping points near the edges of the sample, the pre-

heating blankets and run-in and run-out blocks to guide the shielding gas.

3.7.2 Wire Feed Arrangement

The fundamental purpose of the wire feeding system is to deliver the welding

additive to the molten pool to fill the welding groove at a controllable rate. The feed

rate and the point of delivery have to be controllable accurately. The wire feed rate is

controlled by a wire feeding unit, which uncoils and pushes the wire through a

flexible conduit at a set speed. The conduit is connected to a nozzle at the laser head

which guides the wire to the desired point.

Two different wire feeder units were used during this study. First, a basic stand-

alone wire feeder TecArc F4, Fig. 3.9, was used. The wire was fed from the

stationary feed unit to the moving welding head through a long conduit. This setup

had many drawbacks. The long conduit changed shape as the weld progressed

changing the alignment of the actual tip of the wire. The wire feed speed provided by

the feeder unit was also found to be unreliable, as there was no closed-loop control.

158

Fig. 3.9 TecArc F4 wire feeder and the wire conduit (red).

The second wire feed unit was a more sophisticated, closed-loop control capable

Jetline Engineering series 9600 CWF-50Z, Fig. 3.10. The design of the unit made it

possible to mount the feeder on the robot arm, providing a much shorter conduit with

much less change in the geometry of the wire conduit during the welding process and

therefore much more reliable performance. The control unit, Fig. 3.11, was separate

and was mounted at an easily accessible location. The closed-loop control provided

steady and repeatable wire speed unaffected by friction variations in the wire guiding

system.

159

Fig. 3.10 Jetline Engineering 9600 wire feeder unit mounted on top of the robot

Fig. 3.11 Jetline engineering 9600 wire feeder control unit

At the beginning of the tests, several wire feeding parameters considering the wire

feed nozzle design were settled by doing bead on plate tests using 316L samples and

filler wire. The wire feeding arrangement would be a leading wire type and the wire

feeding angle was determined to be less than 45°. In the bead on plate tests, the

reliability of the arrangement quickly deteriorated with steeper angles causing

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inadequate wire melting and buckling. Maximum working depth was set according

to target material thickness. The challenge was to keep the width of the nozzle as

narrow as possible up to the desired thickness. The interaction point between the

wire and the molten pool had to be adjustable in all three dimensions. Also, the angle

of the wire had to be controllable. These adjustments were designed into the wire

feeding nozzle assembly at the laser head.

The contradicting demands for the wire feeding assembly were challenging to meet.

The rigidity of the wire feeder nozzle was improved continuously and several

different setups were tested. Even the ease of alignment was found to be less

important than adequate rigidity. Several different designs were experimented. One

particular challenge was that the wire tended to re-coil when exiting the nozzle tip.

This would point the wire tip towards an undesirable direction unless a bend was

added near the tip, to align the wire centrally.

The wire feeder (left in Fig. 3.12) was abandoned mainly because the sharp bend

needed at the lower part created unacceptably high friction for the wire, causing

uneven feed speed, vibrations and buckling of the wire at the feed rollers. The wire

feeder nozzle on right in Fig. 3.12 shows stepless sliding brackets which allowed for

adjustment of the height and distance of the wire to achieve proper alignment of the

beam and wire. The aligning of the wire feeder nozzle to the centreline of the weld

was achieved by rotating the bracket. The actual nozzle was built in two components

the flat main copper nozzle supports a thin pipe for guiding the wire. The angle of

the wire αW was adjustable by bending the final bend of the wire feeding pipe. The

angle achieved with this setup was variable at 30° +- 5°. This nozzle was shared with

ferritic steel welding research and was designed to be capable of welding up to 130

161

mm deep grooves. The nozzle thickness was 5 mm with the tip of the wire guide

diameter 3.2 mm.

Fig. 3.12 Two examples of different experimental wire feeder versions. A) A simple

feeder pipe attached to the welding head (arrow) and B) the MTRL NGLW

nozzle which was used in welding the TV-series stainless steel tests

The first wire feeder nozzles were found to have several issues including instability

of the wire tip due to uneven friction, poor adjustability and poor survivability from

reflections of the laser beam. It was also realised that a rapid and reliable means of

adjustment of the Z-axis (FZ) was crucial for the ability to finely adjust the spot size

between passes and for the much wider capping passes. In the next stage, a setup

using precision linear stages was developed.

A commercial wire feed nozzle holder, Fig. 3.13, was obtained. This allowed for

precise, 1 mm per revolution, adjustment of the wire position in three dimensions

(Wy, Wx, WZ). Also, the angle αW of the nozzle was adjustable. The access to the

B

A

A

A

162

deep NGLW groove was solved by introducing a sufficiently long and thin nozzle.

The design weld thickness was 40 mm. The drawback in this system was that the

stroke of adjustment of each axis was only 30 mm. This was sufficient for X- and Y-

directions, where the adjustments usually fell within a few mm, but insufficient for

the Z. The Z-axis or the height of the wire had to be adjustable by at least 100 mm

due to the large changes in wire height during the exceptionally large offset required

for the last passes of the welds. The last setup was augmented with a 300 mm linear

stage to allow for this.

Fig. 3.13 Final wire feeder nozzle bracket (silver) and 300 mm stage (black) used

The wire nozzle tip was developed in three stages. First, a thin copper tube was used

as the nozzle, as per Fig. 3.12 b. This was very versatile and highly adjustable. It

provided acceptable rigidity, being mainly limited by the other components of the

setup used at that time. After a period of use, this tip was found to have a detrimental

flaw. Even the smallest reflections of the laser beam from the melt pool damaged the

163

nozzle immediately, melting several millimetres of its length and depositing the

molten copper into the weld. This also led to extensive repair work, as the whole

nozzle had to be replaced and the setup completely re-aligned.

The second version of the nozzle tip was designed using a slimmed-down

commercial contact tip for GTAW. This was based on a 6 mm diameter stainless

steel tube, which was threaded to accept the replaceable tip. The tip width was 4 mm,

tapering up gradually to the nominal 6 mm of the tube. This setup had three benefits

over the previous design. Firstly, the tip was much heavier, having a larger heat

capacity and therefore it was considerably less prone to detrimental melting.

Secondly, the tip was quickly replaceable in case of any indications of damage,

without the need to do extensive re-alignment. Thirdly, the stainless steel tube was

found to increase the rigidity of the tip, reducing wire vibrations considerably.

Working in the 5 mm wide 4° to 5° V-groove was found viable with the 6 mm tip

setup. An immediate improvement over the previous design was observed. However,

the tip width did not allow for enough adjustment as the wire was often found

bending to one side after exiting the nozzle. A tight-fitting 0.9 mm diameter tip was

used to minimize the curving of the wire after exiting the nozzle. Over time the

friction in the nozzle tip was found to be excessive causing incidents of buckling of

the welding wire at the wire feeder rollers interrupting the welding process. The soft

Alloy 52 welding wire was extremely prone to this.

To overcome the issues with the second design, an improvement was designed with

a smaller overall diameter, see Fig. 3.14.

164

Fig. 3.14 The improved 5 mm diameter wire feed nozzle

This tip holder was designed using a 5 mm tube and a milled down nozzle tip. The

tip had to be fully re-machined to reduce the thread to M4. The nozzle tip thickness

was reduced to 2.8 mm. This arrangement doubled the practically available travel

allowing for precise alignment of the wire-melt pool interaction point.

Also, at the time when the 5 mm tip holder was introduced, a wire straightener was

installed to the system. The wire straightener was placed immediately after the feed

rollers on the robot arm. By carefully adjusting the straightener, only minor

corrections were needed by the tip alignment. This improved the reliability and

usability of the wire weed system significantly.

3.7.3 Shielding Gas Setup

In literature Feng et.al.2017 [46] completed 30 mm SA508 steel NGLW with a

single leading shielding gas nozzle. Zhang et. al.2011 [20] used a single trailing

nozzle. Elmesalamy 2013 [35] welded 316L NGLW with a leading nozzle and a

trailing gas shower tube. In this study, the stainless steel similar metal welds trials

indicated that single fast-flowing shielding gas jets would drag in ambient air and

cause oxidation. This was found harmful for the quality of the similar metal welds,

and unacceptable for the more sensitive DMW. Fast shielding gas jets were also

suspected of causing disturbance to the weld pool.

165

Severe contamination by oxidation, vaporised metal and burnt contaminants was

found in the first bead on plate test runs. This was assessed with a multitude of

arrangements, involving stationary and moving nozzles of various designs. Only the

shielding gas, 100 % argon, was not changed during the development.

First, a single nozzle of 6 mm diameter was placed at 45° from horizontal in trailing

position behind the laser spot to provide the shielding gas. The nozzle was attached

to the welding head and the flow was directed 2 mm above the melt pool according

to literature. The results were unsatisfactory, so improvements were developed.

To keep the oxygen in the surrounding atmosphere from mixing with the shielding

gas jet, two designs of gas shoes were experimented. The shoes were attached to the

laser head and were, therefore, moving with the welding process. Gas shoes are

pictured in Fig. 3.15.

Fig. 3.15 Different gas shoe designs. A) A version constructed for first trials.

B) A proposed design to alleviate dust accumulation issues.

These provided good protection from the oxidation, but they did not allow the fumes

created by the high-intensity laser beam to escape. This caused contamination of the

test pieces and better solutions were needed.

A

A

B

A

166

At the beginning of the shielding shoe development, it was understood that a shoe

type setup couldn’t supply a reliable flow of shielding gas to a deep and narrow

NGLW groove. Therefore a parallel design process was started to create a blade type

welding apparatus to provide both shielding and wire feeding to the bottom of a deep

groove. Fig. 3.16 depicts a blade setup for delivering both the shielding gas and the

welding wire, attached to the laser welding head.

Fig. 3.16 Blade shielding gas and wire feed design

Developing the blade type nozzle was found to be very challenging. The maximum

thickness allowed by the narrow gap was very small and the brass blades would have

to be cut with EDM. The holder bracket for the blades would have to be machined

from steel. Both of these would have to incorporate several separate functions. After

numerous design iterations, the construction was deemed to be too expensive and

time consuming to manufacture. The idea was shelved as overly complicated.

167

Consequently, a far simpler and more flexible setup was developed. It consisted of

two stationary large-diameter shielding gas nozzles and steel blocks for guiding the

gas. These copper nozzles were placed on the welding table at the ends of the

workpiece. The heavier than air shielding gas was guided into the weld groove with

the aid of steel blocks close to the edges of the weld. The nozzle diameter was 13

mm and the length of the large-diameter tube was 130 mm to reduce the turbulence

of the flow. These nozzles were placed on the top of the test piece. This shielding gas

setup is pictured in Fig. 3.17. This setup was found to work satisfactorily with

stainless steel similar metal welding.

Fig. 3.17 Shielding gas setup used during welding of TV4 set up with stationary

nozzles aided by steel blocks to guide the gas flow. The nozzles were

raised according to the progression of the weld.

Gas shielding of the Alloy 52 based dissimilar metal welds proved challenging. The

well-known tendency of Alloy 52 to generate Al and Ti oxides [87, 89, 128] was

detrimental to the quality of fusion. The amount of oxides generated during a single

pass was not found detrimental, but the gradual accumulation of the oxides after 12

to 14 passes caused fusion problems regularly. In some cases, especially when the

laser power was low, the oxide layer was also trapped between passes, causing lack

168

of fusion type defects. Generally, the light oxides, as well as gas bubbles, had

enough time to float to the surface of the weld pool.

Physically removing the oxides was found to be an effective approach to improving

the resulting weld integrity and was regularly used. Each bead was wire brushed

after welding. The dust produced was removed by compressed air.

Using pre-heating also reduced the effectiveness of the stationary nozzles by

generating a strong rising thermal column of air, thus destabilising the gas

atmosphere generated by the nozzles. A more sophisticated design had to be

developed for use in dissimilar metal welding.

The final gas shielding setup was based on the blade principle. It was realised that

the only way to avoid the ejector phenomenon to drag ambient air and therefore

oxygen into the weld was to inject the gas very close to the weld pool and the

cooling weld bead. Nozzles capable of reaching the bottom of the 40 mm welding

groove were sourced and the design was based on these.

Firstly, the commercial nozzles had to be thinned to fit the welding groove. The

originally 5 mm thick nozzles were worked down to 3 mm. A nozzle was placed on a

bracket on the welding head. It was positioned to follow the melt pool as close as

possible without being damaged by the heat. Later another nozzle was added after

that and the nozzles ended up covering nearly 100 mm of the cooling weld bead.

During trials with one trailing nozzle, it was found that a leading nozzle was required

to prevent air from entering the weld from the front. Four different designs were

experimented. Development of the leading nozzles was found to be challenging due

to the wire feed nozzle occupying the space. Gas nozzles placed between the wire

169

feed nozzle and the welding pool were the in the way of the reflections from the melt

pool and were found easily damaged by the reflected laser beam. To avoid this, the

leading nozzle was placed under the wire feed nozzle, effectively combining them to

a single unit. This arrangement was able to avoid being damaged by the reflected

laser beam, but still capable to provide shielding gas within 15 mm of the melt pool.

The distance from the bottom of the nozzles to the bottom of the groove was kept

below 5 mm. The final gas shielding and wire feeding setup is pictured in Fig. 3.18.

Fig. 3.18 Final gas shielding and wire feed setup

3.7.4 Welding Procedure Development

Programming of the robot was continuously developed to improve the welding

procedure. The usability of the robot script was improved. A pre and post gas flow

was applied. A delay after firing the laser was introduced for allowing time for the

melt pool to develop and widen before starting the traverse. The wire was fed to the

stationary melt pool to increase the cross-section. After a good wetting of the side

walls was established the weld traverse was started. This made wetting of the

170

sidewalls more reliable and therefore improved the sidewall fusion, especially at the

beginning of the dissimilar metal welds. Stainless steel welds were not found prone

to problems with this as good quality cross-sections may be made within less than

5 mm from the beginning of the weld.

3.7.5 Alignment of the Weld and Equipment

The alignment of the robot to the sample was originally done by eye and the pilot

laser. By careful execution, this procedure proved satisfactory for stainless steel

welding. During the dissimilar metal welds, this was challenged. The accuracy of the

method was found to be insufficient and the increasing number of nozzles in the

weld groove made observing the actual alignment very difficult.

A coaxial camera system was implemented to improve the precision of the

alignment. This provided a view through the welding head coaxially with the

processing laser. To align the image to the laser a test firing of the laser was

performed producing a spot weld on a sample. Crosshairs were then centred to the

produced spot weld. The crosshair was then available for aligning the weld, with

good precision. The crosshairs are pictured during a test weld in Fig. 3.19. It can be

seen, that in this case, the alignment is wrong by 0.5 mm as the right-hand side

bottom corner of the welding groove is not melted properly. The coaxial camera and

crosshairs allowed the weld to be aligned with a precision matching the repeatability

of the robotic system.

171

Fig. 3.19 Cross-hairs aligned to the centre of the laser spot seen during a trial weld

1) Top of base materials, 2) Welding groove sidewalls, 3) Backing plate.

Later the Phantom Miro IV high-speed camera was used in the coaxial system to

allow recording of the resulting videos. This replaced the original analogue camera

system and the crosshair feature was not available. In this case, the camera itself was

aligned to the weld. This was found less convenient, but also very robust and re-

adjustment was rarely required.

3.8 Monitoring of the Welding Process

The welds conducted were monitored using several techniques. The temperature of

the samples was monitored. The welding distortions were measured after each pass

to develop the welding groove design. The weld pool behaviour was observed using

laser illumination imaging and the coaxial camera system of the welding head. These

methods were crucial to the development of the welding process and weld pool

management.

1 1 2 2 3

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3.8.1 Interpass Temperature Measurement

The pre-heat and interpass temperature measurements were conducted with a

thermocouple thermometer from the middle of the sample. The measurement was

made 20 mm offset from the weld centreline, see Fig. 3.20. In cases of dissimilar

metal welds, the measurements were made on both base materials.

Fig. 3.20 Measurement of the pre-heat and interpass temperature

3.8.2 Measurement of Welding Distortions

The transverse welding distortions were measured using various methods to help the

development of the welding groove design. The groove width was measured directly

between the top corners of the welding groove. An indirect contraction measurement

was calculated from measurements over a wider area. 3D scanning was implemented

to support the mechanical measurements.

Direct groove width measurement by a calliper was applied after each pass, as well

as two indirect measurements between points on the top surfaces of the samples. The

173

measurements were found to be concordant. This observation was later utilised by

concentrating on the direct measurements and one type of point measurement.

Choice of measurement locations was limited by the three pairs of welding clamps

occupying much of the space on top of the sample. Direct manual measurements of

the gap top width were made at two locations at 1/3 and 2/3 of the length of the

weld. Indirect measurements were made between points punched to the top surface

of the samples approximately 20 and 75 mm from the centreline of the weld, see Fig.

3.21 and Fig. 3.22.

Fig. 3.21 Punch mark locations for indirect distortion measurements

174

Fig. 3.22 Conducting the indirect gap measurement between punch marks.

The direct gap width measurement provided reliable data until the groove top corners

would get rounded by heat. In some cases, the edges were damaged prematurely by

contact with the laser beam. This was due to the groove contracting faster than

predicted and the laser beam touching the sidewalls. This was corrected by changing

the groove design.

The indirect measurement using the punch marks was continued until the completion

of the weld. The indirect gap was calculated as per ( 3.2 ), in which Gn = gap at pass

n, G0 = original gap pre welding, dn = measurement at pass n and d0 = original

measurement pre-welding.

𝐺𝑛 = 𝐺0 − (𝑑0 − 𝑑𝑛)

( 3.2 )

A hand-held 3D scanner was used to measure the welding distortions. The butterfly

angle measurement of the top surface was studied. The scanned model was divided

175

into 200 cross sections, Fig. 3.23, which were analysed to create graphs of the

butterfly angles.

Fig. 3.23 200 measurement line pairs for butterfly analysis on a 3D scanned model

of sample TV4. Note the omitted central area to avoid the effect of local

deformation near the weld.

The scans were performed after each pass to monitor the development of the

distortion. The 3D scanning was mainly performed to support the groove design

development. It was regularly performed at the beginning of the tests but later

omitted as it was found too time-consuming for reliable interpass temperature

management.

3.8.3 Laser Illumination Imaging

The laser safety regulations did not allow access to the welding cell during the

process. This made visual monitoring of the welding process impossible. Early in the

tests, the only monitoring system available was a security camera in the cell. It was

soon found that a way to monitor the welding process remotely was required.

176

Monitoring the progressing weld was found to be crucial for observing the weld pool

phenomena and evaluating the equipment in real-time and retrospect.

Laser illumination imaging (LII) is a novel way of monitoring welding processes.

The images are free of glare by the hot glowing gases or electric arc. The digital

camera systems used allow for easy adjustment and recording of the video material.

An LII system consists of a single wavelength laser light source, a narrow bandpass

filter and a digital camera. The working principle is that the target is illuminated by a

powerful light of a particular wavelength and only this wavelength light passes to the

camera sensor through the bandpass filter. All other light is omitted to produce an

image free of disturbance from ambient light sources. The laser used as the light

source is usually pulsed and the camera shutter has to be synchronised to the light

pulses. This is achieved by using a high-speed camera which can be triggered frame

by frame.

First, a laser illumination imaging system was borrowed from EPSRC equipment

loan pool (EIP). The setup consisted of a Cavitar Cavilux illumination laser and a

Phantom MIRO IV high-speed camera. This system provided state-of-the-art

monitoring of the welding process. In the Cavilux system, the illumination laser

aperture unit was a small cylindrical unit at the end of an optical fibre. This was

fitted to the welding head parallel with the processing laser. The camera was fitted

next to the light source. The triggering was provided by the camera system, the

Cavilux acting as a slave. The welding head setup is pictured in Fig. 3.24.

177

Fig. 3.24 Laser illumination imaging setup on the welding head. Experimental LED

light source pictured, Cavilux laser aperture unit similar in size

As the EPSRC EIP closed in mid-2016 a replacement for the Cavitar Cavilux LII

system was needed. The MIRO IV high-speed camera was received as a donation

from the EIP, but a new light source had to be sourced. First a high power single

wavelength LED pictured in Fig. 3.24 was trialled, but it proved to be too low-

powered for satisfactory operation.

Later an Oxford Laser illumination system became available from the EPSRC EIP.

This was then received as a donation. The Oxford Laser illumination system differed

from Cavilux by having no optical fibre to guide the laser. This meant that the large

laser source itself had to be pointed to the weld. The laser unit was mounted

stationary on a tripod. This proved cumbersome but was applied with success to

several welds. The resulting image quality can be seen in Fig. 3.25 depicting a weld

pool 36 mm deep in the welding groove with no glare or other disturbances.

Camera

lens

Light source

178

Fig. 3.25 A frame from Laser illumination imaging video. 3rd

pass of test DS1. 1)

Top of the stainless steel plate, 2) Top of the ferritic steel plate, 3) Weld

bead, 34 mm below the previous, 4) Wire feed nozzle and 5) welding

groove side wall (dark area).

As the gas shielding arrangement became more intrusive by the introduction of the

trailing gas nozzles, the Oxford Laser LII system had to be removed as there was no

way to deliver the illumination light to the welding groove. The previously untried

built-in coaxial camera system was put into use, using ambient light from the

welding process.

The coaxial camera system in the welding head had previously been examined for

use with the LII systems. Two main issues had prevented the application. The main

problem was that in the coaxial optics there was no space for the bandpass filter,

which is essential to the operation. Secondly, it was found that the optical system

was not designed to focus at the exceptionally large focal offset used in conduction

mode welding.

During ambient light coaxial imaging, the inability to focus on the subject was

overcome by using the smallest aperture available in the optics. This provided the

maximum depth of field and the image obtained was found satisfactory for

observation of some important phenomena, see Fig. 3.26.

1

2

3 4

5

179

Fig. 3.26 A frame from the coaxial camera system video. 1st pass of test SA5.

1) Welding groove sidewalls, 2) Filler wire and 3) wetting of the sidewall

One major limitation of all the imaging systems used was the limited memory of the

MIRO IV camera used. The 2 GB memory was only able to record the full length of

the weld at a frame rate of 50 fps. High-speed imaging was possible only for shorter

periods, preventing the recording of the whole welding pass. The inflexibility of the

Oxford Laser light source placement was also a big issue. By using a more modern

larger memory camera and an optical fibre delivered illumination these issues would

be easily overcome.

Wire feeder stability, alignment of the weld and equipment, wetting of the sidewalls,

oxidation, smoke generation (overheating), undercut formation, sidewall or nozzle

damage by reflections, overheating of the groove top corners were among the issues

detected and rectified by using the imaging systems. The resulting recordings were

imperative for the development of a successful welding process.

1

1

2

3

180

3.9 Measurement Uncertainties

3.9.1 Laser Power

In the IPG laser system, the desired laser power was set fundamentally as a

percentage of the full power. For clarity, the percentage was expressed in kilowatts.

Due to the operating principle and the construction of the laser system, there was a

slight difference in the laser power set using the software compared to the emitted

power monitored by the software and furthermore to the actual power at the

workpiece.

The emitted power was monitored by the IPG laser system and it was displayed as

each weld was being made. The IPG system displays the output power based on the

pump current for the diodes. As the efficiency of the diode system varies, the pump

current is not a fully reliable measure of the laser power. Another source of error was

that the system cannot take into account the losses in the beam delivery system, for

example, the losses at the fibre connections and inside the welding head.

To assess the issues mentioned above, a calibration of the system was required. In

Fig. 3.27 a calibration chart is presented. The chart displays the relation of set power

(Analogue Input) to the displayed power (Monitor) and actual power (Actual) on the

workpiece. The power measurements were made with the current setup of the fibre

delivery system, such as the fibre, collimator and the welding head. For the

calibration chart, the power at the workpiece was measured by a Primes

PowerMonitor 100 laser power meter. The calibration was executed during the laser

system acceptance trials before this study begun.

181

Fig. 3.27Laser power calibration chart

The fibre laser source used in this study consists of 16 1 kW fibre modules. These as

a whole are protected for back reflection from the workpiece. Occasionally a single

module may develop a fault due to back reflection or other reasons and is

automatically disconnected from use. This phenomenon was found to occur only

when power near the maximum 16 kW was used, which was not the case in the

welds in this study, although it happened several times over the course of this work.

The laser system reacts to the absence of a module by increasing the power of other

modules, allowing the laser to be used within the reduced power range available. The

reverse happens as the modules are repaired and returned to use. This response,

however, is not entirely perfect and some change in power results. Over the course of

this work, this offset in laser power varied between +5.3% and -9.7%.

The laser power was found stable during a single welding pass. A maximum

variation of 0.7% was recorded. This was considered insignificant. When the laser

power was changed, the offset also changed. At worst, the variation was 3.4% for

weld DS2. For other welds of the study, the variation was within 1.0%.

182

3.9.2 Spot Size Control

The spot size control was compiled from spot on plate tests, the LII videos and was

measured manually from the experiments. Using a calculated spot size for the

records was considered, but the long working distance, working at large offsets, was

found to give considerable discrepancy between the theoretical and practical spot

diameters. One phenomenon not taken into account, if a calculated value would have

been used, was the change of the focal length of the optics due to heat during the

process, also known as focus shift [192, 193]. An increase of 40% in focal spot

diameter was reported just after 50 s at 7 kW laser power [193]. Also, the variation

of laser power, which could’ve caused a change in the effective spot size, was

considered small, especially during the critical filling passes.

The size of the actual effective spot size was kept just small enough to prevent

damaging the upper corners of the welding groove. This gap was then measured with

a calliper and the laser head offset was recorded. The spot on plate tests were used as

a starting point. The measurement accuracy for the spot size was considered to be

±0.2 mm, in the case of measurement from the welding groove top gap.

3.9.3 Accuracy of Welding Consumable Feed

The wire feed speed of the feeder used in early trials was found prone to fluctuation.

This was partly due to the friction in the long conduit required and the bending of the

conduit during a welding pass. The issue was solved by introducing a closed-loop

control feeder unit mounted on the robotic arm, reducing the length of the feed

conduit to 1 m and lessening the changes in bends during welding. The speed of the

wire feed was calibrated by allowing the feed to run 1 min at the speed of 1 m/min.

The feed rate was found reliable to within 2%.

183

The accuracy of the generic gas flow rotameters used was +-10% stated by the

supplier.

3.9.4 Welding Robot Accuracy

The welding speed and positioning of the welding head were controlled by the

KUKA KR 30 HA (high accuracy) robotic system. The manufacturer states ISO

9283 repeatability of 0.05 mm [194].

3.9.5 Accuracy of the Alignment of the Welds

As the first welds were aligned manually using the pilot laser, the alignment error

was considered to be 0.2 mm. Later, when the coaxial camera was used for aligning

the welds as well as monitoring, the accuracy matched the accuracy of the robotic

system.

3.9.6 Welding Groove Dimensional Accuracy

The welding groove angle was machined by Agie Charmilles Fi400CCS multi-axis

wire cutting machine. The machine has a 0.05 µm resolution. The width of the

welding groove varied because of the manual welding required for the attachment of

the backing plate. This variation was measured to be ±0.3 mm.

3.9.7 Accuracy of the Distortion Measurements

The mechanical measurements, such as the direct and indirect distortion

measurements were done by a calliper. Accuracy of the calliper was 0.02 mm [195]

and by comparing repeated measurements on test spots the repeatability of the

measurements was considered to be 0.05 mm.

184

The accuracy of the Handy Scan 3D scanner was 0.030 mm with a resolution of

0.050 mm [196]. These values were stated as maximum, the ISO 10360 volumetric

accuracy was 0.020 mm + 0.060 mm / m.

3.9.8 Summary

Table 3.12 displays typical parameter and measurement values and their errors.

Table 3.12 Typical values of different parameters and measurements with errors

Typical

Value

Error Error %

Laser Power 3500 W ± 35 W 1.0

Spot Size 5 mm ± 0.2 mm 4.0

Welding Wire Feed Rate 1.0 m / min ± 0.01 m / min 1.0

Welding Speed 0.48 m / min m / min

Positioning Accuracy n.a.* 0.050 mm

Argon Flow Rate 20 l / min 2 l / min 10

Welding Groove Angle 2.5 ° 0.05 µm

Welding Groove Gap 5 mm 0.30 mm 0.6

Mechanical Measurements ~50 mm 0.05 mm 0.1

3D Scanning n.a. 0.020 mm

* not applicable

3.10 Analysis Methods for the Weldments

The integrity of the welds produced was evaluated using several methods. The

evaluation was begun with an industry-standard X-ray analysis and complemented

with further assessment of grain structure, residual stresses and mechanical

properties.

All welds were subjected to visual examination. The visual examination results were

supported with process monitoring data, mainly videos and groove behaviour

measurements to assess the outcome. If the weld was found free of flaws it was sent

185

to X-ray analysis for more precise evaluation. The decision for further analysis was

made using these results.

To utilize each sample as effectively as possible, the order in which the analysis

would be conducted had to be planned carefully. The first destructive test would be

contour cutting as this would require cutting an intact sample. Other destructive tests

were conducted after the contour scanning was completed and the data quality found

satisfactory for processing.

3.10.1 X-ray Radiography and Acceptance

Welds in this study were radiographically analysed by X-ray radiography as per BS

EN ISO 17636-1: 2013, ASME V Article 2:2015, British Engineering Services 02-

111-I063 and British Engineering Services 02-111-I085. The welds were subjected

to acceptance by BS EN ISO 13919-1 [177] & ASME IX QW 191: 2015. ASME IX

describes a standard method for analysing welds in nuclear applications and was

conducted to provide an industry-standard assessment of the integrity of the welds.

Radiography was found to be a reliable and quick way to establish the quality of the

welds. Using a non-destructive method was essential to save the sample integrity for

the consequent mechanical analysis, especially for the residual stress measurements.

Issues caused by welding flaws were also avoided by X-ray analysis. For example,

severe lack of fusion or other similar discontinuities were avoided to reduce the risk

of EDM wire breakages during contour cutting.

3.10.2 Microstructural Analysis and Triple Etching

The machining of samples for advanced destructive testing was performed mainly by

EDM. Some simpler samples for microstructural and hardness analysis were cut

186

manually using a band saw and a disc cutter. Usually, the backing plate was left

attached to the sample, see Fig. 3.28. The microstructural analysis was conducted

entirely on full-size cross-sectional samples. These were extracted at points of

interest in the weld.

Fig. 3.28 Typical sample for microstructural and hardness analysis, thickness

varying from 2 mm to 10 mm

The tri-metal nature of the cross-sectional samples required three different etchants

to be used to reveal the microstructures of the three different areas in the welds. The

etching procedure was performed in three separate stages from the weakest etchant

to the strongest.

The first etching stage was to etch the ferritic steel segment of the cross-section. This

was achieved by using 2% Nital. The sample was swabbed with cotton wool

immersed in the etchant for 5 – 7 s and rinsed and dried immediately.

For the subsequent stages, the previously etched parts of the sample were covered

with Lacomit lacquer to protect them from the strong etchants which would

otherwise destroy the surface. The lacquer was applied manually, leaving as narrow

a strip as possible of the already etched material uncovered. This was to ensure

187

etching to reach the fusion line, but simultaneously limit the contamination of the

etchant by excessive by-products of the rapidly etching low-alloyed material.

The second stage was performed to etch the Ni-alloy weld metal. For this 10%

Ammonium persulfate in aqueous solution (Rawdon etchant) was used. The process

was electrolytic etching at 6 V. The processing time was 5 - 10 s.

The third stage of etching was using 2% Oxalic acid in aqueous solution to reveal the

microstructure of the stainless steel. The other areas were covered with Lacomit. The

etchants and processes used are listed in Table 3.13.

Table 3.13 Etchants, compositions and applications

Etchant Composition Type Target Duration

Nital 2% HNO3 –

Ethanol

Chemical Ferritic steel 5 – 7 s

Ammonium

persulfate

10% (NH4)2S2O8-

Water

Electrochemical 6 V Nickel alloy 5 – 10 s

Oxalic acid 10% C2H2O4 -

Water

Electrochemical 6 V Stainless steel 40 – 50 s

3.10.3 Hardness Mapping and Hardness Line Scanning

There are several phenomena causing hardening of steels in welds. Formation of

metastable martensite phase caused by rapid cooling and work hardening due to

strain were considered in this study. Generally, hardening is considered as negative,

as it leads to a decrease in ductility. Hardening of the weld metal and the heat

affected zones were to be expected. The hardness testing was used as a valuable tool

for phase recognition in the microstructural analysis.

Two types of hardness testing were applied to the samples. Firstly, a hardness map

was generated over the whole sample to display the overall hardness change due to

188

the welding process. After the large scale maps, more detailed hardness testing was

conducted to reveal small scale phenomena and to support microstructural analysis.

The hardness tests were made using a Struers DS-80 890 automated hardness tester.

The hardness was measured in the Vickers scale, using small indentation forces

0.3 kg and 1.0 kg, regarded as low-force Vickers hardness. Equation ( 3.3 ) [180]

describes the Vickers number calculation.

𝐻𝑉 =1

9.80665×

2𝐹𝑠𝑖𝑛136°

2𝑑2

( 3.3 )

To avoid interaction between indentations caused by work hardening, the distance

between indents was increased by using an alternating multi-row pattern instead of a

straight line of indents. Standards ASTM E384 [181] and ISO 6507-1 [180] give the

following Table 3.14 for minimum distances.

Table 3.14 Minimum distances to prevent work hardening artefacts in hardness

testing. d = indent diagonal [180, 181]

Distance between

indentations

Distance to the edge of specimen

ASTM E384 2.5 d 2.5 d

ISO 6507-1 > 3 d for steel and copper

alloys > 6 d for light metals

2.5 d for steel and copper alloys >

3 d for light metals

The expected lowest hardness numbers were in 316L stainless base materials, 152

HV [197]. Using HV 0.3 scale this translated to an indent diagonal of 60 µm,

yielding minimum spacing of 150 - 180 µm. The mapping was made to reveal the

effect of the heat input to the material properties. As the welds in question are multi-

pass welds the effect of multi-pass tempering was studied.

189

Due to the large changes in hardness over a small distance in the ferritic steel HAZ a

more detailed analysis was required. The approach was to perform a line scan, with

multiple rows to maintain the minimum allowable distance between indentations.

For S275 mild steel line scan, an intricate multi-line test pattern was created. For

SA508 a map with 200 µm spacing was used.

3.10.4 Contour Method Residual Stress Measurement Procedure

The longitudinal residual stresses were evaluated using the contour method. The

contour method was chosen because of its good correlation to other residual stress

measurement methods like deep hole drilling and neutron diffraction [167, 169, 174,

198], especially in the longitudinal direction [102, 189]. Longitudinal

The method used was based on the procedure developed at The University of

Manchester. Several unique non-trivial improvements and manual work phases were

required due to the three material structure of the DMW samples. The whole process

was completed in-house. The main steps of the contour method are described in Fig.

3.29 and in more detail in Table 3.15.

190

Fig. 3.29 High level overview of contour method analysis

191

Table 3.15 Contour method workflow

Step Name Comments Equipment /

SW

1 Contour Cut Sample cut transversely to the

direction of forces in interest

EDM

2 Surface profilometry

(Contour scan)

Resulting cross-sectional surfaces

scanned for deviation from perfect

straight surface

Nanofocus

laser scanner

3 Surface profile data

processing

Removing areas of no interest and

excess noise

Matlab

4 Automated data cleaning Generates also outline of the

sample

Matlab

5 Surface averaging 2 surfaces averaged into one Matlab

6 Noise removal Averaged surface cleaned Matlab

7 3D model extrusion Outline extruded into a 3D model. Abacus

8 3D model meshing and

surface data import

Linear quadratic elements Abacus

9 Calculation of forces

required to flatten the

surface

Using the model generated and

boundary conditions

Abacus

10 Contour map generation Data filtering plays important role Abacus or

Matlab

The contour cutting for the samples used was performed by EDM. In order to

prevent plasticity artefacts due to cutting the samples were restrained very carefully.

A special jig, Fig. 3.30, was used to restrain the sample during cutting to reduce

plasticity.

192

Fig. 3.30 Contour cutting jig lower half. Top beams removed to reveal the

individually adjustable pads which ensure restrainment avoiding inducing

external stresses. The cut was performed between the beams along the line

drawn in red.

The cutting strategy developed involved die sinking of two holes through the sample,

see cross-section in Fig. 3.31, between which the actual contour cut was performed

using wire EDM. The remaining ligaments supported the sample during the process

to reduce cutting to further reduce the plasticity artefacts. The ligaments were cut last

to separate the halves. The EDM process took altogether for approximately 5 hours.

193

Fig. 3.31 Contour cut sample, cross-sectional view. Contour cut made first, with

ligaments in place to support the sample.

After the contour cut the sample was restrained in a jig, Fig. 3.32, for surface

profilometry. The sample was levelled horizontally using a test dial indicator and a

measurement table. This was to reduce digital data processing and make the

scanning process more reliable. The laser scanner used was NanoFocus µScan. The

resulting raw data was converted from a proprietary format to a .DAT ASCII data

file using the proprietary software.

Fig. 3.32 Sample of SA5 mounted to a jig undergoing the contour scanning

Pilot hole

Ligament

Contour cut

40 mm

194

The profilometry was conducted to both resulting halves of the sample. It provided

two 3D point cloud files; see an example in Fig. 3.33. The data would go through a

series of processing steps to provide a filtered average of the two samples, Fig. 3.34.

Decisions between under and overfitting were made to avoid excess noise while not

removing vital data. The averaged point cloud was also used to generate a silhouette

profile spline with seeds for mesh generation later in the process, Fig. 3.35.

Fig. 3.33 Profilometry raw data example, (µm)

Fig. 3.34 Averaged data with some residual noise, not in scale

195

Fig. 3.35 Profile outline silhouette spline with seeds (circles) for mesh generation

The surface outline spline generated was used to generate a 3D model of the sample,

Fig. 3.36. The outline silhouette was used to extrude a model of the real sample. The

model was divided into three materials using the fusion lines to separate different

material properties. The processed profile data was used with a mesh of linear

quadratic elements to create a solid model. Solid lines with seeds at regular intervals

were added to the model at three locations. This was to generate data points with

regular intervals for line plotting the results at certain predetermined depths.

196

Fig. 3.36 Finite element mesh (left) and the 3D extrusion (single material sample

shown)

Different knot spacings were used for the contour maps, see Fig. 3.37 and Fig. 3.38.

It was noted that different data processing yielded very dissimilar results. The main

observation was the sensitivity of extreme values to the data smoothening.

Fig. 3.37 Contour method residual stress map plotted using dense knot spacing

Fig. 3.38 Contour method residual stress map plotted using less dense knot spacing

197

3.10.5 Tensile Testing

The tensile testing was conducted according to ASTM E8/E8M – 16a [182]. The

tester used was an Instron 4507 two ball screw tension and compression testing

machine. The standard presents 450 mm and 200 mm overall length samples as

standard and specifies a 100 mm coupon as sub size. However, it was decided to use

a smaller size due to limitations by the material available to conserve material for

further tests. The test samples used, Fig. 3.39, were 75 mm long. The samples were

extracted by EDM and tested in as-machined condition.

Fig. 3.39 Tensile test sample ASTM E8/E8M. Longitudinal and composite similar.

Two types of tensile tests were made, longitudinal and transverse (composite). The

longitudinal tests were conducted in both base materials, heat affected zones and the

weld metal. The transverse tests were made across the entire weld. These tests were

made in triplicate. All tensile tests were recorded for DIC processing.

The correlation between tensile strength σ and hardness HV can be described as per

( 3.4 ) [73].

𝐻𝑉 ≈σY

3

( 3.4 )

198

3.10.6 Digital Image Correlation

Digital image correlation was used to analyse the strain localisation in the tensile test

samples. The equipment used was LaVision Strainmaster Portable. The raw data was

processed using DaVis 8.3.1 software to produce strain maps. Further analysis was

performed by Strainmaster DIC 1.0.0 software. The test setup Fig. 3.40 consisted of

a stereo camera arrangement where both cameras were equipped with separate

flashlights.

Fig. 3.40 DIC setup for tensile testing. 1) cameras, 2) flashlights and 3) tensile

test sample.

The samples were prepared by spraying a contrast speckle pattern on one side. This

was done by spraying a speckle pattern with alternating contrasting colours. Black

and white paints were used. The resulting pattern is displayed as Fig. 3.41.

1 1 2 2 3

199

Fig. 3.41 Tensile test samples used for DIC by spraying a speckle pattern

3.10.7 Charpy-V Impact Toughness Analysis

A series of Charpy-V impact tests were conducted to the samples. The tests were

made according to ASTM E23 [199] and ISO 148-1:2016 [178]. Standard size

coupons were used, Fig. 3.42.

Fig. 3.42 Charpy-V impact toughness test coupon as per ASTM E23 and ISO 148-1

Impact toughness tests were concentrated on the HAZ of the ferritic steel. This area

is the most prone to ductility issues in these types of welds. The United States

Nuclear Regulatory Commission sets the minimum impact energy of 102J for ferritic

materials of nuclear reactors [159]. A comparison set was provided by tests in the

ferritic base material. The tests were performed at several different temperatures to

study the ductile-brittle transition temperature. Results were plotted according to BS

Standards Publication PD ISO/TS 7705:2017 [179]. The temperatures used ranged

between 150°C and -120°C.

201

4 Thick Section Narrow Gap Laser Welding of

316L Stainless Steel

4.1 Introduction

At the beginning of this study, a series of stainless steel similar metal NGLW

experiments were made. These thick section welds up to 40 mm were made using

200 mm long samples. They were named Stage I. Their purpose was to serve as the

basis for the development of the welding protocol, equipment, groove geometry,

monitoring and analysis methods for the future dissimilar welding trials. These

experiments were also found valuable for familiarising the staff to the welding

process and the principal analysis methods of the weldments.

The main points of interest and challenges encountered by previous researchers were

mainly in groove design and tolerances [36, 44], and the gas shielding efficiency [17,

18, 20]. Several studies were found considering challenges in NGLW. Jokinen 2004

[17] reported oxidation increasing porosity. Yu et al. 2013 [18] concluded high-

speed shielding gas jets distorting the weld pool shape. Zhang et al. 2011 [20]

discussed using different shielding gases but did not find issues in their 3 pass welds

while using a simple single leading gas jet. They also found two types of lack of

fusion, interpass and sidewall, which were also examined in this study.

During the welding, the process was monitored using different methods. The

preheating and interpass temperatures were recorded. The welding distortion was

observed by mechanical measurements and laser 3D scanning. Each pass was

visually checked for wetting, undercut and fusion quality. As the programme

202

progressed, more sophisticated methods were developed for monitoring the process,

such as laser illumination imaging.

The resulting welds were subjected to visual examination to assess the quality [161].

The successful welds were analysed for through-thickness quality using the industry-

standard ASME IX radiography [176] before proceeding to cross-sectioning and

further microstructural and hardness analysis.

The equipment development, the process characteristics and effect of parameters are

discussed. The welds were cross-sectioned and basic properties like fusion, porosity

and cracking were analysed. These tests are listed in Table 4.1.

Table 4.1 Stage I of thick section similar metal narrow gap laser welds in 316L

ID. Thick.

[mm]

Groove Restraint No. of

clamps

Gas shield Notes

TV1 12 parallel Buttress

clamps

1 pair 2

Stationary

nozzles +

1 moving

nozzle

First

thick

section

NGLW


clamps

2 pairs 2

Stationary

nozzles


clamps

2 pairs 2

Stationary

nozzles

TV4 40 parallel Strongback - 2

Stationary

nozzles

First

40 mm

NGLW

TV5 40 4° Lenskes

clamps and

table

3 pairs 1

Stationary

nozzle

First

test

using

LII

TV6 40 4° Lenskes

clamps and

table

3 pairs 1

Stationary

nozzle

LII used

203

4.1.1 Expectations

The Stage I stainless steel welding was expected to give insight to the narrow gap

welding process. In order to understand the future dissimilar welds, these welds were

to be used as a benchmark. The similar metal welding removed the complexity of

dilution issues and the multitude of microstructures at different metallurgical zones

when using dissimilar metals. Therefore, by studying both similar and dissimilar

welds the phenomena related to the metallurgical properties were expected to be

distinguishable from the properties of the actual process itself.

The objectives were to develop robust welding equipment and procedures for

monitoring, measuring and documenting the process and start analysing the welded

samples using basic analysis equipment.

The 316L material for this stage NGLW was chosen as in literature the most work

had been executed on stainless steels. The non-phase transforming behaviour of

austenitic steels [98] was considered an important factor simplifying the

microstructural behaviour of the welds. Challenges in ferritic steel welding like

issues with cracking encountered by Feng 2016 [45] reinforced the choice, as the aim

was to simplify the effect materials have on the process at this stage as much as

possible.

The welding distortions were a major point of interest. The distortion measurements

were expected to allow detailed development of the welding groove geometry. The

development of the restraint system was central to this work as the equipment was to

be developed from scratch. The mechanisms leading to the contraction of the

welding groove were to be understood. The butterfly distortion angle and linear

204

contraction of the welding groove were to be measured and the correlation between

the grove contraction and the butterfly angle was to be analysed.

The gas shielding efficiency was to be analysed in order to avoid welding flaws. The

performance of the shielding was evaluated visually. The effect of oxidation on the

behaviour of the weld pool was to be monitored using LII. The resulting welds were

to be analysed for flaws by ASME IX radiography and macrography.

4.2 Materials

The materials used for the Stage I similar metal welding were 316L austenitic

stainless steel and ER 316L filler wire. They are described in detail in chapter 3.2.1.

4.3 Experimental Work and Results

The welding setup used in Stage I was based on previous experiences found in the

literature [17, 20, 35, 45] and fast prototyping with thinner materials. The

development of the experimental setup is described in detail in Chapter 3. During the

Stage I a welding protocol was developed. A procedure for accomplishing a 40 mm

multi-pass weld in 8 hours was accomplished. A systematic approach for recording

data from the experiments was developed and constantly improved. Laser

illumination imaging was implemented for the welds. Basic analysis for assessing

the quality of the weldments was executed.

The development of the equipment was one of the key topics at this stage of the

programme. A major part of the work was to reach the positional accuracy for the

wire feeder nozzle matching literature [44, 156]. There were issues like porosity,

reported by Jokinen 2004 [17] and fusion issues Zhang et a.l.2011 [20] found. The

205

gas shielding was improved in numerous ways to avoid them. The work was

complicated by very limited access to observe the welding process. Numerous

different nozzles for the filler wire and the shielding gas were studied and several

monitoring arrangements were developed and investigated.

4.3.1 Experiments 1-3

For experiments TV1, TV2 and TV3 the aim was mainly equipment development.

This included in principal development of the various nozzles required and

integrating the wire feeder control to the robotic system. Also, the gas shielding

arrangement was created and tested during these welds. The thicknesses of these

welds varied from 12 mm to 30 mm to save material and not to exceed the maximum

loading of the clamping system.

These first tests were made using parallel sidewall welding grooves. The filler wire

was 316L at 1.2 mm in diameter. The shielding gas was injected using two 15 mm

nozzles placed at the ends of the welds. The welding parameters for these tests are

displayed in Table 4.2.

Table 4.2 Welding parameters TV1 to TV3

ID. Groove

width

[mm]

Thick.

[mm]

Passes

welded

Power [kW] Focal

offset

[mm]

Speed

[m/min]

Wire

speed

[m/min]

TV1 4 12 4 3 40 0.72 3, 5, 5, 5

TV2 5 30 11 11, 7, ->

8.14

99 0.72 5

TV3 5 30 14 8.14 91 0.48 4 x 5, ->7

As a welding pass cools down, it shrinks and causes welding distortion. The

distortion pulls the sidewalls of the welding preparation together. In NGLW the main

issue resulting from this is that the welding nozzles will not eventually fit in. In

206

conduction mode NGLW the change in welding groove width would also require

adjustment of the laser spot size. To develop a groove design capable of anticipating

the welding distortion in order to keep the bottom of the groove width similar for

each pass, the welding distortions were monitored from the beginning of this study.

The contraction of the welding groove was monitored with measurements of the

welding groove width before the welding and the un-welded ends of the groove after

the welding, see Table 4.3.

Table 4.3 Groove contraction measurements TV1 to TV3

ID. Gap pre

welding

[mm]

Gap post

welding

[mm]

Contraction

[mm]

Contraction

[%]

Calculated

angle

[mm]

Clamps

TV1 4.0 2.9 1.1 28 6.3 Machining

TV2 5.0 damaged* - - - Buttress

TV3 5.5 3.6 1.9 35 3.6 Buttress

* The groove ends were damaged during welding preventing measurement

The welds were cross-sectioned, etched with 2% Nital and macrographed to see the

fusion quality. Generally, the fusion was good with no porosity, cracking or

inclusions, see Fig. 4.1. However, misalignment of the welding apparatus gave poor

results for the capping, see Fig. 4.2.

207

Fig. 4.1 Cross-section of TV1 showing a successful weld with no fusion issues and

good penetration to the backing plate

Fig. 4.2 Cross-section of TV2 showing asymmetry and poor capping due to

misalignment of the sample.

4.3.2 Experiment 4

The experiment TV4 was the first full-thickness weld. At 40 mm thickness, it was

not possible to use the buttress clamping available at the time, but a strongback

208

restraint was welded to the underside of the samples. This weld was the last to use a

parallel sidewall welding preparation and the welding distortion measurements were

used for developing the V-groove capable of accommodating the distortions.

The parameters for the 21 pass TV4 are presented as Table 4.4. The parameters were

kept constant during the entire weld, except for the passes no. 2 and 15 which were

re-melting passes and therefore performed without filler feed. The pass no. 2 was

needed to seal some apparent poor wetting of the first pass. The pass no. 15 was an

attempt to re-melt a large droplet off the sidewall, which was caused by interruption

of the previous pass due to unexpected movement of the sample. The droplet was

melted partially but prevented filling the groove fully as it intersected with the

leading filler wire nozzle. The weld was filled only partially for passes 16 to 21.

Table 4.4 Welding parameters TV4, all passes.

ID. Groove

width

[mm]

Thick.

[mm]

Passes

welded

Power [kW] Focal

offset

[mm]

Speed

[m/min]

Wire

speed

[m/min]

TV4 4.8 40 21 8.14 91 0.72 7 or 0

The distortions in TV4 were recorded manually at each end of the welding sample,

see Fig. 4.3. The gaps in data are due to the gap not being measured for re-melt

passes at the time.

209

Fig. 4.3 Average groove width per pass, measured directly between the top corners

of the welding groove, TV4.

4.3.3 Experiment 5

The parameters for TV5 were as with TV4, refer to Table 4.4. This weld was used

mainly for testing the laser illumination imaging setup in real NGLW conditions. All

welding passes were recorded and the resulting imagery found a valuable tool for

analysing irregularities in the process. A sample image is presented as

TV5 was also cross-sectioned and etched using oxalic acid. The macrographs

indicate a regular weld and good fusion, see Fig. 4.4. There were no cracks,

inclusions or pores.

210

Fig. 4.4 Macrographical cross-section showing no issues with fusion, cracks or

porosity TV5.

The laser illumination imaging was used to observe and record the welding process.

A frame of the video is shown as Fig. 4.5.

Fig. 4.5 A LII video frame, showing wire, nozzle and the welding pool. Note the

stable welding process. Stainless steel similar metal weld TV5

211

4.3.4 Experiment 6

The aim for the last Stage I weld was to investigate the groove contraction behaviour

of a 4° V-angle chosen based on contraction measurements on previous welds. The

groove contraction was measured manually from several points along the weld, both

directly from the groove edges and from punch mark points on the top surface. Also,

3D scanning was improved and the results analysed further. The laser illumination

imaging was also implemented for these welds.

Weld butterfly distortion of TV6 was measured using a hand-held 3D scanner. The

scanning was performed after each pass to examine the development of the

distortion. The butterfly distortion angle was measured from 200 cross-sections in

the 3D models. Fig. 4.6 describes the results for weld TV6.

0 25 50 75 100 125 150 175 200-0.50

-0.25

0.00

0.25

0.50

0.75

1.00

1.25

1.50

1.75

2.00

2.25

2.50

2.75

Released

Pass 15

Pass 14

Pass 13

Pass 12

Pass 11

Pass 10

Pass 9

Pass 8

Pass 7

Pass 6

Pass 5

Pass 4

Pass 3

Pass 2

Pass 1

Initial

Dis

pla

cem

ent A

ngle

(D

egre

es)

Length Along Sample (mm)

Fig. 4.6 The generation of butterfly distortion, 3D scan data analysis, test TV6

In Fig. 4.6 the concentration of the distortion towards the mid-length of the weld is

unmistakably displayed. This was later seen also as the welding groove getting

212

similarly slightly narrower at mid-length, causing the middle of the weld to fill

sooner than the ends.

From the distortion measurement results of experiment TV6, one may conclude that

a variable angle welding groove might be required. In the tests, it was found that the

welding groove width at the bottom of each pass stayed constant enough for reliable

welding when a 5° V-groove straight wall welding groove was used.

The groove top gap contraction was measured manually at 5 points; two direct

measurements and three measurements from punch marks offset ca. 20 mm from the

weld centreline. The distance between punch marks was then converted to a

calculated gap width. The following Table 4.5 and Fig. 4.7 describe the results. It is

noteworthy that the direct measurements match the calculated groove width very

well until the end of the weld. This divergence near the top was found to be due to

local yielding of the edge of the welding groove continuing even when the

completed weld had sufficient thickness and rigidity to counter larger scale butterfly

distortion.

213

Table 4.5 Groove contraction measurements, test TV6

Pass Point 1 Point 2 Point 3 GAP 1 GAP 2

0 38.40 37.82 38.78 8.20 8.26

1 38.11 37.24 38.14 7.85 7.69

2 37.80 36.95 37.80 7.55 7.33

3 37.50 36.73 37.50 7.26 7.04

4 37.26 36.39 37.44 7.00 6.81

5 37.05 36.15 37.32 6.81 6.60

6 36.72 36.06 36.93 6.49 6.37

7 36.48 35.68 36.60 6.13 5.95

8 36.18 35.40 36.34 5.87 5.70

9 36.02 35.14 36.29 5.63 5.47

10 35.23 34.82 36.18 5.51 5.36

11 35.86 34.75 36.12 5.40 5.27

12 35.77 34.84 36.20 5.31 5.14

13 35.81 34.91 36.20 5.20 5.08

14 35.80 34.86 36.10 5.08 4.92

15 35.70 35.01 36.03 -* -

16 35.73 34.78 36.21 - -

* accurate measurement not possible

0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 204

5

6

7

8

Ga

p w

idth

(m

m)

Pass no:

Direct Measurements

Indirect Measurements

Fig. 4.7 Groove top contraction development of a 5° V-groove 40 mm thick 316L

similar metal weld. See the minimum staying over 5 mm. Direct

measurement was not applicable for last 2 passes due to melting of the

groove top edges, test TV6.

214

The results in Fig. 4.6 and Fig. 4.7 show an apparent discrepancy. A rapid increase

of butterfly distortion was seen in the 3D scanning during the passes from 5 to 9. In

the groove top gap contraction measurements, the results were near linear. This can

be explained by the overall distortion happening in different stages. The first stage

was found to be a nearly pure linear translation, as the first passes melted almost the

entire depth of the weld, which did generate contraction but was not able to cause a

change in the butterfly angle. At the second stage, passes 5 – 9, the existing

solidified weld was acting as a fulcrum and as the present pass contracted, a

considerable increase of butterfly angle was generated. When the proportion of the

already existing weld reached a sufficient level, the generation of butterfly distortion

was essentially stopped by this added restraint. From that on, the local yielding of

the welding groove was found to be the major contributor to the groove gap

contraction This can also be seen as a deviation of the direct and indirect

measurements near the end of the weld in Fig. 4.7, where the indirect measurements

from a distance away from the weld indicate no change, but the direct measurement

shows ongoing contraction.

The welding test TV6 was also used for trialling pre-heating by the processing laser

to demonstrate the versatility of laser systems [40]. Four passes were made with no

wire feed, using the parameters for pre-heating. The aim was to create a local pre-

heating effect. The welding parameters used are displayed in Table 4.6. The

parameters were kept constant from the first to the last pass where the welding speed

was slightly lowered to ensure good penetration due to signs of imperfect fusion on

the previous pass and to give a slightly higher deposition rate to fill the pass

adequately.

215

Table 4.6 The preheating and welding parameters for similar metal welding, test

TV6.

weld Preheat unit note

Power 8.14 0.5 kW

Welding speed 0.012 0.004 m/s

0.72 0.24 m/min Last pass 0.10

Focal offset 91 91 mm

Measured Spot Size 5.1 5.1 mm

Power density 6719 2448 W/cm2

Head angle 2 2 °

Wire speed 7 0 m/min

0.117 0 m/s

The preheat and interpass temperatures were monitored before each pass. The

temperatures were measured at the top of the samples at the middle of the weld,

20 mm from the weld centreline. Also, the thickness of the bead accumulated was

measured using a calliper. Results are displayed in Table 4.7.

216

Table 4.7 Temperature and thickness development of TV6 similar metal 316L weld

Pass no: Sample temp.

(°C)

Thickness

(mm)

pre-weld 20.2

preheat 1 21.5 0

preheat 2 22.6 0

preheat 3 23.9 0

preheat 4 25.0 0

1 25.0 3.3

2 34.4 6.6

3 44.4 9.9

4 45.9 13.0

5 46.0 14.0

6 48.1 17.0

7 47.6 20.4

8 27.2 23.6

9 36.6 26.0

10 39.3 28.6

11 42.9 31.2

12 43.1 33.4

13 39.2 35.4

14 38.8 37.3

15 40.2 39.5

16 41.2 n.a.

4.4 Discussion

4.4.1 Welding Restraint, Welding Distortions and Groove Design

Detailed results of welding distortions in NGLW were found scarce in the literature.

Yu et a.l. 2013 [18] discuss the benefits of NGLW as a low heat input process over

arc welding. Elmesalamy et a.l. 2014 [11] conclude that butterfly distortion of a

20 mm NGLW is approximately 1/3 of similar GTAW. They also discuss the crucial

role of restraint conditions on the distortion.

During the welding of TV1, it became evident that a 4 mm wide groove would not

be large enough to allow sufficiently rigid nozzles to be used. To reach the sufficient

217

positional accuracy reported by other researchers [44, 156], a nominally 5 mm wide

groove was used from that on.

The calculated groove angle distortion changed considerably when the restraint

changed from the relatively thin section TV1 restrained with lightweight machining

clamps to the almost 3 times as thick TV2 and TV3 restrained by the buttress

clamping system. This was a clear indication of the effect of restraint conditions on

the welding groove design, as investigated by Elmesalamy et a.l. 2014 [11]. The

conclusion of this was that the groove design would have to be tailored to suit the

restraint conditions on a case by case basis, which in the case of this study required

the decision on the final restraint system to be made before a more detailed design

was possible.

A 5 mm parallel sidewall groove design was used for the first full-thickness 40 mm

stainless steel welding test TV4. As predicted, the welding groove was found to

contract during the welding. It was found wide enough, although the groove width

decreased severely as the weld filled up; the tip of the wire was able to fit in as the

actual nozzle moved above the groove.

The originally 5 mm wide groove top decreased to 3.5 mm for the last passes. The

strongback restraint used was found able to constrain the overall butterfly distortion

to 2° over the width of the sample, but not capable of preventing the local yielding of

the material in proximity of the groove. Therefore the total contraction of the groove

was calculated to require a 4° V-groove to maintain each individual pass 5 mm wide.

The 3D scanning of the weld TV6, pass by pass, was found to yield interesting

insights into the development of welding distortion. However, it was found to be

somewhat unreliable and prohibitively time-consuming. Each scan took several

218

minutes during which time other measurements had to be suspended. This allowed

additional time for the weld to cool, which caused a decrease in the interpass

temperature compared to welds without 3D scanning. For stainless steel welds, this

was known not to cause issues. For dissimilar welding, the extra delay between

passes would have caused an unacceptable drop in the interpass temperature and the

3D scanning was not implemented during these welds.

It was also noted that although a remelt pass was found able to heal irregularities and

slight lack of sidewall fusion it contracted the groove as much as a regular filling

pass.

Results of the weld TV6 confirmed that the 42 kN welding clamps offer similar

restraint as the cumbersome strongback setup and that the 4° V-groove was suitable

for this application also. The slight drop below 5 mm suggested by Fig. 4.7 at the last

2 passes was not an issue since all nozzles were already clear of the groove. Instead,

it was found beneficial for fusion at the top of the weld, allowing the standard 5 mm

wide spot to generate a weld pool wide enough to cover the edges of the base

materials reliably.

4.4.2 Gas Shielding

In literature issues with the oxidation in NGLW were not often reported [17, 18, 20].

This was concluded to be mainly due to the much lower number of passes generally

used by the researchers due to the smaller material thicknesses or two-sided welding.

In Stage I welding, the gas shielding using the stationary nozzles was considered

satisfactory for welding the one-sided 40 mm thick 316L stainless steel similar metal

weld. The welds needed 15 to 19 passes to fill the gap, including re-melts. Some

219

oxidation was observed to accumulate on top of the weld bead, but as the sidewall

and interpass fusion were mostly acceptable it was not considered an issue. The

oxide layer was able to fully float to the surface of the weld pool during the process

leaving no inclusions.

4.4.3 Wire Feeding

The TecArc F4 wire feeder unit, described in 3.7.2, was found not to give

satisfactorily repeatable feed rate, leading to unwanted irregularities and some flaws

in the welds. Wire conduit bending was found to affect the alignment of the actual

wire tip coming out of the nozzle. The interaction point of the wire and melt pool

would move by up to 3 mm during the length of the weld.

An optimal wire feed angle was found to be approximately 30° and an optimal free

wire length 16 - 20 mm. A shorter free wire would cause overheating of the nozzle

and longer would jeopardise positional accuracy. The ultimate limiting factor for

wire feed rate was found to be buckling when unmelted wire penetrated to the

bottom of the weld pool, as happened during the TV4. Because the wire speed was

considerably higher than the welding speed, the wire would buckle and slip under

compression, ejecting molten metal out of the melt pool causing spatter. The spatter

droplets were big enough to prevent the welding nozzle from entering the groove for

the subsequent pass.

The filler wire nozzle used was made from a copper tube. This was found to get

easily damaged by two main mechanisms. The more common issue was the relative

softness of the copper. The nozzle tip was easily bent by mishandling or even the

forces encountered during the welding process. Another very severe flaw was the

220

low heat capacity of the nozzle, which led to instantaneous melting of the nozzle

when exposed to the random reflections of the laser from the melt pool.

In the similar metal tests, it was agreed with the literature [44, 156] that the wire

alignment in the NGLW groove had to be precisely controlled. Horizontal wire

alignment was found to have little effect on the weld pool. It was concluded that the

turbulence in the weld pool was able to counteract the effect of the asymmetry. The

vertical alignment was found to be more critical. In cases where the wire overshot

the weld pool, the wire was heated by direct laser power. This would lead to

overheating of the wire and cause evaporation and oxidation. If the alignment was

too low, the wire would run erratically due to scraping on the previous bead. When

the vertical wire alignment was correct, the heat was able to conduct from the melt

pool to the wire.

The 3D scanning of TV6 showed the welding groove contracting faster at the middle

due to the welding distortions, thus generating a convex profile for the weld. As the

robot trajectory was linear, the convex weld would cause the actual vertical

alignment to change during the pass. This would have a slight influence on the spot

size, which was found insignificant as very long focal length optics were used. The

effect on the wire feeder vertical alignment was found more severe. If the wire

height was too low, the wire would hit the welding groove bottom prematurely and

risk getting stuck to the possible irregularities in the previous bead and buckling. If

the wire overshot the weld pool, the deposition would be compromised as described

earlier. The wire height would have to be set between these limits. In the worst cases,

the bead height would change approximately 2 mm. A compromise in wire

alignment was found in most cases and the effects of this issue were limited to the

top layers of the weld.

221

4.4.4 Laser Illumination Imaging

Laser illumination imaging was used for analysing the welds. An example of the

versatility of the information the system was able to provide is illustrated in Fig. 4.8

and Fig. 4.9. The observations are listed in the captions.

Fig. 4.8 Laser Illumination Image of pass 1 in a similar metal NGLW. 1) Uneven

wetting of side walls due to too small spot size, oxidation of weld bead,

2) Wire nozzle melting due to reflected laser beam, 3) Wire contamination

by particles of nozzle and vibration (in video material), 4) Smoke

generation due to excess laser power

Fig. 4.9 12th

pass of a similar metal NGLW. 1) Good smooth fusion, 2) Gap

contracting excessively, laser beam heats the corners of the groove

prematurely, 3) Wire alignment off-centre and 4) Accumulated oxides

floating on top of the melt pool

1 3

2

4

4

1

2

2

3 4

222

Comparison of Fig. 4.8 and Fig. 4.9 shows the improvement achieved during a single

welding test. The laser power was optimised to prevent evaporation of the molten

pool. The wire nozzle was moved further away from the weld pool to prevent

damage. The accumulation of an oxide layer can also be seen.

4.4.5 Pre-Heating and Interpass Temperature Control

Pre-heating by using the laser beam was examined during weld TV6. Running four

passes at a power density just below melting the bottom of the groove was able to

raise the temperature of the samples only about 3.5°C. This was far less than desired

for the future dissimilar welds. Also, it was noted that the oxidation of small surface

irregularities during the pre-heating had a possibly negative effect on quality. Pre-

heating by using the laser was found ineffective and slow and a more efficient

method of applying heating blankets was chosen for the consequent welds.

Control of the interpass temperature was examined. To keep the interpass

temperature at a certain range, the welding passes were repeated with regular

intervals. In practice, this meant that only limited time was available for taking the

photographs and making the measurements. Often the following pass would have to

be welded as soon as possible after the procedures.

The preliminary stainless steel welds were designed for maximum speed and

deposition rate. To reach the maximum heat input whilst still staying in the

conduction mode the limit power density could not be exceeded. Bearing in mind the

main variables affecting the mode in laser welding are spot size and laser power, the

narrow gap geometry limited the maximum spot size leaving the laser power to be

the only effective parameter to vary.

223

4.5 Conclusions

The Stage I stainless steel NGLW was conducted mainly for developing the welding

procedure, groove design and equipment. The weld geometries and practices

developed were later used as a starting point for dissimilar welding. A significant

product was the development of the welding procedure, documentation and

familiarisation to the practice of welding a multi-pass weld which would take hours

to complete.

The welding process was found robust and not sensitive to welding parameters, for

example, laser power varying from 3 kW to 11 kW were experimented with no

significant change in weld quality as observed from metallographic and visual

inspection, including the undercut, solidification cracking, lack of fusion defects,

oxidation, and porosity.

Sufficient precision in weld and laser alignment and spot size control were

fundamental to achieving good fusion. The requirements were found less stringent

than in laser keyhole welding. For wire alignment, the transverse tolerance Wy was

found very forgiving, as little effect was observed. The wire height Wz and

longitudinal alignment Wx were found more critical as over- or undershooting of the

weld pool caused imperfections. A tolerance of ±2 mm was found to produce

satisfactory results.

Accuracy in the region of ±0.5 mm was required for the equipment positioning and

welding groove preparation. Laser spot size matching the welding groove width was

found to give the best results, too small leading to lack of fusion, too wide to damage

to sidewalls. See Fig. 4.8 and Fig. 4.9. The laser spot size tolerance was ±0.2 mm. A

continuously adjustable spot size using online visual monitoring would’ve alleviated

224

the tolerance related issues. High level of reliability of the equipment and

repeatability of the welding process had to be attained even for experimental work,

as a single momentary imperfection was able to invalidate a substantial amount of

work.

Control of the restraint was found to be crucial as it had direct influence on the

butterfly distortion generated and therefore the welding groove design. Local

yielding of the welding groove contracted the groove in addition to a larger scale

butterfly distortion. Local yielding was responsible for 50 % of the total contraction

and the butterfly for the other 50 %.

The analysis of the welding distortion using parallel side wall groove indicated a

need for a 4° V-angle. A single bevel design was considered accurate enough for

accommodating the distortions and maintaining a constant gap at the bottom of the

groove for each pass and maintaining access to the groove.

Flawless gas shielding was obtained for single passes. However, oxidation was

found accumulating during the course of multi-pass welding.

In comparison to conventional arc welding, the NGLW was found require more

precise geometries in the weld preparation and equipment. The access limitations to

the welding cell placed higher than usual importance to the electronic process

monitoring equipment. LII equipment was found highly beneficial as it was also

capable of recording the welds. It was concluded that a more flexible welding robot,

as usually used for narrow gap welding applications, capable of online parameter

adjustments during welding, would have made work with the developmental setup

and parameters more efficient and faster.

225


Dissimilar Metals S275 and AISI 316L

5.1 Introduction

As the second step in this study, 40 mm thick section dissimilar metal NGLW was

investigated. This stage was performed using 316L stainless and S275 mild steels as

base materials and Inconel Alloy FM52 as the filler. This stage was named Stage II.

The Stage II dissimilar NGLW using 316L and S275 was aimed to verify the

suitability of the welding process and equipment developed during Stage I for the

dissimilar materials. The development of the analysis methods for the welds was

given more emphasis. Hardness mapping, microhardness testing, dilution and

microstructural analysis were developed.

The purpose of this stage was to develop the welding process and analysis methods

using cheaper LAS compared to the more expensive nuclear grade SA508 LAS. This

stage was found necessary due to the very different characteristics of the FM52

nickel filler metal and the environment in the dissimilar metal weld pool compared

with the Stage I stainless steel welding. The viscose weld pool of FM52 [129], the

tendency to DDC [64, 84, 127], liquation [98] and solidification cracking [82, 128]

and the tendency to oxidation [89, 128] were to be studied. The lack of bonding and

lack of fusion reported by McCracken 2010 [128] were to be investigated.

The welds were monitored in a similar way to the previous stage, using mechanical

and optical means, but omitting the 3D scanning. This was mainly due to the time

needed for the scan allowing the interpass temperatures to drop excessively. The

226

quality of the welds was assessed by visual inspection, online monitoring by an LII

system and mechanical measurements of the distortions. The preheat and interpass

temperatures were logged.

To assess the integrity of the welds, they were subjected to the industry-standard

ASME IX radiographical inspection [176]. Cross-sections of the welds were

hardness mapped and microhardness analysed applying BS EN ISO 6507-1:2005

[180] and ASTM E384-17 [181] standards. These results were used together with

microscopy to perform microstructural analysis. Etching of the samples was

conducted according to Chapter 3.10.2. To assess the possible issues with the

compositional changes, the dilution of the weld metal was analysed using EDX.

Stage II samples were named DS1-5, see Table 5.1.

227

Table 5.1 The Stage II of dissimilar thick section narrow gap laser welds using

S275, Alloy 52 and 316L

DS1 DS2 DS3 DS4 DS5

Aim Parameter

development:

preheat

Parameter

development:

preheat

Lack of

fusion

study

Weld pool

management

Undercut

management

Thickness 40 mm 40 mm 40 mm 40 mm 40 mm

Groove 4° 4° 4° 5° 5°

Restraint Lenskes

clamps and

table

Lenskes

clamps and

table

Lenskes

clamps

and table

Lenskes

clamps and

table

Lenskes

clamps and

table

No of

clamps

3 pairs 3 pairs 3 pairs 3 pairs 3 pairs

Gas shield 1 Stationary

nozzle

1 Stationary

nozzle

1

Stationary

nozzle

3 blade

nozzle

3 blade

nozzle

Video Laser

Illumination

Imaging

Laser

Illumination

Imaging

Laser

Illuminati

on

Imaging

Coaxial

camera

Coaxial

camera

Notes Final version

wire feeder

Final version

wire feeder

5.1.1 Expectations

The Stage II dissimilar metal welding using a non-pedigree ferritic steel was

performed to test and compare the effect of the FM52 filler metal with the similar

metal welds performed at Stage I and the findings in the literature. The main result

expected was the ASME IX integrity of the NGLW, considering the various known

challenges found in literature and during the similar metal welding.

The questions to be researched in this stage include:

What is the optimal welding groove angle for these materials?

228

Is NGLW capable of accomplishing ASME IX radiographical requirements?

Will the susceptibility to oxidation of FM52 cause issues? Are there ways to

reduce the possible problems?

How does the material mix affect fusion; how the different materials will

affect weld bead shape and wetting?

How does the weld pool turbulence distribute the elements, is the dilution

homogenous across the weld or between the passes?

Is there DDC, solidification or liquation cracking in FM52?

Could pre-heating be omitted for this type of material?

What is the hardness distribution, are there possibly brittle areas of high

hardness?

How does the multi-pass tempering change the microstructure and hardness

of the materials?

5.2 Materials and Experimental Methods

The first dissimilar metal welds using S275 as the ferritic steel were conducted in

two stages. The main differences were in the welding head setup. First three tests

were made using stationary gas shielding nozzles. For the last two welds, the setup

was developed by adding an improved wire feeding nozzle and a new moving

shielding gas nozzle capable of injecting the shielding gas deep into the welding

groove. Another significant difference between the two stages was the change of

groove angle from 4° to 5°.

229

5.2.1 Materials

Materials used for the Stage II welds were AISI 316L and S275 base materials and

Inconel Alloy 52 filler wire. The weld thickness was 40 mm and the filler wire

diameter 0.9 mm. The materials are described in detail in chapter 3.2.2.

5.2.2 Weld Setup and Parameters

Welding parameters for S275 dissimilar metal welding were based on earlier tests

made with similar metal stainless steel welds and an unpublished technical report

considering dissimilar NG-GTAW. This vast report demonstrated that a thick section

unbuttered weld using materials similar to the ones used in this study can be

completed to the industry-standard specifications. Modifications were made to

achieve similar heat cycle and cooling conditions while using a laser as the heat

source to produce crack-free welds.

The 40 mm thick weld samples were 200 mm long, providing an effective weld

length of 160 mm. The weld preparation was a 4° V-groove with straight sidewalls

and a 5 mm wide root, refer to Fig. 3.1. The samples of 316L and S275 were

100 x 200 x 40 mm in size, providing a weld length of approximately 170 mm. The

sample size was mainly limited by the capacity of the restraint. The weld preparation

design was based on the results of the earlier similar metal welds. It was wire-EDM

machined to have straight sidewalls at a shallow V-angle. 4° and 5° V-angles were

used. The width of the groove was set at 5.0 mm at the root, increasing linearly to the

top to compensate for the welding distortions during the process, see Fig. 5.1. The

aim was to keep the actual width of the welding groove bottom constant for each

pass and allow free access for the wire-feeding nozzle. The welding process was

started from a backing plate with a normal filling parameter set without a separate

230

autogenous root pass. The purpose of the backing plate was to support the first bead

of Ni-alloy weld metal and allow generation of a uniform chemical composition

throughout the thickness. The backing plate was tack welded to the samples and

therefore it also served as a supplementary restraint preventing the samples from

sliding closer together because of the tensile forces during the first passes. The

backing plate was made of 316L and was 6 x 40 mm in cross-section. In real

applications, the backing structure would be machined off prior to use.

Fig. 5.1 Initial 4° welding groove design for S275 dissimilar metal welds

The initial weld setup for S275 dissimilar metal welds is presented in Fig. 5.2. This

setup was used for welds DS1, DS2 and DS3. The laser head, wire feed nozzle,

welding restraint and laser illumination imaging system are displayed.

231

Fig. 5.2 Experimental setup for a DMW aligned for the first pass. 1) Laser head, 2)

coaxial wire feeding and shielding gas nozzle, 3) stationary shielding gas

nozzle, 4) welding clamps and 5) high-speed camera lens. Laser

illumination laser is just outside the picture on right.

5.3 Experimental Work

The work on Stage II built on Stage I results and literature. Relevant work on NGLW

was reported by Jokinen 2004 [17], Elmesalamy 2013 [35], Feng et a.l. 2016 [45],

Zhang et a.l. 2011 [20] and Yu et a.l. 2013 [18] but no published reports on

dissimilar metal NGLW were found.

5.3.1 Experiments 1-3

The first dissimilar NGLW DS1 was welded using a 4° V-groove and 169°C pre-

heat. The welding was started with one pass with high power to ensure melting of

the sidewalls via the separate backing plate. For the rest of the weld, the parameters

were kept constant throughout the process, see Table 5.2.

1

2

3 4

5

232

Table 5.2 Welding parameters DS1.

ID. Groove

width

[mm]

Thick.

[mm]

Passes

welded

Power [kW] Focal

offset

[mm]

Speed

[m/min]

Wire

speed

[m/min]

DS1 5 40 18 8.14 to 2.6 91 0.72 7 or 0

DS1 was hardness analysed by hardness mapping. Altogether 5 maps were produced

along with line scans. The weld was also etched for ferritic steel HAZ to analyse the

microstructure of the HAZ and the areas of high hardness.

Fig. 5.3 DS1 dual etched with Nital and oxalic acid. The hardness mapping

indentations also visible

For DS2 the welding parameters were revised. The unnecessarily high power for the

first pass was omitted and to improve fusion the power was increased for the filling

passes. No pre-heat was applied. The parameters were kept constant throughout the

weld as per Table 5.3. DS2 was used for low force Vickers hardness test

development.

233

Table 5.3 Welding parameters DS2.

ID. Groove

width

[mm]

Thick.

[mm]

Passes

welded

Power [kW] Focal

offset

[mm]

Speed

[m/min]

Wire

speed

[m/min]

DS2 5 40 17 3.5 91 0.72 7 or 0

The experimental setup for DS3 was improved by introducing an improved gas

shielding setup instead of the stationary nozzles. Also, the wire feeder unit was

changed to a more robust one. The welding parameters were revised, the first pass

was welded as previously, but the laser power was decreased to 3 kW for the

following passes. The laser power was reduced further due to signs of overheating at

pass no. 11. The two capping passes were welded with increased power and spot

size.

DS2 required one less pass than DS1 to fill the weld. This was due to the interrupted

wire feed during one pass in DS1. These early welds were hampered with equipment

issues. DS3 required most passes, as bad wetting forced to remelt the 15th

pass and

two re-melt passes were applied to smoothen the capping.

As the previous groove design was developed using stainless steel similar metal

welding, among the interests in DMW was the groove contraction behaviour. This

was monitored during the first tests DS1, DS2 and DS3. The resulting distortions

were averaged to give the data more statistical relevance.

The welds were graded visually for oxidation, symmetry and fusion quality after the

welding. The results are displayed in Table 5.4. This was done mainly to describe the

weld quality in a simple table to justify the choice of samples for X-ray analysis. The

234

weld DS3 was chosen for radiographical analysis and subjected to ASME IX

acceptance.

Table 5.4 Visual examination of welds DS1 to DS3

DS1 DS2 DS3

Oxidation unacceptable unacceptable acceptable

Symmetry acceptable unacceptable acceptable

Fusion unacceptable acceptable very good

5.3.2 Experiments 4 and 5

The severity of the oxidation issues forced the development of a trailing gas nozzle

inserted into the welding groove, see Fig. 5.4. The distortion measurements

suggested a wider welding groove should be applied. Therefore the welding tests

DS4 and DS5 were conducted in a 5° V-groove using the tri-blade shielding gas

nozzle. The laser power was varied, pass by pass, according to the conditions in the

welding groove. Two new phenomena were monitored, pass by pass basis, the

undercutting of the stainless steel due to excess laser power and poor wetting of the

sidewalls due to insufficient power. The analysis was done visually, as the

measurement of the phenomena was found impossible. The undercut was graded

from 0 to 50, 0 being no undercut, 30 meaning unacceptable. The fusion quality was

graded 50 to 0, 50 being flawless and 30 meaning unacceptable.

235

Fig. 5.4 The nozzle with added trailing nozzles as used for DS4.

The parameters used in test DS4 are displayed in Table 5.5.

236

Table 5.5 Welding parameters and notes DS4

Pass Temp

316L

Temp

S275

Depth Power Undercut Fusion

1 64.2 66.6 0.0 3500 0 50

2 65.4 67.3 2.2 3000 0 40

3 79.3 79.1 4.3 3500 10 50

4 86.7 86.5 6.6 3250 10 50

5 88.0 87.6 9.2 3250 10 50

6 68.9 66.3 12.0 3250 0 50

7 86.9 88.1 14.5 3250 2 50

8 86.5 88.1 15.2 3250 2 30

9 91.1 92.0 17.9 3500 15 50

10 90.6 90.0 19.5 3250 10 40

11 81.0 84.3 22.0 3500 15 50

12 85.2 86.1 24.5 3250 10 50

13 82.5 81.8 26.3 3250 10 50

14 82.0 82.3 28.4 3000 10 10

15 86.0 84.6 30.0 3250 * 40

16 70.0 70.5 ** 3250 50

17 66.0 64.0 3400 50

18 68.3 66.7 3500 10

19 76.3 73.5 4000 50

20 78.5 78.9 4000 10

21 70.0 70.0 3500 30

22 47.0 47.0 3000

* Sidewalls damaged, no estimate applicable

** Surface distorted, laser depth measurement not applicable

The weld DS4 suffered from undercutting, which widened the groove and made the

shape irregular near the surface, see Fig. 5.5. Closer inspection of the imagery

reveals lack of fusion at the area of an undercut, illustrated as Fig. 5.6. The issue of

the widened groove was rectified by two parallel capping passes to ensure fusion.

This corrected most of the fusion issues superficially, Fig. 5.7. The weld was X-ray

analysed to analyse the sub-surface integrity of the weld as the ability of penetration

to fix the fusion was unknown.

237

Fig. 5.5 19th

pass of DS4, note undercutting at 130-170 mm.

Fig. 5.6 Close-up of undercut at 19th

pass of DS4, note lack of fusion to the

stainless steel (arrows)

Fig. 5.7 Finished weld DS4. Note the mainly good fusion due to two parallel

capping passes.

238

The weld DS5 was free of undercutting and equipment issues. The parameters used

in test DS5 are displayed as Table 5.6. Photograph of the first pass is presented as

Fig. 5.8. Some oxidation occurred, but it was removed using a wire brush to good

effect as the beads were smooth with good wetting to the sidewalls. Fig. 5.9

illustrates the weakest quality fusion encountered during this weld. Although it was

known from previous work that increasing the power would alleviate the poor

fusion, it was decided not to change the power as the previous experience suggested

the generation of undercut if the power was increased. Fig. 5.10 displays the fusion

and oxidation quality of the finished weld.

Table 5.6 Welding parameters and notes DS5

Pass Temp

316L

Temp

S275

Depth Power Undercut Fusion

1 71.7 79.1 0.0 3500 0 50

2 81.0 78.1 2.1 3250 0 40

3 87.0 87.9 4.4 3250 0 50

4 87.0 88.2 6.4 3250 0 50

5 85.0 85.2 8.4 3250 0 50

6 86.0 87.0 10.5 3250 0 50

7 87.6 89.0 12.9 3250 0 50

8 87.1 88.0 14.9 3250 0 20

9 80.3 79.1 16.8 3250 0 40

10 88.1 87.6 18.4 3250 0 40

11 80.5 80.0 20.4 3250 0 50

12 73.3 73.0 23.0 3250 0 50

13 82.8 85.7 25.2 3250 0 30

14 85.3 85.7 26.4 3250 0 25

15 83.0 84.1 28.4 3250 0 20

16 84.1 85.2 30.4 3350 0 20

17 85.4 85.8 32.4 3500 0 30

18 82.0 82.5 33.4 3500 0 35

19 79.9 77.0 35.1 3500 0 35

20 68.0 68.0 36.4 3600 0 45

239

Fig. 5.8 First pass of DS5. Note smooth, symmetrical wetting, oxidation and no

irregularities

Fig. 5.9 Close-up of the 8th

pass of DS5. Note unacceptable irregular oxidation and

wetting.

240

Fig. 5.10 The finished weld DS5 prior to wire brushing.

The finished welds DS4 and DS5 were analysed visually to justify the radiographical

analysis. The results are displayed as Table 5.7. Both welds were subjected to the

radiographical analysis.

Table 5.7 Visual examination of welds DS4 and DS5

DS4 DS5

Oxidation acceptable very good

Symmetry acceptable very good

Fusion acceptable very good

5.4 Analysis of Process Characteristics

5.4.1 Weld Pool Behaviour Analysis

Due to the laser safety regulations regarding class 4 lasers, access to the welding cell

was prohibited during the welding process. This made conventional real-time

observation of the weld pool behaviour impossible. This limitation was overcome by

obtaining a Cavitar Cavilux laser illumination system and a Phantom Miro 4 high-

241

speed camera, which were introduced before proceeding to the dissimilar metal

welding. Later the Cavilux laser had to be replaced with an Oxford Laser light

source. The illumination laser power was 50 % lower, but was sufficient for low

frame rate imaging, as the laser pulse length was increased correspondingly. The

biggest drawback of this system was that the whole laser source had to be pointed to

the weld since it had no fibre optical beam delivery like the Cavilux.

In an LII system, a diverging laser light source illuminates the target area with

monochromatic light. This light is then received using a camera equipped with an

optical narrow band-pass filter. The arrangement filters out all other light sources

like the welding process glare, providing an unparalleled image of the process, see

Fig. 5.11.

Fig. 5.11 Laser illumination imaging image, note the absence of glare. DS2, second

pass.

In Fig. 5.11 several observations can be made, such as:

- The wire is aligned ca. 0.5 mm to the right of the centreline of the weld

- Weld pool wets the sidewalls with no irregularities

- Oxidation floating on the weld pool is visible as a matte layer

242

Laser illumination imaging material was recorded in video format, which allowed

much easier observations than still photography. Phenomena like vibrations and

irregular behaviour of the filler wire, wetting of the sidewalls and behaviour of the

oxide slag were easily observed. Some of the passes were recorded at high frame

rates in an attempt to see the melt pool turbulence more clearly.

One of the main outcomes of the laser illumination monitoring was the effect of the

oxides to fusion. Especially in video format lack of fusion was possible to detect

with good accuracy. An example of a local inadequate wetting leading to lack of

fusion is shown as Fig. 5.12.

Fig. 5.12 Formation of lack of fusion / inadequate wetting (arrow) and oxidation of

the weld bead viewed by laser illumination imaging

When the need for a more efficient shielding gas system became evident, the laser

illumination system had to be dropped. This was mainly due to the illumination light

being blocked by the trailing shielding gas nozzles. From this point on imaging was

achieved by ambient light coaxial camera optics built in the Precitech laser welding

head, using the Miro IV camera, see Fig. 5.13.

243

Fig. 5.13 Coaxial camera ambient light image. DS4, first pass.

The loss of image quality in using the coaxial system was considerable. The main

reasons for this were the inability of the system to focus at the exceptional focal

offsets used and the glare from the glowing vapours over the melt pool. Also, the

field of view was narrower. The coaxial camera as provided by the laser welding

head manufacturer was designed for focusing at the focal spot of the process laser. In

the welds in this study, the offset was at minimum +91 mm. This was circumvented

greatly by using the smallest aperture available in the system. The high contrast due

to the glare made it practically impossible to observe only a few phenomena,

however, the following conclusions can be made using Fig. 5.13:

- The welding wire is aligned ca. 1 mm to the left of the centreline of the weld

- The weld pool wets the sidewall symmetrically, as the glare from the sidewalls is

symmetrical and continuous across the weld pool

- There is a small piece of debris on the bottom of the weld groove just below the

filler wire interaction point

- The sidewalls are not damaged by the previous passes

The issues in wetting and fusion behaviour of the weld were seen as asymmetry of

the white overexposed reflections. The possible sidewall damage was visible at the

244

correctly exposed portions of the side walls. Also, with the moving video material, at

the beginning of the weld, where the weld pool was still developing, the alignment

of the laser beam, spot size and wetting symmetry was easy to see. Some black dead

pixels can be seen in the images, they should be ignored.

5.4.2 Symmetry of the Weld Bead

Surface tension gradients dictate the direction of flow in a weld pool via the so-

called Marangoni effect. Even small changes in the composition may cause a

significant effect [200]. In DMW this may lead to asymmetric flow and heat transfer

in the weld pool. To investigate the magnitude of the phenomenon in NGLW, along

with the LII videos, the welds were photographed to reveal the symmetry of the bead

surface. No asymmetry of the flow was seen in the LII video material where the

floating oxides acted as indicators. The weld surface photography, see Fig. 5.14,

showed symmetrical distribution of the weld metal agreeing with the LII results.

245

Fig. 5.14 DS4 second pass pictured from the end of the welding groove. Note the

level surface of the bead with no inclination to either metal and the sharp

corners instead of smooth wetting.

5.4.3 Welding Distortion Analysis

The direct and indirect measurements were found to be concordant and consistent as

shown in Fig. 5.15.

246

0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 204

5

6

7

8

Ga

p W

idth

(m

m)

Pass no:

direct measurements

indirect measurements

Fig. 5.15 Welding groove contraction measured directly and calculated from

indirect measurements. Test DS2. Error bars represent min and max

values measured.

The distortion data showed that the groove contraction behaviour matched the

predictions. In S275 dissimilar metal welding, three distinct stages of contraction

were recognised. The first stage was rapid shrinkage when there was little or no

material connecting the halves other than the fully melted weld pool, which would

cause shrinkage partly as horizontal translation. This can be seen at passes 1 to 3 in

Fig. 5.15. At the second stage, there was a fulcrum of solid metal just under the

shrinking and contracting melt pool. At this stage, the contraction slowed down

significantly, as shown in Fig. 5.15 passes 4 - 13. As the solid connection between

the halves thickened, the contraction slowed down even further. Towards the end of

the weld, the groove contraction occurred only very slowly, only due to local

yielding of the materials. This can also be observed as sinking of the top surfaces of

the samples adjacent to the weld.

247

During the comparison of the different weld distortion measurements, it was found

that the results from the indirect measurement were similar whether the indent

distance from the centreline was 20 mm or 70 mm. The difference between these

measurements in DS2 was on average 0.06 mm, which was similar to the accuracy

achieved during the manual measuring process. This finding led to time savings

during the welding as the indirect measurements were later conducted using only the

20 mm offset method.

Averaging the data from all DS welding tests using 4° V-groove proved that the gap

contracts below the desired 5 mm after 12 passes, see Fig. 5.16. This was while the

groove remained approximately 15 mm deep. Therefore there were difficulties with

fitting the nozzle setup in the groove and the laser spot had to be adjusted smaller to

avoid damaging the groove top corners. As a consequence, the groove angle was

increased to 5° to accommodate the distortion.

0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 204

5

6

7

8

Ga

p W

idth

(m

m)

Pass no:

direct measurements


Fig. 5.16 Welding groove top width of DS series of welds using 4° V-groove. Note

the groove contracting to below 5 mm. Average of three tests, DS1-DS3.

Error bars represent min and max values measured.

248

The contraction behaviour of the 5° V-groove is displayed in Fig. 5.17. It can be

seen, that the contraction effectively stops at 14 passes. The resulting width was

5.4 mm. While this is larger than the desired 5.0 mm it was found not to cause

issues. Increasing the spot size was not required, as the wetting and fusion were

found satisfactory. This was attributed to the heat conduction conditions changing

towards the top of the weld. Deep in the groove, the heat had almost 360° of solid

steel to dissipate the heat. Near the surface this was closing in to 180° of steel and

180° air. Therefore it was concluded that there was a beneficial increase in the weld

pool temperature as the weld approached the surface.

0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 204

5

6

7

8

Ga

p W

idth

(m

m)

Pass no:

direct measurements


Fig. 5.17 Welding groove top width of DS series of welds using 5° V-groove. Note

the groove width staying above 5 mm until the end. Average of tests DS4

and DS5. Error bars represent min and max values measured.

The distortion data was used to design a groove shape which would allow the use of

constant spot size. The benefits were obvious, as the laser head with the shielding

gas and filler wire feeding apparatuses could be moved up as a single unit after each

pass without needing to offset the laser and the auxiliaries separately which is

249

required should the spot size be changed. At the specific material thickness,

materials, restraint and welding parameters used, it was found that a V-angle of 5°

with straight sidewalls was sufficient to counter the effect of the distortions. Some

small adjustment of the spot size was implemented during the welding, as the

welding parameters used and the distortion behaviour were not entirely constant.

5.5 Analysis of Weldments

A set of analysis was conducted on selected samples. First, the welds were

characterised using an industry-standard NDT approach. Then further analysis was

conducted to examine the properties of the welds such as hardness, grain structure

and residual stresses.

5.5.1 Radiographical Analysis of Welding Flaws and ASME IX

Acceptance

Of the DS series of dissimilar welds DS3, DS4 and DS5 were selected to be analysed

by radiography. Results were subjected to acceptance by ASME IX. None of the

samples was acceptable over the whole length, see Table 5.8. Nonetheless,

acceptable quality was achieved for certain parts of the welds. DS3 had issues only

at the very end of the weld which could be discarded as a run-off area and the first

50 % of the DS4 was found flawless. The full radiographical reports are presented as

Appendix II.

250

Table 5.8 Summary of radiographical acceptance reports of DS series of welds

ID ASME IX Location 1-2 Location 2-3

DS3 Unacceptable “Lack of sidewall fusion in approximately the last 40 mm of

the weld”

DS4 Unacceptable Intermittent lack of sidewall

fusion

No visible defects

DS5 Unacceptable Intermittent lack of sidewall

fusion

Intermittent lack of sidewall

fusion

The flaws leading to rejection were entirely fusion related. The X-ray reports note

intermittent lack of fusion as the defect. LoF related issues in multi-pass HWLW had

also been reported by Todo et. al. in 2015 [54]. It was concluded that some of the

fusion issues were equipment or operator related. The alignment of the laser and

filler wire was found to be critical for good wetting and fusion. Excess oxidation was

found to lead to very poor wetting of the side walls which led to lack of fusion to the

sidewalls and the occasional case of lack of fusion/oxide inclusion between the

passes in the weld. The equipment was later improved by developing a precision

adjustable wire feed nozzle and the tri-blade gas shielding system.

5.5.2 Hardness Mapping

The mechanical analysis of the DS series of welds was started by performing

hardness mapping. Performing the hardness analysis first was to give valuable data

for further investigations. Several hardness maps were made. Different Vickers

hardness scales were used and different grid spacing experimented for displaying the

hardness profile. First hardness maps were plotted using HV 1.0 scale, see Fig. 5.18.

The resulting indentations at the softest areas detected were 123 µm in diagonal

corresponding to 124 HV1. This limited the grid spacing and edge distance to a level

which was not determined to be accurate enough. To minimise the issue, the

251

following tests were made using HV0.3. The testing was conducted in as welded

condition.

Fig. 5.18 Hardness map of DS1. 69 (HV1) S275 left, FM52 middle, 316L right.

Fusion lines in red. A main observation is the strain hardening of the

austenitic materials.

Maximum hardness in HV1 was 250 HV. An area of high hardness was found in the

work-hardened stainless steel. Individual hot spots at the S275 fusion line were

indicating a narrow area of high hardness in the HAZ. The hardness of the weld

metal was at maximum 240 HV.

The hardness map was found lacking in width in the stainless steel work hardening

area. Therefore the test was repeated using a uniform grid of test points, see Fig.

5.19. This grid was spread evenly over the whole sample to show the hardness

distribution over the whole sample. This test was performed using the more flexible

HV 0.3 indentation force. The large area of interest required the use of a less dense

252

grid to limit the number of test points to reduce the test time. The same sample was

used re-grinded and polished.

Fig. 5.19 Hardness map of DS1 69 (HV0.3. Approximate fusion lines in red. Note

the narrow area of high hardness at the S275 FL.

Comparisons of the maps showed, that the apparent area of each point measurement

was much larger in the latter test. This was due to the sparser grid spacing allocating

a larger surface area for each point, making single points look more pronounced.

In the hardness map in HV 0.3, the hardening of the stainless steel was fully covered.

The analysis revealed a smooth hardness transition to the base material. The area of

highest hardness was found at strain hardened 316L on the right and the lowest in

S275 base material. The profile of the area of SS strain hardening was explained by

the welding distortions being the largest near the top of the weld, including the very

surface of the sample.

253

In the ferritic steel, the hardness of the HAZ was low, as expected from a low

hardenability mild steel. Single hard points were detected near the S275 fusion line,

having the maximum value of 245 HV1. The seemingly random spots near the mild

steel fusion line were concluded to require closer inspection.

In both HV1 and HV0.3, the top pass of the fusion zone was found to be softer than

the underlying passes. The reason for this difference was determined to be strain

hardening of the austenitic Ni alloy weld metal, similar to the strain hardening of the

stainless steel. The last bead did not undergo similar compressive stress as the

previous passes were subjected to. The same phenomenon was also observed in

316L in the HAZ caused by the last weld bead.

5.5.3 Hardness Line Scan Measurements

The relatively sparse grid used for the mapping was found not to give a detailed

enough picture of the area of main interest, the HAZ closest to the S275 fusion line.

It was also concluded that the plotting had generated some artefacts. These made the

map less accurate than desirable, making further conclusions difficult.

Dense line plots were made to establish the hardness variation across the weld in

more detail and to gain a more detailed understanding of the hardness profile. Fig.

5.20 shows hardness variation across the weld with a 0.1 mm interval. As the maps

predicted, the line indicates that a local hard region was generated in S275 HAZ near

the fusion line. This hardened area was found to be 1.5 mm wide. A much softer

area, starting just before the fusion line and spreading 0.5 mm to the weld metal was

found, see Fig. 5.20. The hardening of the S275 HAZ was a result of the bainitic

microstructure generated in fine grain zone (FGZ), see Fig. 5.24.

254

-20 -15 -10 -5 0 5 10 15 20 25

125

150

175

200

225

250

275

300

FM52 SA508

Ha

rdness (

HV

1)

Distance from S275 fusion line (mm)

S275

Fig. 5.20 Hardness line plot across the weld, second to top pass. Note the narrow

area of high hardness in S275. Fusion lines marked on the graph. Sample

DS1 69.

Solid-state carbon diffusion to the Cr rich fusion zone (decarburisation) during

cooling was considered to contribute to the premature drop of hardness. The

apparent drop of hardness before the fusion line was determined to be possibly an

artefact due to hardness measurement indents partially reaching the adjacent softer

weld metal. It was also recognised, that although this test did not fill the

requirements for indent spacing set by standards, the measured hardness values were

in good accordance with the results from the first map. The measurement was

concluded to give a good overview of the hardness changes over the weld.

Further hardness investigation was made for sample DS2 116 in order to investigate

the S275 fusion line in fine detail and assess the multipass tempering effect. To

comply with the indent spacing minimums, these tests were made using a grid of

indents to space the measurements further apart. The locations of the indents on the

sample are presented in Fig. 5.21. In the figure, the first grid of indentations is

already measured and the indents are visible.

255

Fig. 5.21 Hardness test grid and locations. Note top bead being considerably wider

than previous passes to ensure smooth capping. Sample DS2 116.

The actual distance of each indent perpendicular to the fusion line was measured and

the results were plotted. Points from the less dense measurements at the same height

were added to extend the graphs to BM and FZ. The four topmost line plots are

presented in Fig. 5.22. Abnormally high hardness values were occasionally measured

in the FZ near the fusion line. These were due to the automatically distributed test

point striking an unmixed ferrite stringer. They were removed from the line plot.

256

-2000 -1500 -1000 -500 0 500 1000140

160

180

200

220

240

260

280

300

320

FM52

Hard

ne

ss (

HV

0.3

)

Distance from FL (µm)

TOP

-1

-2

-3

S275

Fig. 5.22 Hardness of S275 HAZ at different states of multipass tempering.

Multipass tempering can easily be noticed from the results. This tempering has a

strong beneficial reductive effect on hardness. The maximum hardness of the S275

HAZ reduced when approaching passes which have gone through tempering cycles.

By comparing results from different weld beads it can be seen that three heat cycles

temper almost all the hardening caused by the welding.

A drop in hardness in the S275 HAZ, 50 µm from the fusion line can be seen in a

close up of the FL in Fig. 5.23. The HV 0.3 indents in the HAZ near the FL were

approximately 50 µm in diagonal. This would suggest horizontal accuracy of a

similar scale. Decarburization of the S275 HAZ was a suggested reason. The drop in

the S275 HAZ hardness before the FL can be explained by the indents being affected

by the proximity of the softer FZ.

257

-500 0 500140

160

180

200

220

240

260

280

300

320

Hard

ne

ss (

HV

0.3

)

Distance from FL (µm)

TOP

-1

-2

-3

S275 FM52

Fig. 5.23 Hardness of the S275 HAZ. Close up near the fusion line.

5.5.4 Microstructural Analysis

Microstructural analysis was conducted on the cross-sectional samples of the weld.

The main areas of interest were the austenitic to ferritic/bainitic fusion line of the

S275 and overall grain size.

A cross-section of the whole S275 HAZ is displayed as Fig. 5.24. The width of the

HAZ was approximately 2.5 mm. The different grain structures at different areas of

the welds were examined in further detail.

Fig. 5.24 S275 HAZ of a tempered filling pass, BM-tempered-ICHAZ-FGHAZ-

CGHAZ-WM

In the S275 base material, Fig. 5.25, the grain size was found to be 15 – 20 µm. In

the grain growth zone of the HAZ, the grain size was practically unchanged, Fig.

5.26. The HAZ grain size was found to have increased to 30 µm near the top where

258

higher heat inputs were used owing to longer time spent over the austenisation

temperature, Fig. 5.27. Together with the results from the hardness analysis, the

grain structure was derived bainitic, as the hardness was not high enough for

martensite in 0.25 C steel.

Fig. 5.25 Unaffected S275 base material

Fig. 5.26 Grain structure at a tempered filling pass

Bainite, tempered

Decarburised zone

Weld metal

Pearlite, dark

Ferrite, light

259

Fig. 5.27 Untempered CGHAZ at the top pass

A lighter area of suspected decarburisation was seen in some instances, like Fig.

5.26. The tempering effect of multi-pass process was found to round the grain in

S275, but not to lead to grain growth as the maximum temperature was not high

enough for that to occur.

In the Alloy 52 weld metal, the grain size was 10 µm. At the larger heat input top,

the grain size varied between 10 – 20 µm. The whole WM was austenitic in

structure, apart from a 20 µm partially mixed zone, Fig. 5.28.

Fig. 5.28 DS1 weld metal microstructures

The multi-pass tempering of the weld was found to lead to complex thermal histories

which were aggravated by the slight changes of parameters like heat input

necessitated by controlling of the melt pool and wetting. The overall conclusion of

the microstructural analysis was that there were no detrimental structures being

100 µm

Columnar- cellular microstructures

Equiaxial growth

Banding

Bainite

Weld metal

Fusion line

260

developed in the weld. The grain size was small in every area of the weld. The multi-

pass tempering effect was seen as banding in the HAZ microstructure, Fig. 5.29.

Fig. 5.29 S275 HAZ banding due to multi-pass tempering

5.5.5 EDX Dilution Analysis

The weld DS1 was analysed for weld metal dilution using EDX. Scans were made at

three depths in the weld. One at the root pass, one at the 3rd

pass and one at the top

pass.

The root pass line scan is presented in Fig. 5.30. The left-hand side of the scan

presents the S275 mild steel fusion line, while the 316L fusion line is just outside of

the field of view. The first observation was that the weld metal composition was very

consistent. There were no distinct signs of banding or grouping of elements. In the

261

area where the composition changes from S275 to weld metal a gradual change was

seen. The gradient in Fe and Ni content to the average level of the weld metal was

300 µm wide. The Cr content reached the average weld metal level in 150 µm.

0 500 1000 1500 2000 2500 3000 3500 4000 4500

0

20

40

60

80

100

wt.

%

X (µm)

Cr Wt%

Fe Wt%

Ni Wt%

Fig. 5.30 EDX line plot of the weld metal and fusion lines. Note the rapid transition

from ferritic BM to WM composition

The weld metal was found to have been diluted noticeably by the base materials, see

Table 5.9. The effect of the location of the weld in the thickness was found to have

no significance, as the average content of the three main alloying elements were

found to be stable through the thickness of the weld.

Table 5.9 Dilution of weld DS1

Fe Wt% Cr Wt% Ni Wt%

Alloy 52, analysis 9 29 60

Root pass 24 33 43

3rd

pass 23 34 41

Top pass 24 33 43

262

The S275 fusion line at top bead is displayed as Fig. 5.31. The change of

composition to the average weld metal can be seen happening in a shorter distance

than in the root pass. The width of this transitional zone is 30 µm. Ni is again the

element reaching the weld metal composition the quickest. This shorter transitional

zone can be attributed to the larger heat input at the top pass contributing to a more

turbulent weld pool. Unmixed ferrite can be seen 180 µm from the FL as a drop in

Ni and an increase in Fe content.

200 400 600 800 1000 12000

20

40

60

80

100

wt.

%

X (µm)

Fe Wt%

Ni Wt%

Cr Wt%

Fig. 5.31 EDX line plot for top bead at the S275 fusion line. The transition in the

composition is prompt.

In Fig. 5.32 of the 316L fusion line a more gradual change in chemical composition

was seen. The transition of Fe spreads over a distance of 200 µm. The Cr has no

distinct transition zone, as the change in content spreads to a wide area and the

change in content is relatively small. In this case, also, the Ni content has the steepest

gradient which is only 10 – 20 µm wide.

263

2600 2800 3000 3200 34000

20

40

60

80

100

wt.

%

X (µm)

Cr Wt %

Fe Wt %

Ni Wt %

Fig. 5.32 EDX line plot for top bead at the 316L fusion line.

5.6 Discussion

Weld quality acceptable by ASME IX examination was produced, but not achieved

along an entire length of a weld. The acceptance was performed radiographically and

inspected for linear indications, generally for incomplete fusion or slag inclusions,

and rounded indications, usually porosity. Very few issues with rounded indications

were detected. The main challenge was control of lack of fusion defects. See

radiographical acceptance examination reports for welds DS3, DS4 and DS5, in

Appendix II. Two mechanisms leading to poor fusion were found; a floating layer of

accumulated oxides affecting wetting due to inadequate gas shielding and the oxide

layer not separating from the previous pass due to lack of penetration.

The fusion issues in the dissimilar metal welds were caused partly by the limited

laser power that could be applied to the weld pool. The laser power, the directly

proportional heat input and hence the penetration was limited by the appearance of

undercut in the stainless steel sidewall when using excessively high power. This was

detrimental to the symmetry of the weld and fusion, see Fig. 5.5. and Fig. 5.6.

264

The weldability of FM52 filler metal was known to be challenging [80, 89, 128]. The

melt pool behaved sluggishly, evident from the angle of wetting in Fig. 5.14, and the

spot size control had to be precise to ensure acceptable wetting and fusion. The

higher viscosity of FM52 compared to S275 was known to cause a Marangoni flow

away from the S275 base material, making the wetting more difficult. Although

observations of the floating oxide particles on the LII video material did not indicate

a visible effect, this may have had an effect at the immediate vicinity of the material

interface.

The filler metal was also found to oxidise strongly, as suggested by literature [89,

128], the de-oxidising alloying elements, oxide floaters, forming a crust of Ti and Al

oxides on top of the weld pool. This was found to be an issue especially as in the

multi-pass welding the oxides accumulated over the passes. Improved shielding gas

nozzles were able to alleviate the problem, but the issue persisted.

One major challenge in NGLW of S275 to 316L using FM52 was the precision

required from the process. The spot size and alignment of the weld had to be correct

within approximately 0.2 mm. If the spot size was too small, the wetting of the

sidewalls was compromised, see Fig. 5.14. If it was too large, the sidewalls were

damaged by reflections, see conclusions in Chapter 5.4.1, page 240. Filler wire

alignment was found critical in the vertical direction, overshooting being more

detrimental. The overshooting filler wire would be heated in direct interaction with

the laser beam, leading to droplet formation at the tip of the wire. This droplet would

oxidise keenly and drop into the melt pool causing irregular bead surface, the effects

shown in Fig. 5.9. The required precision was approximately ±1 mm, tolerances

being similar to Jokinen’s findings [17]. The horizontal alignment was not found to

265

be critical, as the metal was distributed evenly to the area being heated by the laser

spot, despite the wire interaction point deviations.

The limiting factor for the maximum thickness was the accumulation of oxides, see

Fig. 5.9. Reliable welding of the passes deep in the welding groove was

accomplished, but the accumulation of the oxide layer made welding after 12 passes

unreliable. The thicker the oxide layer accumulated, the poorer the wetting was. The

thick oxide layer prevented the curvature of the surface and made the sidewall

interface angular, similar to Fig. 5.14. Mechanical removal of the oxide layer

alleviated the problem to a degree.

The dilution characteristics across the weld were very uniform. This was found

different from the conventional multi-pass GTAW results of Chung et a.l. 2011 [85],

where the dilution gradients were very smooth instead of sharp changes. The dilution

was considerable compared to the centreline of a wide conventional weld. The

increase in Cr was beneficial but the decrease in Ni was undesirable. Overall this

reduces the SCC resistance of the weld metal, but may be considered a minor issue

as the highest susceptibility to SCC is at the weld interface [85] and not in the weld

metal.

The results can be summarised as follows:

Producing ASME IX acceptable welds was found challenging due to

formation of lack of fusion defects

o Much of the lack of fusion was attributed to the sluggish, viscous melt

pool of FM52

The gas shielding of FM52 was known to be challenging. In the multi-pass

welds, the oxides were found to accumulate to a detrimental effect.

266

o Accumulation of oxides was found to be a major contributor to lack

of fusion to sidewalls and between the welding passes

o Considerable development of the gas shielding system was required

to achieve the repeatability required in multi-pass welding

Asymmetrical undercutting of the stainless steel was observed

o The undercut distorted the weld geometry, especially at the top passes

o The undercut contributed to lack of fusion defects

The welding distortions were high enough to warrant an increase of the V-

groove angle to 5°

The two different indirect groove contraction measurements investigated

produced similar results

Re-melting a pass was found able to repair small fusion defects and abnormal

bead shape to some extent

A very narrow area of high hardness was found in the S275 mild steel HAZ.

This is typical for low-carbon mild steels, where hardenability is low and

only the most intense heat cycle just adjacent to the fusion line can cause

hardening.

Microstructural analysis confirmed an area of fine-grain bainite in S275

HAZ, see Fig. 5.26. This was formed by the heat cycle of the welding and the

heat treatment caused by the multi-pass welding

The fusion zone grain structure was fully austenitic as expected for a high-

nickel alloy, as per the Schaeffler diagram, Fig. 2.26. It was found free of

solidification and liquation cracking, which often are issues with highly

restrained welds [98].

Multi-pass tempering effect produced strain hardening in the austenitic weld

metal and stainless steel HAZ, see Fig. 5.19.

267

5.7 Conclusions

The dissimilar metal welding using EN S275 mild steel and AISI 316L austenitic

stainless steel with Inconel filler metal FM52 was found to pose considerably more

difficult challenges compared to the previous similar metal NGLW. Although nearly

all the questions set at the beginning were answered, the ductility dip cracking

investigation was found to require more detailed work. The following conclusions

were reached:

From the direct and indirect welding groove distortion measurement results

of the DMW tests with a 4° V-groove it was seen that the welding distortions

contracted the gap below the nominal 5 mm. In conclusion, a 5° V-groove

was introduced for the following tests. The increased contraction was

concluded to be caused by the increased number of passes and heat input

compared to the Stage I welds.

ASME IX radiographical acceptance was not achieved for an entire length of

a weld due to lack of fusion type defects. It was concluded that adding a run-

in and run-out sections to the weld would help to solve the issue, as most

issues were concentrated at the ends of the welds. However, lengths of

acceptable joint were accomplished.

The accumulation of oxides on the FM52 was recognised increasing the

propensity for lack of fusion as the welds progressed. Controlling the gas

shielding environment and removal of the oxide layer between passes were

found crucial to the success of the welds. A sufficient purity shielding gas

environment was established after introducing large low-velocity gas nozzles

as close as possible to the weld pool inside the groove. Using small diameter

high-velocity gas jets was found to cause turbulence, which was found to

cause contamination of the gas by the ambient air.

268

Weld symmetry was investigated visually between passes. No significant

asymmetry was found. However wetting issues occurred mainly at the 316L

fusion boundary.

No solidification or liquation cracks were found in the FM52 fusion zone.

The dilution of base materials to the FM52 weld metal analysed by EDX was

noticeable. This was not considered an issue, as literature [85] states that in a

DMW the weakest SCC properties are found at the fusion interface in any

case, and depend on the complex local microstructure. The elemental

composition of the entire fusion zone was found evenly distributed. This was

confirmed both across the weld and vertically through the thickness. This

indicates good mixing of the weld pool which leads to uniform properties.

Hardness mapping suggested phase transformation hardening at a very

narrow region next to the FL in the S275 mild steel HAZ. This was

confirmed by detailed microhardness analysis and found 1.5 mm wide. The

resulting grain structure was found bainitic.

In hardness mapping strain hardening was observed in the FM52weld metal

and 316L stainless steel. The last welding pass and the resulting HAZ were

found softer than the underlying areas due to lack of additional heat cycles

and the consequent strain hardening.

Hardness mapping showed uniform hardness through the thickness of the

weld. This supports the uniformity of the properties.

The geometrical positioning tolerances were found similar to the Stage I

similar metal welding, the dissimilar materials having no notable effect.

Very careful control of the laser power was required, on the contrary to the

Stage I findings. By comparing the records of welding parameters and cross-

section macrographs of the welds, the limits of laser power were established.

The laser power lower limit was set by the interpass lack of fusion. The upper

limit was set by the generation of asymmetrical undercutting of the stainless

steel.

269

A beneficial multi-pass tempering effect was found to produce small grain

size in the S275 HAZ. This suggests good mechanical properties and impact

toughness. There was little change in microstructure after 3 heat cycles. From

there on the weld was uniform.

270

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271


Dissimilar Metals SA508 Gr3 Cl2 and 316L

6.1 Introduction

Stage III and the main aim of this study was dissimilar NGLW of nuclear grade

SA508 LAS to 316L SS. These welds were made using high chromium-nickel alloy

filler Inconel FM52. The material thickness used was 40 mm representing the wall

thickness of a pressuriser surge line piping to pressuriser vessel dissimilar metal

weld. The plate form samples were 200 mm long.

Stage III was performed to produce and analyse welds resembling actual applications

in light water reactors as closely as possible within the budget and in laboratory

conditions. The welds were subjected to the same acceptance criteria as real-life

production welds. These were further supported by several additional analyses to

further assess the quality and to compare the properties achieved to literature.

The Stage III welding trials were conducted using all the knowledge developed

during this programme. The pre-heat and interpass temperatures were controlled and

monitored. Each welding pass was monitored visually and distortions measured

mechanically as per the experiences of the previous stages. The welding process was

recorded using the coaxial camera video system.

Visually sound, according to BS EN ISO 13919 [177], samples were subjected to

ASME IX radiographical acceptance as a production piece would be [176]. Industry-

standard, U.S. NRC Regulation 10 CFR Appendix G, [159] Charpy-V impact

toughness tests were conducted, with additional SEM analysis of the fracture

272

surfaces. Residual stress distribution was analysed using the contour method [162] to

assess the susceptibility to stress corrosion cracking [10]. Macrographical and

hardness investigations were conducted on full-size cross-sectional samples to

support microstructural analysis. Longitudinal and transverse (composite) tensile

tests were made after BS EN ISO 13919 [177] with added DIC analysis.

The sample design was as per Chapter 3. Two different V-groove angles, 4° and 5°

were examined; all tests are displayed in Table 6.1.

Table 6.1 Stage III of dissimilar metal narrow gap laser welds,

SA508 to 316L using Alloy 52

Target Groove Gas shield Notes

SA1 4° Stationary nozzles Thermocoupled

SA2 Gas shielding &

Groove angle study

5° Tri-blade nozzle

SA3 Sample generation 5° Tri-blade nozzle

SA4 Undercut control study 5° Tri-blade nozzle





6.1.1 Expectations

The Stage III trials were expected to provide samples with sufficient integrity for the

various analyses and comparisons planned, although there were several known

challenges. One challenge was the sluggish weld pool of FM52 [129] and the

susceptibility to solidification, ductility dip and liquation cracking [98, 128].

Accumulation of Al and Ti oxides to the weld during the repeated passes was also

reported causing lack of fusion [87, 128]. Observations from the previous stages in

this study were in line with the literature; achieving good fusion being the main

challenge. There were no indications in the literature to suspect inferior properties

273

for successful NGLW samples compared with conventional welding, although it was

recognized that the smaller weld pool would be subjected to faster cooling than the

larger melt pools of arc welding processes.

Many characteristics of the FM52 filler were already embarked upon on Stage II, this

stage being concentrated on the properties of the final pedigree materials and the

weld as a whole. Special emphasis was put on investigating the multi-pass tempering

effect; could it replace the costly PWHT usually required? The research questions

set for this stage were:

Is there a difference in weldability between S275 and SA508 Gr3 Cl2? Does

the change in materials affect the tendency or mechanisms leading to lack of

fusion?

Is ASME IX radiographical acceptance achievable using the actual

application materials?

How does the different hardenability in the different low-alloy steels affect

the hardness of the HAZ?

What are the resulting microstructures in HAZ and weld metal, are there

potentially brittle areas?

Are the welds capable of exceeding the U.S. NRC impact toughness

requirements for PCC boundary components?

Are there other known flaws like FM52 DDC, liquation or solidification

cracking or porosity?

What is the effect of multi-pass tempering to the hardness, microstructure,

grain size, impact toughness and tensile properties?

Is the multi-pass tempering effect uniform? Is there variation in the properties

through the thickness?

How do the residual stress properties compare to literature?

274

6.2 Materials and Experimental Methods

The experimental setup was based on the findings of the earlier stages of this study.

The pedigree materials were found to set new challenges for the welding setup.

Improvements to the equipment were conducted up to the fifth welding test and from

then on the equipment was kept unchanged.

6.2.1 Materials

The base materials of the Stage III welds were ASME SA508 Gr3 Cl2, a nuclear

grade pressure vessel steel and AISI 316L austenitic stainless steel. The filler wire

used was a high chromium-nickel alloy Inconel FM 52 (Alloy 52, ERNiCrFe+7).

The materials used for this stage of the study are presented in detail in chapter 3.2.3.

6.2.2 Experimental Setup and Welding Parameters

The experiences from the previous welding trials had narrowed the usable parameter

window. Therefore all the pedigree welds were started with similar parameters, see

Appendix I. The welding speed was maintained at 0.08 m / min. The starting laser

power was 3500 W. During the welding the laser power was adjusted according to

observations of fusion and undercut in stainless steel. Offset was set to provide a

spot size similar to the width of the welding groove bottom, although Feng et a.l.

2106 [45] recommended a slightly wider for their ferritic steel welds. Increasing the

spot size to touch the sidewalls was found to generate undercut. Pre-heat was varied

between room temperature and 125°C. Shielding gas flow rates varied up to

14 l / min.

The first pedigree metal weld experiment SA1 was conducted with 4° V-groove.

Based on the observations during the previous welds, stationary shielding gas

275

nozzles with brackets guiding the gas flow were applied to provide gas shielding, see

Fig. 6.1. The wire feed nozzle arrangement was changed to an adjustable holder, see

Fig. 6.2. Otherwise, the setup was similar to the previous similar metal and S275

dissimilar metal welds. One reason for using a stationary shielding gas arrangement

was that it allowed the crucial view to the weld with the laser illumination imaging

system. The weld was also thermocoupled with 16 thermocouples.

Fig. 6.1 Experimental setup for SA1 pedigree metal weld. Thermocoupling,

stationary gas shielding arrangement, pre-heating blankets and Lenskes

clamps.

276

Fig. 6.2 First application of the commercial wire nozzle holder, as used for SA1.

The most significant improvement over previous arrangements in Stage III of the

tests was that they were conducted using a micro-adjustable commercial wire feed

nozzle holder, see Fig. 6.2. The nozzle holder was attached to the welding head using

a round pipe, which allowed adjustment in rotation and height.

The welds SA2 to SA8 were conducted using a moving gas shielding nozzle capable

of injecting the gas to the bottom of the welding groove, see Fig. 6.3. While this was

indispensable for the gas shielding, it prevented the use of LII imaging. Therefore

the subsequent welds were monitored with an ambient light camera system using the

coaxial optics built into the welding head. This change resulted in a significant loss

of quality in the information recorded. At the same time, the V-groove was widened

to 5° as per distortion measurement results from SA1.

277

Fig. 6.3 Wire feed setup with a linear stage for height adjustment and tri-blade gas

nozzle arrangement. As used from SA2 onwards.

From SA2 onwards the design was improved by implementing a linear stage for

quicker and more precise height adjustment, avoiding unnecessary rotation, see

Fig. 6.3. This allowed reliable adjustment of the nozzle height needed when

changing the focal offset to adjust the spot size. The nozzle was first fitted with a

4 mm wide tip and later improved to a design with a 2.8 mm tip allowing more

transverse adjustability in the narrow welding groove. When the range of

adjustability of the nozzle was sufficient and the alignment of the wire was

satisfactory the results were good regardless of the tip design.

6.3 Experimental Work and Results

The experiments on Stage III were based on previous work. The welding parameters

were initially the same as for the DMW in Stage II, as only the ferritic LAS was

changed.

278

6.3.1 Experiment 1

The first experiment of SA508 DMW was named SA1. This weld was designed to

include temperature monitoring by 16 thermocouples. Pre-heating was used under

the SA508 LAS. To allow the use of the LII recording system the gas shielding was

done by using brackets at either end of the welding groove.

Fig. 6.4 The welding setup for SA1. SA508 LAS nearer the camera.

The welding of SA1 was started with 3500 W laser power and 1 m/min wire feed

rate. The SA1 took 16 filling passes which were then covered with one cosmetic

pass. All the passes used the mentioned wire feed rate, but the laser power was

varied according to Table 6.2. The gas shielding performed very poorly, so this

experiment was not given much emphasis. The thermocoupling data proved valuable

for the following work, the resulting graphs are displayed as Appendix VI and VII.

279

Table 6.2 Welding parameters and notes SA1.

Pass Temp

316L

Temp

S275

Thickness Power

[°C] [°C] [mm] [W]

pre 125 202 2.1 3500

1 128 155.0 5.1 3500

2 133.0 148.0 7.5 3000

3 142.0 152.0 10 3000

4 147.0 155.0 12.2 3000

5 138.0 144.0 14.7 3000

6 131.0 134.0 17 2900

7 81.0 81.0 20 2900

8 101.0 116.0 23 2900

9 121.0 124.0 25.7 2900

10 135.0 142.0 27.8 2900

11 130.0 146.0 30 2900

12 118.0 126.0 32.5 2900

13 119.0 131.0 35.5 2900

14 124.0 136.0 36.5 3500

15 38.7 3500

16 39.7 3500

17 5200

6.3.2 Experiment 2

The test SA2 was conducted with the latest gas shielding arrangement which had

three nozzles delivering the gas to the bottom of the groove, see Fig. 6.5. The weld

was started without pre-heat to limit the turbulent airflow over the sample.

280

Fig. 6.5 The gas and wire nozzle used for SA2

The filler wire rate was 1 m/min, except for the second pass, which was a re-melt to

rectify an apparent centreline crack. No crack was found in the analysis of the weld.

The starting welding power of 3500 W was later lowered to decrease the undercut,

which was removed by the pass 10, see Table 6.3. This, however, caused the wetting

of the sidewalls to become less reliable, which can be seen as poor fusion at passes

13 and 16.

281


Pass Temp

316L

Temp

S275

Thickness Power Feed

Rate

Fusion Under-

cut

[°C] [°C] [mm] [W] [m/min]

1 21.5 21.5 3500 1.0 50 20

2 42.5 43.5 1.2 3000 0.0 50 40

3 67.8 69.6 2.5 3000 1.0 50 20

4 63.0 63.4 4.2 3000 1.0 50 30

5 67.3 69.1 6.8 3000 1.0 50 40

6 66.0 68.8 9.4 3000 1.0 50 50

7 68.2 70.0 11.5 2500 1.0 50 30

8 72.5 76.7 14.0 2500 1.0 50 10

9 69.9 69.7 15.5 2500 1.0 50 5

10 77.0 75.9 17.7 2500 1.0 50 0

11 72.2 76.2 19.9 2500 1.0 50 0

12 65.2 69.5 21.8 2500 1.0 50 0

13 60.6 65.6 24.0 2500 1.0 10 0

14 69.0 68.7 26.0 2700 1.0 50 10

15 62.8 68.1 28.0 2500 1.0 50 5

16 71.5 72.9 30.0 2500 1.0 10 0

17 63.5 63.0 31.5 2700 1.0 50 5

18 68.8 71.6 33.5 2700 1.0 50 0

19 74.1 74.6 35.0 2700 1.0 50 0

20 50.0 50.5 37.5 3000 1.0 50 0

21 55.7 62.6 57.0 3500 1.0 50 0

22 60.0 64.5 77.0 5000 1.0 50 0

The weld suffered from undercutting in 316L from the beginning and laser power

had to be reduced to alleviate the problem. The occurrence of small undercut inside

the weld was not found detrimental as the geometry of the weld was not

compromised. The challenge was mainly to keep the undercut from growing too big

causing the type of sidewall collapse shown earlier in Fig. 5.6. Fig. 6.6 shows a

typical shielding and wetting performance for the weld SA2.

282

Fig. 6.6 SA2 first pass. The smooth wetting, symmetrical concave bead shape and

good gas shielding were characteristic for the first passes of all Stage III

welds. A and B are separate blocks to guide the shielding gas.

6.3.3 Experiment 3

The use of separate blocks as gas guides at the ends of the samples required manual

work for setting after each pass. For SA3 the gas shielding was further improved by

introducing fixed run in and –out grooves, instead of separate loose arrangements,

see Fig. 6.7. The weld was started 20 mm inside the bracket to allow for a longer

sample to be created. The gas shielding was good, see Fig. 6.8. The parameters used

are shown in Table 6.4

283

Fig. 6.7 The welding setup for SA3 with welded run in and –out brackets. Welding

direction down. Stainless steel on the left.

Fig. 6.8 3rd

pass of SA3. Note the oxide-free weld extending to the run on bracket

at start.

284


Pass Temp

316L

Temp

S275

Power Feed

Rate

Fusion Under-

cut

[°C] [°C] [W] [m/min]

1 54.7 89.2 3500 1.0 20 50

2 75.3 123.5 3500 0.0 30 50

3 102.4 131.0 3000 1.0 30 45

4 117.0 136.5 3000 1.0 20 50

5 131.3 135.6 3000 1.0 15 50

6 126.8 139.1 3000 1.0 30 50

7 127.8 134.9 2800 1.0 25 50

8 130.0 132.6 2800 1.0 25 50

9 128.2 129.5 2800 1.0 25 50

10 131.0 144.4 2800 1.0 25 50*

11 67.6 67.6 3500 25 50

12 101.4 101.6 3000 1.0 10 50

13 123.8 128.8 3000 1.0 10 50

14 116.6 109.2 2800 1.0 15 50

15 122.0 128.0 2800 1.0 25 40

16 114.0 117.4 2800 1.0 25 40

17 108.2 110.4 2800 1.0 25 40

18 102.3 104.3 2800 1.0 10 10

19 109.1 112.5 2800 0.0 10 0

20 98.2 101.7 2800 1.5

21 87.1 91.0 5500 1.5

22 81.0 85.0 7500 1.5

23 77.7 76.4 7500 1.5

24 114.9 114.5 7500 0.0

* Nozzle pinched due to run in plate breaking

The poor fusion of the first pass was remedied with a re-melt pass. The fusion

improved the fusion to what was considered satisfactory. The weld was interrupted

at pass no 10 as the run in arrangement contracted rapidly as a tack weld cracked.

The same issue stopped the pass no 11 after 50 mm had been welded. This can be

seen as a loss of interpass temperature control in the data.

285

Initial gas shielding experiences for the first passes were good. After the weld had

gathered some thickness, the gas was able to escape more easily via the rear corner

of the trailing nozzles at the starting position of the pass. This was seen as increased

oxidation until all the nozzles were directly above the previous bead. This was

solved in the following welds by manually lifting also the run in bracket bottom as

the weld progressed.

6.3.4 Experiment 4

SA 4 was welded using the best practices developed. A pre-heat of 75°C was

applied. The welding parameters were the same as previously, laser power being

varied according to the observations on undercut and fusion. The shielding gas setup

was improved by filling in the gap under the trailing nozzles at the starting position

to ensure even gas flow distribution. The course of welding is displayed in Table 6.5.

286


Pass Temp

316L

Temp

S275

Power Feed

Rate

Fusion Under-

cut

[°C] [°C] [W] [m/min]

1 75.7 75.0 3500 1.0 50 0

2 85.6 87.4 3000 1.0 50 20

3 87.7 91.1 3000 1.0 50 20

4 85.7 87.9 3000 1.0 50 20

5 85.6 92.0 3000 1.0 50 20

6 75.0 76.3 2900 1.0 50 5

7 86.1 89.4 2900 1.0 50 1

8 87.6 90.4 2900 1.0 50 1

9 79.9 81.3 2900 1.0 50 0

10 84.3 86.0 2900 1.0 50 5

11 87.9 90.5 2900 1.0 50 5

12 86.6 88.9 2900 1.0 50 5

13 85.9 86.7 2900 1.0 50 10

14* 78.9 80.8 2900 1.0 50 30

15** 70.0 70.0 2900 1.0 - -

16** 93.0 92.0 2900 1.0 - -

17** 79.5 78.7 2900 1.0 - -

18 69.5 69.5 2900 1.0 50 30

19 75.2 78.2 2650 1.0 40 40

20 76.2 71.8 2650 5.2 50 -

21 71.6 73.5 2650 3.5 - -

22 66.2 66.2 3100 1.5 - -

23 61.0 61.0 3100 0

24

* Filler wire stuck

** Partial welds to fill the layer

The SA4 started with good fusion quality and acceptable undercut characteristics.

However at pass no. 14 the wire feeder system failed and the pass was interrupted.

The repair of the pass was unsuccessful, requiring three separate runs to fill the layer

leaving lots of discontinuities and flaws in the weld. Although the small undercut

near the top allowed a smooth, Fig. 6.9, and symmetrical, Fig. 6.10, capping, this

weld was not analysed further.

287

Fig. 6.9 Smooth capping of SA4 with just some minor lack of fusion type defects.

Fig. 6.10 Symmetrical capping of SA4.

6.3.5 Experiment 5

The experiment SA5 was welded as the previous passes. After careful examination

of the previous interpass temperatures, the pre-heat temperature was set at 95°C. The

interpass temperature was held between 92°C and 105°C. There were no issues with

interruptions to the welding. The shielding gas performance was acceptable and

augmented by mechanical cleaning after each pass. The resulting weld was very

good. The parameters for the weld are displayed as Table 6.6.

288


Pass Temp

316L

Temp

S275


Rate

Fusion Under-

cut

[°C] [°C] [mm] [W] [m/min]

1 94.7 86.1 0.0 3500 1.0 50 0

2 99.5 97.2 1.9 3250 1.0 49 0

3 97.7 95.9 4.0 3500 1.0 50 0

4 101.4 103.8 5.5 3500 1.0 50 0

5 102.4 106.8 7.5 3500 1.0 50 0

6 88.8 87.3 9.7 3500 1.0 50 0

7 101.4 99.7 12.0 3500 1.0 50 0

8 104.8 106.9 13.5 3500 1.0 50 0

9 104.3 105.6 15.5 3500 1.0 50 0

10 96.8 97.9 17.5 3500 1.0 50 0

11 91.9 93.0 19.5 3500 1.0 50 0

12 99.0 100.8 21.5 3500 1.0 50 0

13 103.4 105.7 23.5 3500 1.0 50 0

14 103.8 106.0 25.5 3500 1.0 50 0

15 97.3 98.4 27.5 3500 1.0 50 0

16 94.3 96.1 29.5 3500 1.0 50 0

17 93.3 95.0 30.5 3500 1.0 50 15

18 93.5 93.2 33.0 3350 1.0 40 3

19 88.4 89.2 34.7 3500 1.0 45 0

20 89.8 91.3 - 3500 1.0 49 0

21 78.3 78.6 - 3800 1.3 50 0

The fusion quality of SA5 was good throughout the weld, compare Fig. 6.11 and Fig.

6.12. The undercut was kept under control except for a brief occurrence at pass no.

17, which was remedied by a slight reduction of power for the subsequent pass. Fig.

6.13 presents the finished weld showing little oxidation and good fusion.

289

Fig. 6.11 First pass of SA5. Smooth wetting and bead surface with low oxidation.

Fig. 6.12 12th

pass of SA5. Weld quality very similar to pass 1. Slight colouration of

the SA508 surface due to heating starting to appear.

Fig. 6.13 Finished weld SA5 uncleaned, as welded. Note lack of oxidation.

290

6.3.6 Experiment 6

The SA6 was welded to replicate the good results of SA5. The pre-heat temperature

was matched as closely as possible at 92°C. The welding setup was the same. The

progress of the welding was good with no significant issues. Minute undercut was

noticed on several passes, but the extent was considered insignificant. The weld

filled at the same rate as previously and there were no interruptions. The parameters

are displayed in Table 6.7.


Pass Temp

316L

Temp

S275


Rate

Fusion Under-

cut

[°C] [°C] [mm] [W] [m/min]

1 92.2 88.3 0.0 3500 1.0 50 0

2 100.2 98.6 1.7 3500 1.0 50 0

3 99.5 100.7 3.5 3500 1.0 50 1

4 99.8 101.3 5.3 3500 1.0 50 4

5 97.7 98.5 7.2 3500 1.0 50 4

6 97.0 96.7 9.7 3500 1.0 50 2

7 99.2 98.3 12.0 3500 1.0 50 1

8 100.0 99.7 13.7 3500 1.0 50 0.5

9 100.3 101.6 15.3 3500 1.0 50 0

10 97.6 98.2 17.4 3500 1.0 50 0

11 93.3 94.0 19.2 3500 1.0 50 0

12 99.6 102.8 21.4 3500 1.0 50 4

13 100.5 101.9 23.2 3500 1.0 50 6

14 100.2 102.2 25.5 3450 1.0 50 1

15 97.8 100.6 26.9 3500 1.0 50 0

16 92.7 92.8 28.9 3500 1.0 50 0

17 93.4 96.6 30.9 3500 1.0 50 0

18 99.9 102.8 32.5 3500 1.0 50 0

19 99.3 100.3 34.6 3500 1.0 50 0

20 95.0 96.3 36.3 3500 1.0 40 0

21 80.3 83.1 * 3850 1.3 50 0

291

A sample of the good weld quality of SA6 is presented as Fig. 6.14. The finished

weld is presented as Fig. 6.15. The minor underfilling near the end of the weld was

attributed to filler feed issues at some of the passes.

Fig. 6.14 SA6, 11th

pass, low oxidation, smooth bead and good wetting of the

sidewalls

Fig. 6.15 Finished weld SA6. Minor underfilling (arrow)

6.3.7 Experiment 7

The test SA7 was conducted the same way as the previous welds. Due to the slight

undercut detected in SA6 the welding power was reduced from the 2nd

pass on. A re-

melt was applied at pass 3. The power reduction resulted in issues with fusion as

Table 6.8 shows. Two re-melt passes were applied to remedy the issues with poor

292

fusion after passes 2 and 10. The re-melt was able to smoothen the bead and improve

fusion.


Pass Temp

316L

Temp

S275


Rate

Fusion Under-

cut

[°C] [°C] [mm] [W] [m/min]

1 87.7 80 0.0 3500 1.0 50 0

2 86.6 85.3 2.4 3000 1.0 20 0

3 84.2 84.6 4.9 3000 0 30 5

4 89.3 92.0 4.9 3000 1.0 45 0

5 89.9 93.0 7.5 3000 1.0 45 0

6 66.8 66.6 10.0 3200 1.0 50 0

7 89.3 92.9 12.0 3000 1.0 30 0

8 88.4 92.0 14.0 3100 1.0 40 0

9 88.2 89.8 16.0 3100 1.0 40 0

10 89.0 89.0 18.5 3100 1.0 20 0

11 78.4 78.6 19.5 3500 0 50 15

12 89.5 91.1 20.0 3500 1.0 50 10

13 92.7 94.7 22.5 3500 1.0 50 5

14 101.8 104.8 25.0 3500 1.0 50 2

15 102.1 104.2 27.0 3500 1.0 50 4

16 88.8 88.4 28.5 3500 1.0 50 20

17 88.5 92.3 3500 1.0 30 25

18 93.9 97.9 3500 1.0 10 50

19 90.4 94.0 3700 1.0 30 *

20 98.0 98.5 3700 1.3 50

21 98.5 101.7 3700 1.3 50

22 93.5 96.0 4200 1.3 50

23 4500 0 50

* undercut collapsed into the weld pool

The usually good parameter setup produced unexpected undercut starting from pass

16. This resulted in a catastrophic collapse of the sidewall at pass 19, see Fig. 6.16.

The finished weld SA7 is presented as Fig. 6.17. This weld was not analysed further.

293

Fig. 6.16 Collapse of the sidewall (arrow) SA7, 19th

pass.

Fig. 6.17 SA7 finished, with fusion flaws.

6.3.8 Experiment 8

Experiment SA8 was a repetition test to replicate the results of SA5 for sample

production. The welding parameters and data are presented in Table 6.9.

294


Pass Temp

316L

Temp

S275


Rate

Fusion Under-

cut

[°C] [°C] [mm] [W] [m/min]

1 99.2 96.3 0.0 3500 1.0 50 1

2 95.2 92.1 2.3 3500 1.0 50 5

3 99.1 98.3 2.3 3500 1.0 50 3

4 100.1 100.0 4.6 3500 1.0 50 3

5 103.4 101.9 6.8 3500 1.0 50 3

6 86.2 84.8 9.3 3500 1.0 50 8

7 98.8 100.3 11.3 3500 1.0 50 7

8 102.0 103.2 13.1 3500 1.0 50 7

9 101.3 102.3 15.3 3500 1.0 50 7

10 101.6 102.4 17.3 3500 1.0 50 9

11 90.0 89.8 19.5 3500 1.0 50 9

12 96.6 100.7 21.7 3450 1.0 50 11

13 100.2 101.5 23.5 3400 1.0 40 15

14 88.3 89.5 25.8 3500 1.0 45 * 15

15 94.5 97.2 27.8 3500 1.0 30 -

16 102.4 106.3 29.3 3500 1.0 40 -

17 90.2 89.5 29.9 3500 1.0 45 -

18 72.5 71.0 3800 1.3 43 -

19 82.5 84.4 3800 1.2 45 -

20 89.0 91.3 4000 1.2 45 -

21 95.1 97.1 3900 1.0 45 -

* A single deep depression appeared at the 316L fusion boundary

The SA8 progressed well until pass no. 13 when the fusion quality unexpectedly

decreased. Although the grading for the average fusion quality does not show it, a

deep lack of fusion related recession appeared in the weld at pass no. 14 see Fig.

6.18. This rendered the weld unusable for further analysis.

295

Fig. 6.18 A depression in the 14th

pass, SA8.

6.4 Analysis of Process Characteristics

6.4.1 Welding Process Monitoring

Due to the limitations set by the shielding gas arrangement the laser illumination

imaging was only used for monitoring SA1. The rest of the welds were monitored

using the coaxial camera system.

The SA1 LII video material shows a stable melt pool with good wetting to the

sidewalls, but with severe oxidation, see Fig. 6.19. The wire feed was functioning

well and alignment was considered good. The melt pool did not appear to have a

strong Marangoni flow to either side, as the oxides spread evenly on the weld, with

only a slight tendency to drift towards the stainless steel.

296

Fig. 6.19 SA1, first pass. LII snapshot. Good wetting and alignment. Excessive

oxidation.

The 13th

pass of SA1 is displayed in Fig. 6.20. The groove has contracted below

4.37 mm and the laser is severely overheating the stainless steel top corner. Small

indications of this were already observed during the previous pass where the top gap

was measured at 4.44 mm after the welding. The sidewalls were not damaged

previously due to the conicality of the defocused laser beam.

Fig. 6.20 SA1 13th

pass. LII snapshot. Groove contracting excessively, laser beam

melting the stainless steel prematurely, very severe oxidation.

LII imaging of SA1 displayed that the parameters used generated a stable melt pool

and acceptable wetting of the side walls. The observations of the large amounts of

oxides with the stationary gas shielding arrangement emphasized the need for the

297

development of gas shielding. The wire feed setup used was proven reliable,

adjustable and free from vibrations.

The second phase of optical monitoring of the welds was the use of the coaxial

camera system provided by the laser head manufacturer. The camera provided was

changed to the one used with LII for allowing the recording of the videos. The

decrease in image quality was significant, but some crucial observations could still

be made.

6.4.2 Mechanical Distortion Measurements

The need for a wider V-groove was confirmed by the mechanical measurements of

the weld SA1. The resulting graph is displayed as Fig. 6.21. The minimum gap

measured and calculated was 4.25 mm. Geometrical calculations indicated that a 5°

V-groove would be required. The discontinuity in the graph was due to an

exceptionally long break between passes, during which the interpass temperature

control was lost and the sample cooled excessively.

298

0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 204

5

6

7

8

Ga

p W

idth

(m

m)

Pass no:

direct measurements


Fig. 6.21 Welding groove top width of SA1 using 4° V-groove. Note gap contracting

below 5 mm. Error bars represent min and max values measured.

The average of successful tests using the new 5° groove angle is displayed in Fig.

6.22. It indicates that the groove contracts very close to the set 5 mm as the groove

fills up. This allows the use of constant spot size, leading to using constant focal

offset. This translates to no need for wire or shielding gas nozzle height changes

during the welding.

299

0 2 4 6 8 10 12 14 16 18 204

5

6

7

8

Ga

p W

idth

(m

m)

Pass no:

direct measurements


Fig. 6.22 Welding groove top width of SA series using 5° V-groove. Note gap

remaining over 5 mm until the end of the weld. Average of 6 tests. Error

bars represent min and max values measured.

Unlike in the mild steel dissimilar tests the groove contraction did not appear to have

distinct phases, but it contracted near linearly. As a conclusion, the 5° V-groove was

proven to provide near-constant welding pass width for each pass of the weld.

6.4.3 Effect of Remelting

Re-melting was applied in certain cases to repair an uneven or poorly fused welding

pass. Generally, the re-melting was able to heal some of the irregularities, but not all.

A poor beginning of the second pass of SA7 is presented as Fig. 6.23. The same bead

after a re-melting pass is displayed as Fig. 6.24. The re-melt effect on the bead

roughness and fusion can be seen as a reduction of the length of the irregular weld.

300

Fig. 6.23 2nd

pass SA7, poor wetting and fusion of the first 75 mm, see arrow.

Fig. 6.24 3rd

pass SA7, a re-melt applied to the previous pass has reduced the length

of the poorly fused area by 50 %, see arrow.

6.4.4 Temperature Distribution

The weld SA1 was monitored by thermocoupling. The data is presented as

Appendices VI and VII. The thermocouples were welded to the top and root sides of

the samples as Fig. 6.25 and Table 6.10.

301

Fig. 6.25 Thermocoupling of SA1, top side. Bottom similar.

Table 6.10 Thermocouple numbering SA1

TC Location Material

1-4 Root SA508

5-8 Root 316L

9-12 Top SA508

13-16 Top 316L

From the thermocoupling data, it can be seen that the samples need one hour for the

temperature to be fully soaked. The weld in question was pre-heated from the ferritic

steel side only. The data indicates that using this setup the asymmetric pre-heating

temperature evens out after just 3 passes. As the issues with the welding emerged

typically after 10 or more passes, the conclusion was that a single-sided pre-heating

did not produce benefits.

6.5 Analysis of Weldments

As with the previous dissimilar welds, the successful Stage III welds were subjected

to the industry-standard ASME IX radiographical acceptance. The radiographical

analysis was followed by cutting for contour method residual stress analysis. The

302

contour cut had to be made to an intact sample to preserve the residual stress

distribution. After the data for the contour method was found dependable, the

samples were EDM machined to produce test coupons for hardness, microstructural

and other mechanical analysis.

6.5.1 Radiographical Analysis and ASME IX Acceptance

The welds were analysed visually as per literature [161]. A summary of the results is

presented in Table 6.11. With this data and observations of the weld pool

behaviour during welding the welds SA2, SA5 and SA6 were considered as

successful and were analysed by radiography. The resulting

radiographical reports are presented in Appendix II. The weld SA5 was

found acceptable by ASME IX, the others having issues with lack of fusion,

mainly to the sidewalls, see

Table 6.12. Most issues were located at the beginning of the weld.

Table 6.11 Stage III weld quality visual inspection

SA1 SA2 SA3 SA4 SA5 SA6 SA7 SA8

Oxidation U VG VG VG VG VG VG VG

Symmetry A VG VG VG VG VG VG VG

Fusion U A U U VG VG U U

VG-Very good, A-Acceptable, U-Unacceptable

Table 6.12 Results and interpretation of radiographical analysis of Stage III welds

Sample ASME IX First half Second half

SA2 Unacceptable Intermittent lack of fusion Intermittent lack of fusion

SA5 Acceptable Localised root undercut No visible defects

SA6 Unacceptable Intermittent lack of side

wall fusion

Intermittent lack of side

wall fusion

303

Lack of fusion was found to be the main form of defect in the welds. Very few other

issues like pores or cracks were detected. None of these was severe enough to

warrant rejection by the ASME IX.

6.5.2 Macrographical Analysis

The samples SA2 and SA5 were cross-sectioned for macrographical analysis after

being contour cut. The weld SA2 was used for analysing the lack of fusion defects

found in radiography. The weld SA5 was used for the analysis of an acceptable

ASME IX quality weld. Different etching processes were required to reveal the

microstructures of different materials. Photomontage of SA5 after Nital and

ammonium persulfate etching is presented as Fig. 6.26. In the figure, the 316L is left

unetched as oxalic acid was not found to reveal additional information in

macrography.

304

Fig. 6.26 Photomontage of Nital (for SA508) and ammonium persulfate (FM52)

etched weld cross-section. Sample SA5 161.

The main challenges in the welding were the propensity to lack of fusion and

undercutting of the stainless steel. These were found to be related to the welding

parameters used. The fusion results for each individual welding pass and parameters

can be interpreted from the cross-sectional macrographs. The fusion line shape will

correspond to the amount of undercut occurring.

6.5.2.1 SA2

To explain the resulting weld quality of each pass, the welding parameters used must

be known. The welding of SA2 is presented as an example. The fusion quality of

each pass was monitored during the welding process. Fusion to side walls was given

a score out of 50 where 50 indicates flawless fusion. A completely flat bead profile

305

would be graded as 0 and smooth stepless fusion as 50. A verbal comment was

added to address properties like the regularity of the fusion and the trend of change.

Undercutting of stainless steel was also monitored. The undercut was rated on a

similar scale, 0 indicating no undercut and 50 severe undercut bound to cause issues

later in the weld. For scale, a 1 mm deep undercut was considered detrimental and

graded 50.

An abridged presentation of the main welding parameters and the resulting weld

quality for each pass, as commented during welding, are presented as Table 6.13.

306

Table 6.13 Record of basic welding parameters, fusion quality and undercut in weld

SA2

Pass Feed

rate

Power Fusion /

bead shape

Score Undercut Score

1 1 3500 Good 50 Symmetrical 20

2* 0 3000 Good 50 Increased on SS side,

quite large

40

3 1 3000 Good 50 Smaller than

previous

20

4 1 3000 Good 50 Slightly larger 30

5 1 3000 Good 50 Slightly larger 40

6 1 3000 Good 50 Very large 50

7 1 2500 Good 50 Decreasing 30

8 1 2500 Good 50 Very small 10

9 1 2500 Good 50 Almost none 5

10 1 2500 Good 50 No undercut 0

11** 1 2500 Good 50 No undercut 0


13 1 2500 Flat 15 No undercut 0

14 1 2700 Good 50 Small 15


16 1 2500 Flat 10 No undercut 0





21 1 3500 Good 50 No undercut, weld

almost full

0

22 1 5000 Good, some lack

of coverage

50 Full weld 0

* note: re-melt with no wire

** intermittent matte bead surface due to oxidation from this bead on

The weld was started with the standard 3.5 kW laser power for the first pass. The

second pass was run as a re-melt to rectify signs of a superficial centreline

solidification crack. For the second pass, this was reduced to 3.0 kW. The power was

reduced further to 2.5 kW at pass 7 and onwards to counter the formation of the

undercut. This power setting produced reliable welds until pass 13, which suffered

from poor wetting of the side walls. For pass 14 the power was increased to 2.7 kW

307

to improve the wetting. Power had to be reduced back to 2.5 kW immediately to

counter emerging undercut. Power was increased again at pass 17 due to poor

wetting. Last filling pass using normal focal spot size was no. 20 and it was covered

by two passes with increased spot size.

Intermittent lack of fusion was found in SA2 in the radiographical analysis. No

cracking was reported. These results can be confirmed by the cross-sectional

samples. Nital was found to be the best etchant to show up the lack of fusion,

although the stainless steel fusion line is not visible, see SA2 98 presented as Fig.

6.27. The fusion issues detected are several interpass lack of fusion defects

beginning at pass no. 11, lack of fusion to both base materials at pass 17, lack of

fusion to SA508 at pass 19.

Fig. 6.27 Cross-section of an ASME IX rejected dissimilar metal weld showing

various fusion defects. Nital etch. 316L fusion line superimposed as dotted

line. Sample SA2 98

308

When comparing Table 6.13 and Fig. 6.27 several connections were made. The re-

melt pass was seen as a slight widening of the fusion zone above the first pass. The

interpass fusion issues found after pass 11 began at a point corresponding to

comments of the bead surface starting to show signs of excess accumulated

oxidation. The 16th

pass was noted to have a nearly flat bead shape, which resulted in

lack of fusion to the sidewalls during the following pass.

The interpass lack of fusion defects were concluded to have their origin in the

accumulation of oxides on the weld bead. No indications of fusion issues were

observed until several repetitions of individual welding passes and they appear

simultaneously with the first notes of signs of excess surface oxidation. As the best

available gas nozzle design was being used different gas flow rates were trialled to

counter this in subsequent welding trials. A significant improvement was also

achieved by reducing the accumulation of oxides by meticulous cleaning of each

pass with a wire brush.

A connection between poor wetting and lack of fusion to the sidewalls was observed.

Poor wetting caused square-cornered shape for the weld bead. The following pass

would not be able to penetrate and melt the extremes of the corners reliably. In SA2

the 16th

pass was found not to wet the sidewalls properly resulting in the 17th

pass

having lack of fusion to the side walls. As the radiographical report indicated that the

lack of fusion was intermittent, it is possible that the effects of a similar flat bead

shape at pass 13 were missed at the location of this particular cross-section.

The poor wetting observed in welding of SA2 was found to be related to too low

laser power. The limiting factor for laser power was the undercut generation in the

stainless steel. The laser power reduction at pass 7 was conducted after the undercut

309

was found to have increased to an unacceptable level. The reduction of power kept

the undercut well under control. However, when, at pass 14, the power was increased

by just 200 W to remedy the effect of poor wetting, the undercut began to increase

immediately.

310

6.5.2.2 SA5

For comparison, an ASME IX acceptable weld SA5 is presented. An abbreviated

version of the welding parameters and observations is presented as Table 6.14. The

macrograph of the weld is displayed as Fig. 6.28. The LAS HAZ in SA5 is more

uniform than in SA2 due to the welding parameters being kept more constant over

the progression of the weld.

Table 6.14 Record of basic parameters and fusion quality in test SA5

Pass Feed rate Power Fusion / wetting Score Undercut Score

1* 1 3500 Good 50 No undercut 0

2* 1 3250 Good 49 No undercut 0















17** 1 3500 Good 50 Small 15

18 1 3350 Poor 40 No undercut 3

19 1 3500 Acceptable 45 No undercut 0


21 1.3 3800 Good 50 No undercut 0

* poor wetting for ~30 mm at the beginning

** shielding nozzle failure at the beginning of the weld

311

Fig. 6.28 Cross-section of an ASME IX accepted dissimilar metal weld showing

near-perfect fusion. Nital etch. Fusion line superimposed in red. Sample

SA5 161

Despite the ASME IX acceptance, one incident of interpass lack of fusion was found

at pass no. 18. This was right after an equipment failure had disabled the shielding

gas system from operation, causing a severely oxidised bead. Although the oxidized

surface of the contaminated pass no. 17 was cleaned to the best ability, some

interpass LoF still occurred. The wetting and shape of the resulting bead were not

affected.

In the experimental setup, each pass would be welded using fixed parameters. The

parameter adjustments had to be made between passes. A welding setup with real-

time adjustment would ease the process in real applications. No weaving or pulsing

of the wire or beam was available.

312

The weld SA2 shows the challenges involved in NGLW of dissimilar metals SA508

Gr2 Cl2 and 316L using FM52 filler. SA5 demonstrates that near-perfect fusion was

achieved. The main issue in the welding was lack of fusion. Two modes, LoF to the

base materials and interpass LoF were observed.

One major contributor to fusion challenges was found to be the propensity to

oxidation of the FM52 nickel alloy. Possible solutions include improving the gas

shielding and removing the accumulated oxide layer mechanically after each pass.

Another viable option is using FM52 only to the thickness considered necessary for

achieving desired corrosion properties and filling the rest of the weld using FM82.

Alloy 82 is much less prone to oxidation during welding. Although inferior in

corrosion resistance, Alloy 82 has very similar mechanical properties as the CTE to

Alloy 52 and would yield a mechanically similar weld.

6.5.3 Hardness Evaluation

Hardness testing was performed to assess the effect of NGLW on the hardness of the

dissimilar metal weld sample. Sample SA5 157 was hardness mapped using Vickers

method HV 0.3. The choice of Vickers load chosen was a compromise between the

sensitivity for small local variations in the material and minimum spacing of the

indents to achieve good resolution. A base grid of 2 x 2 mm was used to cover the

whole sample. The grid density of the map was increased to 0.4 x 1 mm at the areas

with steep gradients. The dense grid was used near and inside the fusion zone, where

hardening phenomena were expected. Base materials were of less interest and were

mapped mainly to provide large enough area of even hardness points to minimise

interpolation artefacts in the produced maps. The resulting hardness map for SA5

157 is presented as Fig. 6.29.

313

Fig. 6.29 Hardness map of SA5 157. SA508 on the left, FM52 middle and 316L

right. Approximate fusion line in red. (HV 0.3)

In a dissimilar metal weld, the opposing sides were known to generate very different

hardness profiles. For base materials, the highest hardness in SA508 was 275 HV

and for the 316L it was 175 HV. Significant hardening was found in SA508 heat

affected zone, the hardest single point measurement being 480 HV. This was reduced

by multi-pass tempering to 350 HV. The 316L HAZ was found to harden via strain

hardening to a maximum of 275 HV.

From the SA5 157 hardness profile, it can be seen that the high hardness HAZ is

2.4 mm wide. The area is very uniform along with the depth of the weld. The highest

hardness was measured only at a small area near the top of the weld. The area of

high hardness is only 3.5 mm deep, matching the HAZ caused by the last pass.

314

Two mechanisms leading to hardening were found. The ferritic steel hardened by

phase transformation hardening via the martensite reaction. This was found to

generate the highest hardness observed in the sample. The austenitic stainless steel

and the austenitic weld metal were hardened by strain hardening due to the stresses

generated by the contraction of the solidified but still cooling weld metal.

The increase in hardness in the SA508LAS HAZ was generated by the rapid cooling

initiating the martensite reaction in the LAS. A detailed map, see Fig. 6.30, using a

grid of 200 x 200 µm was generated to provide more precise information of the

extent of the hardening. The map was generated on the same sample previously

mapped with a less dense grid. Some points were manually offset to avoid previous

indents and discontinuities. Only the HAZ generated by the last pass was observed to

have high hardness, the HAZ of underlying passes having been tempered by the

subsequent welding passes, indicating an effective multi-pass tempering. One factor

contributing to different HAZ hardness near the top was that the last pass was

welded using a larger laser spot size and with higher power increasing the heat input.

This, however, does not manifest itself as widening of the HAZ in hardness or

microstructure as the hardened area is uniform up to the top surface of the sample.

315

Fig. 6.30 Detail of the high hardness HAZ in SA508. Test point grid 200 x 200 µm.

Sample SA5 157

The area of high hardness observed had a distinctive boundary. Inside the area, the

hardness varied from 400 to 480 HV. The highest value 480 HV, according to

Grange [201], indicates to an untempered martensitic grain structure in 0.18 %

carbon steels, such as the SA508 Gr3 Cl2.

The weld metal shows opposite hardness gradient to the LAS HAZ. The

phenomenon is illustrated in Fig. 6.31. For clarity, a scale more suitable for

presenting the variation in hardness due to strain hardening is selected. To allow

comparison, the same scale of mapping used previously for the S275 mild steel

welds is used. Magenta areas have hardness exceeding 250 HV.

316

Fig. 6.31 Hardness map with an alternate scale to emphasize strain hardening.

Sample SA5 157

The last pass was found the softest and hardening was detected in the preceding

passes. The underlying passes had been subjected to the compressive stresses

generated by the cooling of the subsequent passes. This caused strain hardening of

the austenitic FM52 material.

In the 316L stainless steel, the strain hardening followed the welding distortions near

the welding groove. The tensile stress generated by the cooling and contracting weld

strained the stainless steel causing a field of elevated hardness near the fusion line

and further into the base metal. Near the top, in the 316L HAZ a similar softer area

as in the topmost weld bead was observed.

The scale of the welding distortions causing the stainless steel hardening is presented

in a profile of the sample top surface contour. As an example of the phenomenon,

317

Fig. 6.32 displays the local distortions on the top surface of the sample SA5 157. The

stainless steel was distorted in the vertical direction by 1.5 mm. The distortion spread

to at least the 17 mm area observed in the sample, gradually tapering off. The

amount of distortion in the much higher strength SA508 is negligible in comparison.

Fig. 6.32 Outline of top of the weld SA5.

6.5.4 Microstructural Analysis

Hardness analysis indicated considerable hardening in all three metals involved in

the weld. The grain structural changes in non-phase transforming austenitic materials

were diffusion and grain size analysis. The complicated thermal histories

experienced by the ferritic phase transformation hardened SA508 steel was predicted

to give more varied results.

The SA508 Gr3 Cl2 is a quenched and tempered alloy. The grain structure of the

base material is martensite with cementite precipitates, see Fig. 6.33.

318

Fig. 6.33 SA508 Base material, tempered martensite (light). Carbon precipitated as

cementite particles (grey) austenite grain boundaries (dark), SA5 159

Nital etch.

The HAZ confirms the hardenability of SA508 Gr3 Cl2. The primary grain structure

in the coarse grain zone is martensite, Fig. 6.34. This agrees with the hardness

measurements. The area of highest hardness was found near the fusion line at the top

pass HAZ, Fig. 6.35. The grain structure is again martensitic with some residual

austenite. Fig. 6.36 displays the transitionary zone between CGHAZ and FGHAZ.

20 µm

319

Fig. 6.34 Martensitic (dark) HAZ with bainite (light) adjacent to FL of the top pass,

SA5 159 Nital etch.

Fig. 6.35 Martensitic/bainitic structure at the area of highest hardness, 200 µm

perpendicular) from FL, top pass, SA5 159 Nital etch.

20 µm

20 µm

320

Fig. 6.36 Transition from CGHAZ (right) to FGHAZ (left), approx. 1 mm from FL,

top pass, SA5 159 Nital etch.

In the FGZ the grain is very small, Fig. 6.37, and the hardness is still high. The

structure was a mixture of bainite and martensite. The IGHAZ is presented as

Fig. 6.38.

20 µm

321

Fig. 6.37 FGHAZ, approx. 1.5 mm from FL top pass, SA5 159 Nital etch.

The intercritical HAZ is displayed as Fig. 6.38

Fig. 6.38 ICHAZ, approx. 2 mm from FL, top pass, SA5 159 Nital etch.

20 µm

20 µm

322

The multi-pass tempering effect can best be seen in Fig. 6.39 as banding. The HAZ

caused by top pass is visible with much coarser grain structure.

Fig. 6.39 Vertical stitched image of SA508 HAZ. SA5 159 Nital etch.

323

The multi-pass effect is uniform through the thickness of the HAZ. No significant

visible change can be observed beyond three tempering cycles. The largest change is

caused by the first tempering as predicted by the hardness maps. Although the multi-

pass tempering produces a uniform structure in general, significant banding can be

observed. The fine spheroidised grain structure is banded by more coarse regions.

After two tempering cycles, the CGHAZ in the weld can be seen refined to a fine

equiaxial grain structure. The tempered SA508 HAZ is presented as Fig. 6.40. The

pictured location is the 7th

to the last pass, which has gone through 6 tempering

cycles by the overlaying passes. A significant change in microstructure in

comparison to the martensitic HAZ at top pass Fig. 6.34 can be seen. The grain is

small which suggests good impact toughness. Phase recognition was found

challenging as the tempering history was complex.

Fig. 6.40 Multi-pass tempered CGHAZ, 7th

to last pass, 15 mm from top, SA5 161

Nital etch.

20 µm

324

Using ammonium persulfate etching the weld metal microstructure was revealed.

The nickel Alloy 52 solidifies as austenite and there is no solid-state phase

transformation during cooling [75]. The microstructure of the weld metal was found

austenitic, with the distinct microstructure of fully austenitic Type A solidification.

Three main grain boundary types, solidification subgrain boundaries (SSGB),

solidification grain boundaries (SGB), Fig. 6.41 and migrated grain boundaries

(MGB), Fig. 6.42 and Fig. 6.43, were detected.

Fig. 6.41 Austenitic weld metal at 316L FL. 1) SSGB and 2) SGB. SA5 161

Ammonium persulfate etch.

1

2 100 µm

325

Fig. 6.42 Austenitic weld metal at 316L. 1) SSGB, 2) SGB and 3) MGB. SA5 161


Fig. 6.43 FM52 weld metal at SA508 FL. 1) SSGB, 2) MGB and 3) martensitic

region. SA5 161 Ammonium persulfate etch.

2 1

3

1

2

100 µm

200 µm

3

326

In Fig. 6.43, arrow 3, a white band of weld metal next to the fusion line is observed.

This area is martensitic and exhibits high hardness compared to the adjacent WM

and HAZ. The dark lines parallel to the fusion line, offset by 30 µm, resembling

Type II boundaries can also be seen. They were not considered pure Type II

boundaries as they don’t have the characteristic pattern of the prior austenite grain

and are too far from the FL for being generated by the Type II mechanism. They

were observed infrequently and were found discontinuous.

The SGBs that form have the visible compositional component as well as a

crystallographic component. In certain conditions, the crystallographic component

migrates over the compositional structure. The driving force for this is the lowering

of boundary energy which happens by straightening the grain boundary. The newly

formed boundary is called migrated grain boundary (MGB). Reheating such as

multi-pass heat treatment can increase grain boundary migration. The MGBs in weld

metal were of great interest as solidification cracking nearly always occurs along

with them [75].

The weldability of Alloy 52 was known to be an issue. The flaw types presented in

the following chapters were encountered. The flaws were random in nature and did

not appear at all welding passes. Two types of lack of fusion defects were observed.

Some suspected ductility dip cracking was found as well as some small pores. It was

noted that solidification cracking or liquation cracking was not found.

In some cases, the crystallographic component of the SGB migrate from the

compositional component to form an MGB. The driving force for MGB formation is

the lowering of the boundary energy by straightening the originally tortuous SGB.

This makes them more prone to cracking [75]. In comparison to Alloy 82, the MGBs

327

of the high chromium Alloy 52 are straighter which increases the susceptibility to

cracking [75]. At certain locations observations of MGB widening, suggesting DDC

were made. The worst cases found are presented in Fig. 6.44 and Fig. 6.45. The

images show the slight widening of some MGBs and no cracking of the SGBs. In

weld metal, DDC always occurs at MGBs, [200] ruling out the possibility of the

cracks being solidification cracking (hot tearing). Compared to examples in literature

[75], and the rarity of these findings in all the material observed, these findings were

not considered severe. The DDC was mainly found near the fusion boundaries,

which is consistent with the literature suggesting high restraint levels and thermal

tensile stresses during cooling contributing to susceptibility to DDC cracking [75].

Fig. 6.44 Suspected DDC (arrow) at MGB near 316L fusion line. SA5 161


50 µm

328

Fig. 6.45 Suspected DDC in weld metal near SA508 fusion line with possible

recrystallization (arrow) SA5 161 Ammonium persulfate etch.

Local lack of fusion type defects were also found in weld SA5 in spite of

radiographical ASME IX acceptance. They are presented in Fig. 6.46 and Fig. 6.47.

The interpass LoF detected at the 18th

pass in weld SA5 was determined to be caused

by the excess oxidation in the previous pass caused by an equipment failure

interrupting the shielding gas flow mid weld. Sidewall LoF was found to have had its

origin in the square shape of the previous pass. The bottom of the defect was found

to coincidence with the corner. This can be seen in Fig. 6.46.

50 µm

329

Fig. 6.46 SA508 sidewall LoF at 2nd

to last pass. SA5 159 Nital etch.

Fig. 6.47 Interpass LoF in weld metal at pass no 18, SA5 161 Ammonium persulfate

etch.

Altogether two pores were found in all of the cross-sectional analysis of the SA

series of welds. The larger is presented in Fig. 6.48. Porosity was found acceptable

by the BS EN ISO 13919-1 [177] standard for laser beam welded joints. The pores

250 µm

250 µm

330

were also found smaller than the ASME IX [176] radiographical acceptance takes in

to consideration.

Fig. 6.48 Pore near fusion boundary of SA508. Note over-etched SA508. SA5 161


The 316L stainless steel microstructure was revealed with Oxalic acid electrolytical

etching. The base material grain structure is presented as Fig. 6.49.

200 µm

331

Fig. 6.49 316L base material. Austenitic grain structure with δ-ferrite stringers in

rolling direction. SA5 159.

The ferrite distribution in the 316L base material was found uneven: the ferrite

stringers being absent from the top and bottom 3 mm. It was also found, that at some

locations the grain size was exceptionally large, see Fig. 6.50.

500 µm

332

Fig. 6.50 Local cluster of large grain in 316L base material. SA5 159 Oxalic acid

etch.

The austenitic steel HAZ was known to be susceptible to HAZ hot cracking. This

was observed at the surface, where solidification stresses and welding distortions

were the highest, see Fig. 6.51. No other cracking was observed.

300 µm

333

Fig. 6.51 Hot crack in 316L HAZ near the top SA5. Note heavily over-etched weld

metal SA5 159 Oxalic acid etch.

Where the δ-ferrite stringers were present; the fusion line had experienced skeletal

ferrite formation, see Fig. 6.52 and Fig. 6.53.

100 µm

334

Fig. 6.52 Intergranular ferrite formation at fusion line. SA5 159 Oxalic acid etch.

Fig. 6.53 Skeletal δ-ferrite formation at fusion boundary. SA5 159 Oxalic acid etch.

Overall grain size in 316L HAZ was very similar to the base material. Grain growth

was suspected to be partly limited by the ferrite formation.

335

6.5.5 Residual Stress Analysis

Samples of the SA pedigree series of welds were tested for residual stresses using the

contour method. The welds were analysed in the as welded condition and the results

were not affected by post weld heat treatment. As the contour method required an

intact sample, the selection was based on the radiographical analysis. The samples

selected were SA2 and SA5.

The contour cut was performed using EDM, see Fig. 6.54. The sample was

restrained in a jig during the cutting. The jig was designed not to input any external

forces on the sample, but to hold it stationary. The round marks left by the restraint

used during the cutting are visible on the surface.

Fig. 6.54 Sample SA5 showing the contour cut

336

The non-trivial approach required for the dissimilar metal weld required many non-

standard considerations and work phases. The scripts originally created by Prime

[162] were developed further and manual modification of the data was required to

process the multi-material sample. The tri-metal composition of the weld required

the model to be partitioned in three sections with three different material elasticity

properties. The material properties were then manually edited in the resulting files.

The values for Poisson’s ratio (ν) and modulus of elasticity (E) were chosen from the

literature. Values from literature are displayed in Table 6.15.

Table 6.15 Poisson’s ratios and coefficients of elasticity for PCC materials

E (GPa) ν Source Notes

SA508 Gr3 Cl1 210 0.27 Feng [46] As used for CM

SA508 Gr3 Cl1 212 0.276 Hamelin [202] As used for CM

SA508 Gr3 Cl1 209 0.3 Vasileiou [203] As used for CM

SA508 206 0.294 Unpublished data *

Alloy 690 211 0.289 Chemically ≈ FM52 Special Metals [204]

Alloy 690 211 0.29 Chemically ≈ FM52 Special Metals [205]

FM52 213 0.3 Unpublished data *

316L 193 0.25 Aalco [206], NAS [71] *

* used in this study

The resulting contour maps are presented as Fig. 6.55 and Fig. 6.56. The welding

parameters of the welds differed slightly and an entirely matching result was not

expected. To analyse the effect of the cutting induced plasticity error different

cutting directions were used. This allowed comparing the gradient of the stresses

near the weld with an approaching and a receding cut. The areas of highest tensile

stresses, over 320 MPa can be seen drifted towards the start of the cut due to

plasticity. The areas of low tensile stresses below 160 MPa can be considered very

similar. This behaviour was found to be as expected by the literature [174] and

indicates a good correlation between the measurements.

337

Fig. 6.55 SA2 residual stress map

Fig. 6.56 SA5 residual stress map

A directly comparable study of residual stress formation in DMW was not found;

however, some comparisons were made. The distribution of the stress along the

thickness in the weld metal was found matching with findings by Kankanala

et al. 2011 [207], with a gradual increase from near zero level towards the top. They

studied ferritic austenitic dissimilar NG-GTAW using FM52 to for pipe form

samples of 40 mm wall thickness. The largest tensile hoop stress was 264 MPa. The

residual stress levels they reported were lower as the welds were in post weld heat

treated condition.

When compared to contour method analysis performed to pressuriser relief nozzles

from a cancelled power plant project by DeWald et al. [208] a similar concentration

338

of tensile stresses to the top or outer surface can be seen. These nozzles were made

by conventional MMA welding with Alloy 182 welding rod and were post weld heat

treated. The maximum tensile stresses were similarly inclined towards the stainless

steel and were approximately 300 MPa.

The maximum longitudinal tensile stresses in a GTAW DMW bead on plate study by

Kerr et al. [189] were found considerably higher. They reported tensile stresses up to

500 MPa using the contour method and neutron diffraction. The filler material was

Alloy82 and the base material 304L and the weld was investigated in as welded

condition.

A very good resolution of the decreasing tensile stress gradient (considering the

cutting direction) can be observed in both cases. The results from SA2 indicate a

steep gradient of the residual stress distribution at the SA508 material. The SA5 map

shows a gradual decrease of tensile stresses towards the 316L base material. The

sharp decrease of tensile stresses at the SA508 fusion line was attributed to the solid-

state phase transformation causing expansion of the LAS HAZ during cooling [209].

The expansion is due to the face centred cubic (FCC) crystal lattice austenite

transitioning to body centred cubic (BCC) allotropies, which causes an increase in

volume.

The sample SA5 was suspected to have slight corrosion due to water-based EDM

dielectric fluid. This was assumed to have contributed to an area lower tensile stress

in the results. This was later found to have cancelled during the averaging of the two

samples from the same cut. However, it was not entirely certain, as averaging only

filters out the expansion by corrosion if both samples have the same issue.

339

6.5.6 Longitudinal Tensile Strength Analysis

Samples SA5 and SA6 were used for tensile strength analysis. Altogether 33 test

coupons were fabricated and tested. Of these 15 were located longitudinally and 18

transverse across the weld. The samples used were sub-size and manufactured as per

Fig. 3.39.

Longitudinal tensile samples were cut from SA6. The aim was to investigate the

properties of the heat affected zones and compare the results to the base materials

and to assess the properties of the nickel alloy fusion zone.

The test coupons were located in both base materials, heat affected zones and weld

metal. The coupons were cut at three different depths representing the top, middle

and bottom of the thickness. This was done to investigate the multi-pass tempering

effect caused by the welding process. The HAZ samples were located 1 mm into the

base material measured from the fusion line. The test coupons were labelled and

numbered starting from CPL1 to CPL15. Locations of the coupons are presented in

Fig. 6.57. The sample numbering is clarified in Table 6.16. A photograph of the

samples is presented as Fig. 6.58. The heat affected zone samples were aligned to the

actual fusion line, which was slightly angled due to welding distortions.

340

Fig. 6.57 Locations of longitudinal tensile test coupons numbered CPL1 to CPL15

Table 6.16 Longitudinal tensile test coupon locations in sample SA6 by location

No: Location

CPL1,CPL2,CPL3 316L BM

4, 5, 6 316L HAZ

7, 8, 9 WM

10, 11, 12 SA508 HAZ

13, 14, 15 SA508 BM

341

Fig. 6.58 Longitudinal tensile test coupons; CPL1-3 316L BM, CPL4-6 316L HAZ,

CPL7-9 FM52 WM, CPL10-12 SA508 HAZ, CPL13-15 SA508BM

When comparing the locations of the heat affected zone and weld metal test coupons

to the hardness maps of the weld SA5 it is evident that also the webs of the samples

located near the top are within the multi-pass tempered region of the weld. Both

SA508 and 316L materials had a similar hardness at all three test depths.

The samples after testing are displayed in Fig. 6.59 and the results are displayed in

Table 6.17. The UTS (Ultimate Tensile Strength) values for the materials were

within the specifications, except the FM52 weld metal, in which the measured value

fell short by less than 1 %. The strain hardened 316L HAZ was found to have gone

through a slight increase in strength, but the change was much more prominent in

SA508 HAZ. The hardness evaluation indicated a significant hardening at the LAS

HAZ, which directly translated to a significant increase in tensile strength.

342

Fig. 6.59 Longitudinal tensile test coupons after testing

Table 6.17 Longitudinal tensile properties. Sample SA6

Location UTS

(MPa)

Elonga

tion

(mm)

Elongation at 36 mm gage

length

(%)

Material specification

UTS

(MPa)

316L BM 560 22.7 58.1 530 – 680

316L HAZ 610 14.6 37.4

WM 535 14.4 36.9 536 min

SA508 HAZ 910 6.0 15.4

SA508 BM 710 6.5 16.7 701 min

average of 3 coupons

The 316L base material was found to behave as expected producing a ductile

material stress-strain curve typical to austenitic steels, see Fig. 6.60. The stress strain

graph is characteristic of ductile materials with no distinctive transition from elastic

to plastic deformation. Non-uniform plastic deformation after the start of necking

was found short.

Strain rate used for the tests was 10 mm / min, except for the coupon CPL1 which

was tested at 2 mm / min. The slower strain rate led to a much larger elongation than

the faster rate due to strain hardening.

343

0 10 20 30 400

100

200

300

400

500

600

Str

ess (

MP

a)

Strain (mm)

CPL1

CPL2

CPL3

Fig. 6.60 316L BM stress-strain curves, test coupons CPL1 to 3

The 316L HAZ stress-strain curve is presented in Fig. 6.61. A more pronounced

yield point is present. Elongation is significantly reduced. Strength increases

averaging 610 MPa.

0 10 200

100

200

300

400

500

600

Str

ess (

MP

a)

Strain (mm)

CPL4

CPL5

CPL6

Fig. 6.61 316L HAZ stress-strain curves, test coupons CPL4 to 6

The ultimate tensile strength for Alloy 52 material is stated as 356 MPa, see Fig.

6.62. This was found slightly, although insignificantly, inferior in the actual weld

344

metal. The reasons for this may be the less than optimal heat treatment of the as

welded material and the inevitable dilution by the base materials.

0 10 200

100

200

300

400

500

600

Str

ess (

MP

a)

Strain (mm)

CPL7

CPL8

CPL9

Fig. 6.62 Filler Metal 52 weld metal stress-strain curves, test coupons CPL7 to 9

An area of significant interest in the DMW was the SA508 HAZ. This was found to

have gone through a substantial increase in strength, compare Fig. 6.63 and Fig.

6.64. This was supported by the fine grain microstructure found in metallography.

This also manifested itself as a substantial increase in hardness. Decrease of

plasticity as elongation was found small.

345

0 5 100

100

200

300

400

500

600

700

800

900

1000

Str

ess (

MP

a)

Strain (mm)

CPL10

CPL11

CPL12

Fig. 6.63 SA 508 Gr3 Cl2 HAZ stress-strain curves, test coupons CPL10 to 12

0 5 100

100

200

300

400

500

600

700

800

900

1000

Str

ess (

MP

a)

Strain (mm)

CPL13

CPL14

CPL15

Fig. 6.64 SA 508 Gr3 Cl2 BM stress-strain curves, test coupons CPL13 to 15

Overall the results of the longitudinal tests made at three different depths were not

significantly different. Only coupons 9, weld metal bottom, and 10, SA508 HAZ,

were found to have a slightly smaller elongation to their counterparts. There was no

indication of a change of properties by location of the test coupons through the

thickness. This is equivalent to a uniform tensile property distribution across the

depth of the weld.

346

6.5.7 Composite Tensile Analysis

In composite tensile tests, the aim was to establish the tensile strength of the weld as

a whole. Three coupons were cut at three different depths. Test coupons and their

locations are displayed in Fig. 6.65 and described in Table 6.18. The fractured

coupons are displayed in Fig. 6.66. The composite tests were compared to SA508

base material. The averaged results of the tests are displayed in Table 6.19.

Fig. 6.65 Charpy-V coupon and composite tensile test coupons CPT1-18 locations

in sample SA5

Table 6.18 Transverse tensile test locations

No: Location Depth

CPT1, CPT4, CPT7 Weld Top

2,5,8 Weld Mid

3,6,9 Weld Root

10,13,16 SA508 BM Top

11,14,17 SA508 BM Mid

12,15, 18 SA508 BM Root

347

Fig. 6.66 Composite (CPT1-9) and transverse 508 base material (CPT10-18) test

coupons, 316L steel top

Table 6.19 UTS and elongation of transverse tensile tests, SA5, average of 9

coupons

Location UTS

(MPa)

UTS

Max

UTS

Min

Elongation

(mm)

Elongation at 39 mm gage

length (%)

SA 508

BM

720 733 712 6.3 16.0

Composite 595 608 582 10.0 25.7

The variation of tensile properties was found to be small. The measured UTS of the

base material coupons were within 2.9 % of each other. For the composite samples,

the variation was 4.5 %. The fracture location was in the FM52 weld metal, except

samples 1 and 8, where the fracture occurred at the 316L stainless steel base

material. This variation can be expected as the nominal UTS of these materials are

similar.

The multi-pass tempering effect was investigated by comparing the top, middle and

bottom samples. The results are presented in Table 6.20. The variation in the

measurements was found too small to allow for judgements of benefits or drawbacks

of the multi-pass tempering effect. This agrees well with the hardness test results.

348

Table 6.20 Variation of transverse tensile properties according to location of the

sample, SA5

Location UTS

(MPa)

Elongation

(mm)

Elongation at 39 mm gage length

(%)

Composite Top 595 11.4 29.1

Composite Middle 600 11.0 28.1

Composite Bottom 585 11.1 28.5

BM Top 725 6.4 16.5

BM Middle 715 6.2 15.9

BM bottom 720 6.1 15.7

Average of 3 coupons

The stress-strain graphs for the composite tensile tests are presented in Fig. 6.67, Fig.

6.68 and Fig. 6.69.

0 5 10 150

100

200

300

400

500

600 CPT1

CPT2

CPT3

Str

ess (

MP

a)

Strain (mm)

Fig. 6.67 Composite test stress-strain curves, test coupons CPT1 to 3.

349

0 5 10 150

100

200

300

400

500

600 CPT4

CPT5

CPT6

Str

ess (

MP

a)

Strain (mm)


0 5 10 150

100

200

300

400

500

600

Str

ess (

MP

a)

Strain (mm)

CPT7

CPT8

CPT9


For comparison, the transverse direction stress-strain curve of the SA508 base

material is provided as Fig. 6.70. The 9 BM tests conducted display very consistent

results and are in accordance with the previous longitudinal analysis.

350

0 5 10 150

100

200

300

400

500

600

700

800

Str

ess (

MP

a)

Strain (mm)

CPT10

CPT11

CPT12

CPT13

CPT14

CPT15

CPT16

CPT17

CPT18

Fig. 6.70 Transverse 508 BM test stress-strain curves, test coupons CPT10 to 18.

6.5.8 Digital Image Correlation Analysis

DIC analysis was conducted on the tensile test samples for strain localisation

mapping for the composite tests. Most composite samples failed at the weld metal,

which was expected as the FM52 nickel alloy had the lowest UTS of the materials

involved refer to longitudinal test results in Table 6.17. During the composite tensile

testing 7 of the 9 tested samples failed at the middle of the weld metal.

The minimum stated yield strength for both FM52 and 316L is 240 MPa. In the

coupons failing at the nickel weld metal a considerable amount of strain was

observed also at the stainless steel. In all these cases the first area to yield was the

stainless steel, which then work hardened and caused the ultimate failure to occur at

the weld metal. Fig. 6.71 displays this behaviour. In Fig. 6.71 A, the first indications

of strain are observed. At first, the strain is entirely in the 316L base material. The

next stage is when the weld metal starts yielding more than the stainless, see Fig.

6.71 B. These stages fluctuated, both materials straining in turn. Fig. 6.71 C is the

point just before breakage when the yielding has transferred entirely to the weld

351

metal. The 316L part of the coupon was found to neck in width by 0.90 mm (18 %)

without failure. No measurable necking was observed in SA508 portion of the

coupons, which was due to the yield strength of SA508 (600 MPa) being higher than

the UTS of the austenitic materials.


Stainless steel at bottom.

Samples 1 and 8 failed at the 316L stainless steel. Sample 1 was located near the top

and subjected to a higher than average heat input due to the capping passes being

welded at the top. The sample 8 was located in the middle of the thickness and was

representative of an average heat cycle. The small number of samples did not allow

for further conclusions to be made of the origin of this behaviour. A three-stage

description of the strain history of Sample 8 is presented in Fig. 6.72. The stainless

steel yields first Fig. 6.72 A. This is followed by a stage where both the 316L and

weld metal strain, Fig. 6.72 B. Towards the end of the test, the already heavily

deformed weld metal hardens and seizes to yield, Fig. 6.72 C. A detail showing the

considerable deformation of the weld metal is displayed as Fig. 6.73. The weld metal

was found necked by 0.45 mm or 9 % in width.

A B C

352


Stainless steel at bottom.

Fig. 6.73 Deformation of the fusion zone in SA5, coupon 8, arrow

6.5.9 Impact Toughness Analysis of SA508 Base Material

The HAZ of ferritic steel weld is susceptible to embrittlement via phase

transformation hardening. Therefore the emphasis of the fracture toughness tests was

on examining the SA508 and its heat affected zone. Two sets of test coupons were

EDM machined to allow comparison between the SA508 HAZ and the SA508 base

material. The test coupons were standard size according to ASTM E23 [199] and

ISO 148-7:2016 [178]. First set of samples, numbered from 1 to 9 were machined

from the base material of the sample. The HAZ samples were aligned perpendicular

to the LAS fusion line and the V-notches were aligned 1 mm into the BM from the

fusion line. This was to provide the crack propagation start in the coarse grain zone

A B C

353

of the HAZ. The samples were extracted from sample SA5 B, the end half of the

contour cut sample. The actual sample locations are described above in Fig. 6.65.

The base material samples were aligned parallel to the plate. The HAZ samples were

tilted slightly to match the fusion line. The v-notches were aligned 1 mm from the

fusion line into the HAZ of the SA508 base material.

To heat and cool the samples a heat plate and liquid nitrogen were used. The samples

to be tested at other than room temperature were either heated on a hot plate or

cooled by immersing in liquid nitrogen. Thermocouples were spot welded to the

samples in order to monitor the temperature. The thermocouples were located to the

side of the sample as close to the V-notch as possible. It was known that the

thermocouples were only able to indicate the temperature of the surface of the

sample. To minimise the heat gradient within the sample, the samples were over

cooled or overheated by 20°C. This allowed approximately 2 minutes for the

temperature to soak. It also allowed time for placing the sample on the anvil. All

temperatures indicated are surface temperature readings and the procedure above

means that the core sample temperatures are lower than the temperatures indicated.

The results of the SA508 Gr3 Cl2 base material impact test are displayed in Fig.

6.74. Nine impact tests were conducted. The ductile to brittle transformation

temperature and the upper and lower shelf energies are displayed, except for the two

test points at 15°C, which were considered as scatter.

354

-30 -20 -10 0 10 20 3030

40

50

60

70

80

90

100

I

H

G

F

E

DC

B

Impact

Energ

y (

J)

Test Temperature (°C)

A

Fig. 6.74 SA508 base material Charpy-V impact energy, J

The change in the proportion of ductile shear fracture can be seen in Fig. 6.75.

Samples are displayed in rising order of temperature. Samples A to D located at the

lower shelf of the ductile to brittle transition curve. The fracture mode is almost

entirely brittle with a characteristic flat surface and very little necking of the sample.

Sample E shows a significant increase in the ductile fraction of the fracture area as it

approaches the upper shelf of the transition diagram. Ductile fracture is

characteristically more uneven and the necking observed in samples is significant.

Samples F and G are located on the upper shelf of energy and considered ductile.

355

Fig. 6.75 SA508 Base material fracture surfaces. A – D, G and H brittle fracture, E

transition and I ductile facture

The scattered nature of the Charpy test results is observed in samples G and H,

which were found brittle at near room temperature, where adjacent samples F and I

A B

I

F E

D C

G H

5 mm

356

were found ductile. The reasons for that were concluded to be involved with the

suspected material flaw in sample H and the temperature measurement inability to

represent the core temperature as the thermocouple location was on the surface of the

sample. Impact energies and the proportion of ductile fracture are displayed in Table

6.21.

Table 6.21 Impact energies and proportion of ductile fracture in SA508 Charpy-V

tests

Temp

(°C)

Impact energy

(J)

Proportion of

ductile fracture

(%)

Notes

A -30 37 2.1

B -20 42 2.4

C 0 52 4.4

D +5 54 5.3

E +7.5 73 14.7

F +10 83 18.4

G +15 61 2.2

H +15 43 4.3 suspected inclusion / discontinuity

I +22.5 92 29.1

For comparison, ASME A508/A508M – 17 specification for Gr3 Cl2 [67] is

minimum 48 J at 21°C as average of three, minimum single measurement 41 J. The

material certificate specifies values of three samples of the SA508 material used,

which were 46 J, 66 J and 84 J, at room temperature, see material certificate in

Appendix IV.

The fracture mode was further examined using SEM. In the SA508 base material, the

ductile and brittle fracture areas had a clear distinction. The ductile area had the

typical morphology, see Fig. 6.76, with spherical dimples caused by the tensile load

near the Charpy notch. The brittle fracture, Fig. 6.77, area had the characteristic

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morphology of the cleavage mechanism. The SEM analysis confirmed the findings

of the optical microscopy.

Fig. 6.76 Ductile fracture area in SA508 base material, 400x, 1000x and 5000x

magnification, sample F

358

Fig. 6.77 Brittle fracture area in SA508 base material, 400x, 1000x and 5000x

magnification, sample F

359

6.5.10 Impact Toughness Analysis of SA508 Heat Affected Zone

The HAZ of the SA508 base material was tested for embrittlement. The notches of

the Charpy-V test coupons were located 1 mm from the fusion line as per the

U.S. NRC code [159]. The tests were conducted with the same geometry and

procedure as the SA508 base material tests, temperatures ranging from -120°C to

+150°C. The results are displayed in Fig. 6.78.

-100 -50 0 50 100 150

100

120

140

160

180

200

Imp

act

Ene

rgy (

J)

Test Temperature (°C)

Fig. 6.78 SA508 HAZ Charpy-V impact energy, J

The resulting fracture surfaces are displayed in Fig. 6.79. All surfaces were

considerably more uneven than in the base material tests. This suggested a ductile

fracture. There was no clear distinction between ductile and brittle areas as in the

previous tests. According to optical microscopy, the brittle areas were found to be

scattered in the surface.

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Fig. 6.79 SA508 HAZ fracture surfaces. A and B brittle, C to I ductile.

Judging the fracture mode was found less obvious in the HAZ tests than the base

material. Samples A and B which were tested at the lowest temperatures had

significantly lower impact energy than the rest. The difference between B and C was

37.6 %. Necking of the surface was also considerable in C – I, which were tested at -

60°C and above. One significant change was also that the sample halves did not

separate in the higher impact energy tests, see Table 6.22. SEM was used to analyse

the fracture mechanisms in more detail.

A

D

C B

E F

G H I

5 mm

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Table 6.22 Impact energies and proportion of ductile fracture in SA508 HAZ

Charpy-V tests

Sample A B C D E F G H I

Temperature (°C) -120 -100 -60 -40 -20 0 25.3 100 150

Impact energy (J) 95 98 157 146 197 155 160 174 152

Separation of samples Yes Yes No No No No No No No

To confirm the results of the optical analysis, the SA508 HAZ fracture surfaces were

analysed by SEM. The high-temperature samples are represented by images from the

room temperature Sample G, Fig. 6.80. The fracture was found to be entirely ductile.

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Fig. 6.80 SEM images of ductile fracture. 400x, 1000x and 5000x magnification,

Sample G

363

Sample C, partially presented in Fig. 6.81, was the lowest temperature sample with

high impact energy. The morphologies are much less obvious than in the base

material tests. The increase of brittle fracture in the area below the red line was

confirmed by higher magnification SEM. The images are presented in Fig. 6.82 and

Fig. 6.83.

Fig. 6.81 SEM image of Sample C, low magnification

364

Fig. 6.82 SEM images of Sample C. 400x, 1000x and 5000x magnification, ductile

area

365

Fig. 6.83 SEM images of Sample C. 400x, 1000x and 5000x magnification, brittle

area

366

Results for the brittle sample A is presented in Fig. 6.84. The fracture surface

morphologies were more varied, with less clear distinction of ductile or brittle areas.

Fig. 6.84 SEM images of HAZ Sample A, three different locations A, B and C. 400x,

1000x and 5000x magnification.

The SEM for sample A, location A shows some “cup and cone” morphology due to

ductile tensile loading but also the “river pattern” of a brittle structure. The location

D has indications of parabolic dimples, an indication of shear loading. However, the

majority of the fracture surface of sample A was like the location G, a fine brittle

structure.

The main outcome was that the coldest samples A (-120°C) and B (-100°C) with

impact energies below 100 J were mainly brittle in behaviour. The samples from C

A1 A2 A3

B1

B2 B3

C1 C2 C3

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to I were mostly ductile. This indicated the ductile to brittle transition to occur

between -100 and -60°C.

Overall the impact energies of the SA508 HAZ were consistently high at test

temperatures including and above -60°C. Below this temperature, the fracture

mechanism was confirmed brittle using SEM. The average energy on what could be

considered the upper shelf was 163 J, which is above the U.S. NRC 10 CFR 50

Appendix G requirement of USE (102 J) [159] by a fair margin.

6.6 Discussion

Achieving the ASME IX radiographical acceptance for the welds proved

challenging. The welding process was found sensitive to alignment and laser spot

size related fusion issues. Lack of fusion defects were the cause for all rejected

welds. However, an acceptable quality was reached in one full weld and in several

segments of welds.

Lack of fusion was contributed partly to the undercut phenomenon which was

observed to affect mostly the stainless steel BM and partly to the accumulation of a

layer of oxides altering the wetting properties of the molten pool. The undercut

phenomenon was considered to be affected by the lower heat diffusivity of the

stainless steel compared to the LAS, the mechanism being that the stainless steel

surface would overheat to the melting point rather than dissipating the heat into the

cool mass of the material causing undercut. The effect of the oxides floating on the

surface of the melt pool was seen in the LII imagery Fig. 6.20 and visual inspection

of incomplete welds, Fig. 6.18.

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The macrographical analysis showed that when the undercutting of 316L was

controlled, the welds formed symmetrical and the resulting fusion lines were near

parallel. The 316L was found to contract locally, the top of the sample being drawn

into the weld, causing considerable distortion of the top profile. The fusion issues

found in radiography became evident in the macrographical analysis. They were

found occurring only after 10 – 13 passes had been completed. Often they were

found after comments of excess oxidation being observed on the weld bead during

welding. This evidence along with the weld video monitoring observations led to a

conclusion that the oxide accumulation is one major source of fusion issues.

An area of high hardness in SA508 HAZ caused by the last pass was detected in

hardness analysis. The hardness was generated by phase transformation hardening.

The welding heat input had risen the base material temperature up to austenisation

temperature and the material subsequently had quenched by the heat sink of the

material itself. This was only local, as the multi-pass tempering effect reduced the

hardness significantly after only a single additional heat cycle. The reduction of

hardness was from 480 HV to below 350 HV. The hardness profile of the weld

below the last HAZ was found uniform and continuous. This shows the potent

capability of the multi-pass tempering in restoring toughness of the SA508 HAZ.

Hardness testing results in the 316L were concordant with the observation of

significant strain as noticeable strain hardening was found in the material. Strain

hardening due to the stresses caused by the multi-pass welding was also found in the

FM52 weld metal. This was typical behaviour for the austenitic weld metals used.

Microstructural analysis confirmed the existence of a martensitic HAZ of the last

bead in SA508, which was expected to form in a LAS due to the rapid cooling phase

369

of the welding heat-cycle. Effective multi-pass tempering was found to improve the

grain structures. The effect of the grain refining caused by the re-heating of the

martensitic HAZ just over the austenisation temperature, dissolving the existing

phases, but not allowing time for grain growth of the austenite, was observed. A

beneficial small grain size was found in multi-pass heat treated parts of the SA508

HAZ, beginning at the once heat treated second to last pass.

In the weld metal, a very characteristic primary austenitic structure was observed.

The grain growth from the 316L fusion line and the previous weld beads was

epitaxial, developing to competitive growth later on. SSGB, SGB and MGB grain

boundaries were observed. This is typical for the materials used [75, 122]. MGBs

were not found at every weld pass. Some very small suspected DDC cracks were

found in two of the 20 weld passes. No other types of cracks were observed. Other

flaws found were interpass and sidewall lack of fusion and occasional minor pores.

Generally, the welds were solid for such a highly restrained weld. This is an

indication of the potential benefits of NGLW in this type of weld geometries.

Very little grain growth was found in the 316L HAZ, this was contributed to the

relatively rapid heat cycle not allowing time for the growth. The grain structure and

size were found characteristic to an austenitic stainless steel weld. Small amounts of

δ-ferrite had formed on fusion line. The 316L base material was found relatively

non-uniform with uneven distribution of ferrite stringers and some areas of

exceptionally large grain. A solidification crack was found at the top surface in the

HAZ. These irregularities were given no emphasis as in a real application such issues

would be ironed out.

370

The residual stress analysis the two slightly differently executed measurements were

found concordant, giving validity to the findings. The effect of cutting-induced

plasticity was found to distort the high-scale stress map towards the beginning of the

contour cut, as was expected by the literature. Mid-range stresses were not found

affected. Resolution of the maps was concluded to best at a decreasing gradient of

stresses.

In longitudinal tensile analysis, five resulting zones were analysed. All zones were

found matching the nominal material properties. The scatter was small. The multi-

pass tempering effect on the tensile properties was found to be similar irrespective of

the location of the sample in depth; top, mid or root. This was concluded to be due to

the beneficial small grain size being generated already after a single multi-pass

tempering cycle, as the top coupon was cut just below the untempered top pass.

In composite tensile testing, the strength of the weld was proven similar or better

than the properties of the weakest metals used. In most cases, the breakage occurred

in the weld metal, but only after a significant reduction of area in the 316L base

material. The two other breakages occurred in the stainless steel base material. The

UTS of the composite tests was found 595 MPa, which exceeded that of the

longitudinal tests in the FM52 weld metal and 316L base material and material

specifications. Overall this indicates a sound weld.

In the DIC analysis of the composite tensile tests, the deformation of 316L BM and

FM52 FZ was found to alternate between the two materials before breakage. This

was to be expected as in the longitudinal tensile testing the UTS of the FM52 FZ,

535 MPa, was close to the 316L, 560 MPa. As the UTS of the 316L was higher, the

breakage occurred most often at the FM52 fusion zone. The difference in UTS came

371

apparent as considerable strain hardening (during welding) was observed in the

316L HAZ and it showed very little strain during the tensile testing. The SA508

remained practically unchanged as it has the highest yield strength of the materials

involved, results indicating it being higher than the UTS of the other metals.

The results for the Charpy-V analysis of the SA508 Gr3 Cl2 base material provide a

good correlation between observations of the fracture surface and the impact energy

results. The plotted graph shows a distinct ductile to brittle transition. The

temperatures indicated are surface temperatures measured from a warming sample.

This means that the core temperature was lower than the surface temperature stated

in the results. This was compensated by avoiding overcooling the samples and also

allowing the temperature to soak in for a time just before the test. The impact

toughness properties of the SA508 Gr3 Cl2 base material tested was found to be in

good agreement with the material certificate provided with the material and within

the specifications of the A508/A508M – 17 standard.

According to the Charpy-V impact test and the material certificate, Appendix IV, the

SA508 base material upper shelf impact energy varied between 42 and 92 J. Post

welding the impact energy was found increased, the absorbed USE varied from 146 J

to 197 J. This is a good indication of the sound uniform properties NGLW and the

multi-pass heat treatment inherently involved in it produces and a direct consequence

of the fine grain size observed in other analysis.

The impact toughness requirement of 102 J by the U.S. NRC for the SA508 HAZ

was exceeded in tests even without post weld heat treatment. Although the number

of samples was low, relatively little scatter was observed. A clear ductile to brittle

transition zone was established and confirmed using optical and SEM microscopy.

372

Overall the effect of multi-pass tempering was found beneficial, as the tempered

portion of the weld and HAZs had good properties, often exceeding those of the base

materials. Although the first passes were subjected to a larger amount of heat

treatment cycles than the ones nearer to the top, no difference was found between

samples cut from the top, middle or bottom of the weld thickness in any of the

analysis. The top sample was cut at a pass which was heat treated only once. This

shows the uniformity of the resulting weld.

6.7 Conclusions

Answers to the research questions set for this stage were found through a detailed

analysis of the welding process and the resulting welds. The following summarises

the key findings and conclusions in Stage III DMW:

The lack of fusion defects were the main challenge in welding. Lack of fusion was

detected using radiography and analysed further using macrography. This was in

agreement with both literature and previous work in this study. In practical work, it

was found that the SA508 material was found to be more prone to LoF defects than

the S275 previously used, but the development of welding practices was able to

counter the challenges.

Weld quality complying with the industry-standard ASME IX radiographical

acceptance was reached. The weld SA5 was accepted by the inspection for its entire

length. Several other partial sections were also acceptable, issues often concentrating

at the ends of the welds. In cross-sectioning the issues were found to be concentrated

towards the top of the weld, similarly to the findings in previous work in this study.

373

The U.S. NRC impact energy requirement for the ferritic steel HAZ was exceeded in

as welded condition. The multi-pass tempered HAZ was found to comply with the

requirements without PWHT. This suggests that the PWHT is not required for

achieving a sufficient level of ductility for safe operation.

The multi-pass tempering was proven effective in restoring toughness in the LAS

HAZ. The higher hardenability of SA508, compared to S275, produced a maximum

hardness of 480 HV in the HAZ which was tempered to 350 HV suggesting a

significant improvement in ductility. The multi-pass tempering was found to temper

the martensitic microstructure very effectively. Using detailed hardness mapping,

only a single heat cycle was found capable of reducing the hardness of the

martensitic SA508 HAZ to a constant level. The tempering of the microstructure was

demonstrated using microscopy. The multi-pass tempering effect was confirmed to

restore the toughness by using Charpy-V pendulum impact testing.

The multi-pass tempering effect to the SA508 HAZ toughness was shown uniform.

Only the untempered HAZ of the last pass had a potentially brittle microstructure

and hardness. The rest of the weld had much lower hardness and little variation. The

hardness mapping and microstructural analysis indicate that the ductility should be

independent of the depth of the sample. The results agree with previous work in this

study.

The microstructural variation in SA508 HAZ was very small after three heat cycles.

In microstructural analysis, no significant variation in the microstructure was

detected deeper in the weld. This is similar to the Stage II findings, regardless of the

very different hardenability of the SA508 used. The variation of microstructure due

to the multi-pass tempering in the weld metal and the 316L was small.

374

The mechanical properties of the weld were found uniform through the thickness.

The uniformity of the mechanical properties was measured cutting tensile strength

and impact energy test coupons from three depths in the welds. The mechanical

properties were found independent of the location of which the sample was cut. This

further reinforces the notion of a uniform multi-pass tempering effect.

Unexpectedly few indications of cracking or porosity were observed. Very small

signs of DDC were observed in the FM52 weld metal, but no solidification or

liquation cracking were detected. The rarity of this issue was considered as a benefit

of the NGLW process as the FM52 is known for high susceptibility to DDC in high

restraint welding conditions, such as thick section welding.

The residual stress distribution was found analogous to literature. Using the contour

method the distribution of the residual stresses was found similar to a comparable

MMA weld, but the maximum values measured in this study were higher. This was

concluded to be due to the lack of PWHT. This poses an issue, as the higher residual

stress levels make the weld more susceptible to PWSCC. This indicates the need for

further study on the actual EAC properties if the weld is to be used without PWHT.

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7 General Discussion on Dissimilar Metal NGLW

7.1 Introduction

The welding in this study was conducted in three distinct steps. Different materials

were used at each stage. The materials had very different properties resulting in

equally different results. However, many of the phenomena discovered were

universal and found regardless of the materials used. This chapter discusses the

commonalities and differences in NGLW of different material combinations.

7.2 Process Characteristics

The concept of NGLW is based on trying to reach the narrowest possible welding

groove. This creates a unique environment for welding. All the equipment has to be

concentrated into exceptionally small confines. This generates new types of

challenges compared to conventional welding processes. The following was found

characteristical to NGLW in this study.

Visibility into the process was very limited due to the narrow groove and the various

nozzles involved. Monitoring the melt pool without a coaxial camera system was

found virtually impossible after the trailing shielding gas nozzles were added to the

welding head. Also achieving adequate accuracy in the programming of the weld

path pre welding was inaccurate until the introduction of the camera. Using the

coaxial camera system was used to increase the accuracy of weld programming to

match the precision of the robotic system.

376

Laser illumination imaging was used at the beginning of the study. The resulting

imagery was free of all disturbances and found highly valuable for the process

development. This system was not compatible with the coaxial camera optics as the

band-pass filter could not be applied; hence a much inferior ambient light imaging

had to be used.

Significant reduction of welding passes and filler material required over

conventional GTAW compared to literature [35, 37] was reached. Considerable filler

material savings were reached also over NG-GTAW. This also leads to secondary

benefits such as a reduction in inspection steps. Material and time savings

calculations indicate the capability of over 90 % reduction of welding time in

comparison to GTAW of the reactor pressure vessel nozzle at 80 mm thickness.

Results show over 60 % reduction on the filler material required compared to NG-

GTAW.

7.3 Geometrical Tolerances

Although less stringent than in keyhole laser welding, the narrow gap arrangement

was known to have relatively strict fit and alignment requirements as Jokinen [17,

44] had reported. It was understood that at an industrial scale the tolerances of

welding groove manufacturing would be much inferior to the laboratory setup.

The approach to welding groove design was to design a geometry which would keep

the width of the bottom of the welding groove constant for each pass. The aim was to

minimise the need to adjust the laser spot size between passes.

Some of the tests, see Chapter 5.4.3, in this study indicated a three-stage groove

contraction behaviour, which might have benefitted from a three bevel groove shape.

377

However, this was not seen in SA508 welds, see Chapter 6.4.2, where the

contraction behaviour was found near-linear until tapering off near the top of the

weld.

Due to fabrication inaccuracies in tag welding, the actual width of the welding

groove varied ±0.20 mm. Using 3D scanning the groove angle was also observed to

contract 0.25° more at the middle of the weld than at the ends. This together with the

welding distortions which also varied case by case, led to less than optimal spot size

fit. Considering the apparent simplicity of the experimental setup, it was concluded

that the groove shape design could never quite match the real distortions accurately

enough to provide trouble-free welding using fixed parameters. In all cases, some

tailoring of the parameters was required. This had to be done in between the passes

as the welding system would not allow otherwise. It was concluded that a welding

robot system designed for conventional welding, capable of features like in-process

power and alignment (weaving, spot size etc.) adjustment, should be used to improve

the results and alleviate the stringent geometrical tolerances.

The tolerances for the alignment of the weld and filler wire were found similar

throughout the study. The materials did not have a significant effect on the required

precision on the wire alignment.

7.4 Oxidation of the Weld Metal

The similar metal NGLW was conducted by flooding the welding groove with one or

two large-diameter slow-flowing stationary nozzles at the ends of the sample. This

was found to produce adequate results, although the accumulation of oxides over

multiple passes was observed to slightly increase the occurrence of flaws, like lack

of fusion.

378

The FM52 filler relies heavily on Al and Ti to float oxygen out of the weld pool as

oxides [87, 89, 128]. This was seen as the accumulation of an oxide layer to the weld

bead, which was one of the major problems encountered in this study.

One major difference in the similar metal welds over the dissimilar welds was that

the deposition rate was higher and therefore the number of passes required to fill the

up to 40 mm welds was lower, which reduced the accumulation of oxides and

reduced the issue.

7.4.1 Remedies for Oxidation Issues

The oxidation was reduced by improving the gas shielding and introducing a regular

mechanical removal of the oxide layer. One way suggested to address the oxidation

propensity was that the amount of FM52 could be kept to a minimum while most of

the weld should be filled using mechanically similar FM82. The thickness of the

FM52 being decided on the corrosion properties. Changing the FM52 to more

modern variants like FM52M [87] or FM52MSS [89], which are developed specially

to reduce the floating oxides, is recommended.

7.5 Undercutting of Stainless Steel

Undercutting was first observed in the Stage I stainless steel similar metal welds.

The small amount of symmetrical undercut was found to have little adverse effect,

causing only negligible widening of the fusion zone. Signs of undercut were seen in

all stainless steel similar metal welds, but only re-melting would cause measurable

undercut. A cross-sectional sample of undercut is presented as Fig. 7.1.

379

Fig. 7.1 Near symmetrical undercut (arrow) due to re-melting of a pass in an

incomplete stainless steel weld. Note also overheated top corners of the

welding groove. Oxalic acid etch.

The asymmetrical undercutting experienced in dissimilar metal welds was found

much more detrimental. Some amount of undercutting of the stainless steel was

observed in most of the dissimilar metal welds. This was determined to be caused by

local overheating and excessive melting of the stainless steel due to the smaller heat

diffusivity compared to the ferritic steels, Chapter 6.6. The undercutting

phenomenon was found to be self-sustaining by shaping the weld geometry in a way

that the laser reflections were directed more towards the already overheating

stainless steel.

Despite the slight asymmetry generated by the undercut phenomenon, significant

geometrical problems in the weld did not occur until the last 3 or 4 passes. The

overhang above the undercut would collapse at this stage and that would create a

much wider, irregularly shaped welding groove. This would happen erratically, see

Fig. 7.2, causing unsolvable problems to the welding of the final passes. In the end,

10 mm

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the controlling of the undercut was the most critical factor to control in order to

create a successful weld.

Fig. 7.2 Local widening of the welding groove caused by irregular collapse of

asymmetric undercut in the stainless steel, sample DS4

The consequences of the collapse were countered by offsetting the laser and

increasing the spot size accordingly to guarantee fusion to the ferritic steel. However,

the collapse was always intermittent and it would generate a groove with a very

irregular shape. This led to issues with weld symmetry, fusion, dilution and gas

shielding. Welds with undercut collapse were distorted and often severely faulty at

the top.

Due to the deep and narrow welding groove, the asymmetric undercut was difficult

to detect. The growth of the undercut was monitored by examining the geometry

after each pass visually. To prevent the undercut several approaches were

investigated including: controlling the interpass temperature, reduction of laser

power, increasing wire feed rate, changing the alignment of the laser spot inside the

welding groove, etc. Careful control of laser power was found to be the most

effective solution.

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7.6 Mechanisms Leading to Formation of Lack of Fusion

The repeatability of the NGLW was found challenging mainly due to the well-known

issues related to fusion [18, 41, 46]. Two types of lack of fusion were detected,

interpass LoF and LoF to the sidewalls. The fusion issues were found to be partly to

be caused by the challenging weldability of the FM52 filler material [84, 129] and

partly due to the tendency of 316L to generate undercut.

Lack of fusion, especially to the stainless steel, was considered the main challenge in

the dissimilar metal welding, being the cause of all rejections in the ASME IX X-ray

inspection. Sidewall lack of fusion was found to occur due to two phenomena:

oxidation and undercut. Firstly, the layer of oxide slag floating on the melt pool was

observed to distort the shape of the surface by preventing the capillary effect from

wetting the sidewalls. Secondly, the stainless steel undercut prevented the laser beam

from heating the fusion boundary sufficiently. Interpass lack of fusion was found to

form via lack of penetration due to low laser power the previous oxide layer being

trapped between passes without melting.

7.6.1 LoF Due to Oxidation

Alloy 52 was known to be prone to issues with oxidation [87, 89, 128]. The Al and

Ti in the alloy act as de-oxidisers generating light floating oxides [64]. Generally, the

oxides floated to the top of the weld pool leaving no inclusions in the welds. The

problem was that the layer of oxides floating on the weld pool accumulated pass

after pass. In the first passes of the weld, there were seldom difficulties with

oxidation, but after a certain amount of passes, depending on the gas shielding setup,

the problems became obvious. Using LII imaging videos the layer of oxides was

found to force the surface of the melt pool flat, see example Fig. 7.3. This prevented

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the capillary effect from pulling the liquid metal upwards to the side walls to form a

smooth concave weld bead, forming a square flat shape with sharp corners instead,

see Fig. 7.4.

Fig. 7.3 LII image of DS2, pass 11. The weld bead generated is flat. Note heavy

oxidation.

Fig. 7.4 Desirable and undesirable bead shapes

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The following pass was unable to melt the corners of this flat, square shape reliably

to achieve good fusion and uniform penetration to the previous bead, Fig. 7.5.

Instead, the shape generated a high risk of lack of fusion, see Fig. 7.6.

Fig. 7.5 Optimal penetration between passes and reliable fusion

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Fig. 7.6 Consequences of inadequate wetting increasing the risk of lack of fusion

This type of lack of fusion was a common challenge and remedies were developed.

The generation of oxide slag was minimised by improving gas shielding. It was

supplemented by manual wire brushing after each pass to reduce the accumulation of

oxides. Laser power was kept as high as possible to keep the penetration at a

maximum level. Together these methods were sufficient to produce for 40 mm

thickness welds fulfilling the requirements of ASME IX radiographical acceptance.

7.6.2 LoF Due to Undercut

The undercut induced lack of fusion defects formed through a mechanism of

asymmetric widening of the bottom of the welding groove towards the stainless steel

base material. Undercutting would put part of the bottom out of reach of the laser

beam, effectively generating a shadow. The shaded area would not receive any direct

heat from the laser beam, and the viscous FM52 weld pool would not be capable to

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flow there to heat and wet the sidewall. This prevented normal wetting of the

stainless steel sidewall. In many cases, the resulting bead surface would be convex

which would qualify as lack of fusion in itself, see Fig. 7.7.

Fig. 7.7 Undercut induced lack of fusion principle and two possible bead shapes

A flat bead shape would cause problems during the following pass. The corner

would not be in reach of reliable penetration of the subsequent pass and it would,

therefore, be left in the weld unmelted as a flaw, see Fig. 7.8.

316 L SA508

386

Fig. 7.8 The effect of undercut induced lack of fusion to the following welding pass

To prevent this type of lack of fusion careful monitoring of the undercut

development was used. If the undercut was not developed too wide, reduction in

laser power was found an effective way of reducing the undercut for the subsequent

passes.

7.6.3 Interpass LoF

Interpass lack of fusion was found to form when the laser power was too low,

confirming the findings in the literature [41, 46]. The weld SA2 was welded with a

power of 2500 to 3000 W. The sidewalls wetted very well and the appearance of the

beads was generally good. This weld, however, was riddled with interpass LoF.

To reduce the susceptibility to interpass LoF a higher laser power was applied to

increase penetration. The weld SA5 was filled with laser power of 3500 W. It was

free of interpass LoF. The power was however limited by the increase in the

316 L SA508

387

asymmetric undercut of the stainless steel, which tended to lead detrimental

distortion of the weld geometry. A careful balance had to be maintained.

One contributing mechanism was the oxide layer on top of the previous bead

stopping the weld penetrating the surface. The hard oxide layer was found acting

effectively as an insulator. In many cases, the interpass lack of fusion was found

corresponding to observations of excess oxidation on the previous bead. The

underlying mechanism was considered to be the less turbulent melt pool of the low

power passes not being able to convey enough heat to dissolve the oxidised surface.

The sluggish nature of the FM52 melt pool [84] was suspected to contribute to this

issue.

7.6.4 Location of Lack of Fusion

The issues with lack of fusion were seen to concentrate to the ends of the samples in

all test stages. Mechanisms leading to this were considered. The smaller contraction

of the welding groove at the ends, see Fig. 4.6, was making the groove relatively

wider nearer the ends. This effectively made the laser spot smaller than optimal

leading to lack of sidewall fusion. Simultaneously the thickness of the bead

previously deposited was relatively lower causing the wire to reach the melt pool at

the top range of the tolerance range, sometimes overshooting the pool and being

melted directly by the laser beam, being deposited discontinuously as droplets, refer

to Fig. 7.13. This was seen to increase oxidation and unevenness of the bead, both of

which were recognised as promoters of lack of fusion.

These issues were challenging to combat using the fixed trajectory robotic system.

The linear trajectory of the robot had to be a compromise between the lower wider

ends of the weld and the higher narrower middle part. The laser spot size to groove

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width ratio could be optimised if a more flexible robotic system was used. It would

also align the wire correctly, keeping the interaction point with the melt pool

constant.

7.7 Laser Beam Reflections from the Weld Pool Surface

Although literature states that the absorption of a laser beam to metals generally

increases very close to 100% as the temperature rises to temperatures required for

welding [40] a considerable amount of reflections were observed during the trials.

The reflections were generally directed to the welding direction, see Fig. 7.9. Due to

the unstable surface of the melt pool, the direction of the reflections was oscillating

randomly and the power was usually harmlessly distributed over a large area.

Fig. 7.9 Reflection of the laser beam towards the direction of welding

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The angle of the reflections was dependent on the thickness of the bead being

generated; the thicker the bead, the more obtuse the angle of reflection. During the

low deposition rate bead on plate tests where the weld pool was near horizontal the

almost vertically reflected light occasionally damaged unprotected components such

as the shielding gas piping in the laser head. In the similar metal narrow gap welds,

in which thicker bead was being laid and the weld pool surface was more inclined,

the reflections were directed more forward, out of the groove hitting generally the

filler wire nozzle area. The angle of the reflections was especially harmful for the

dissimilar NGLW, where LII imaging was able to record the melting of the wire

feeder nozzle and the convection of the resulting debris into the weld pool, see Fig.

7.10 and Fig. 7.11.

Fig. 7.10 Overheating of wire nozzle (arrow) due to reflections. DS2 pass 2, 38 s

from the start.

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Fig. 7.11 Copper debris (two right arrows) from the damaged nozzle (left arrow)

being carried to the melt pool by the filler wire. DS2 pass 2, 250 ms after

Fig. 7.10.

The issue of the harmful reflections was countered by experimenting with different

fore-aft tilt angles for the laser beam, as is the industry-standard practice [40]. This

was found ineffective as the melt pool surface oscillated as did the reflections.

During the first trials, the nozzle was made from a relatively thin copper piping,

which was very easily damaged by the heat. The molten bits of the copper nozzle

would contaminate the weld. Construction of a new nozzle would take time and the

interpass temperature would drop, further jeopardising the quality of the weld. This

was successfully counteracted by designing new nozzles with quickly replaceable

nozzle tips. The nozzles were also designed to have much higher mass to be capable

of absorbing more heat without damage. The heavier mass was found to be an

efficient means of reducing the damage, virtually resolving the issue. After the

redesign, the main effect observed was harmless heating of the equipment, although

a protective brass shield was devised to protect the precision Z-stage from damage.

If asymmetry was generated in the weld, the reflections from the melt pool would

direct to the undercut side base material, see Fig. 7.12. This would lead to increased

undercut and raise the temperature of the material.

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Fig. 7.12 The detrimental effects of asymmetric reflections

In severe cases the laser power was sufficient to melt the BM sidewall several

millimetres above the undercut, damaging it by melting grooves to it. This pitting

would later cause problems as the side wall would lose its intended shape. When the

weld progressed to the level of these undulations, they effectively acted as undercut

themselves and increased the undercut and asymmetry even further.

7.8 Alignment of the Filler Wire

The wire alignment requirements were found to be concordant with the observations

of Jokinen 2004 [17] and Salminen 2003 [156]. The alignment was found to be near

immune to transverse offset. Even a severely offset wire would produce a

symmetrical weld. This was determined to be related to the weld pool being

turbulent enough to be able to distribute the material evenly. Undershooting the weld

pool slightly caused usually no issues as the previous bead would guide the wire to

the melt pool satisfactorily. Only in severe cases, where the previous bead was

316 L SA508

392

uneven and wire interaction point several mm in front of the melt pool, some

tangling and vibration of the wire were observed. Overshooting the weld pool was

found detrimental to the quality of the weld as the wire would be heated prematurely

by the direct interaction with the laser beam above the melt pool; see Fig. 7.13 and

Fig. 7.14. This would lead to droplet formation at the tip of the wire leading to

increased oxidation of the increased interaction surface of the molten metal and the

local atmosphere. There would also be unevenness of the bead as the filler material

feed would be discontinuous, which would cause indentations prone to lack of

fusion.

Fig. 7.13 Droplet formation due to filler wire inadvertently hitting the laser beam

above the melt pool. Oxidation of the droplet visible in video material.

Maximum droplet size shown, just before gravity pulls the droplet down to

the meltpool. DS2, pass 1.

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Fig. 7.14 Droplet at the moment of being absorbed by the meltpool, 40 ms after Fig.

7.13. DS2, pass 1.

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8 Conclusions

8.1 Introduction

Current welding technologies for PWR reactor PCC dissimilar welding are dated,

costly and may not be able to produce the longest possible lifespan. In order to

alleviate these issues, a novel process called narrow gap laser welding is suggested.

In this study, a welding setup capable of 40 mm thick dissimilar NGLW was

developed. A three-stage welding programme was carried out with the intention to

generate dissimilar metal narrow gap laser welds for extensive analysis. The main

goal was to evaluate the welds against two main industry-standard evaluation

criteria, the ASME IX radiographical acceptance and the U.S. NRC impact

toughness requirements for light water reactor coolant boundary components.

Altogether 19 thick section narrow gap laser welds were completed. Of these, 6 were

stainless steel similar metal welds and 13 were dissimilar metal welds using FM52

filler metal. 5 of the dissimilar metal welds were made between S275 mild steel and

316L stainless steel. The 8 pedigree DMW performed were made using

SA508 Gr3 Cl2 low alloyed high strength steel and 316L.

8.2 Conclusions

The following conclusions were reached:

Dissimilar laser narrow gap welding was found to be feasible. This was established

by subjecting the welds to the industry-standard quality requirements. The welds

were first subjected to visual inspection. The welds found flawless were

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radiographed to ASME IX radiographical acceptance. Flawless sections of a weld

were cut for BS EN ISO 148 Charpy-V impact toughness analysis. The impact

energy results were subjected to U.S. NRC acceptance criteria for light water reactor

PCC boundary components. Weld quality fulfilling both the ASME IX

radiographical integrity and U.S. NRC fracture toughness requirements was

achieved. Consequently, a level of quality acceptable for light water nuclear power

reactor coolant boundary components was reached.

Two mechanisms leading to contraction of the welding groove were identified. To

allow access to the welding groove, the width of the groove should be wider than the

thickness of the various nozzles involved. The distortions were found heavily

dependent on the rigidity of the restraint system. The large scale butterfly distortion

was responsible for 50% of the contraction and local yielding of the material close to

the groove another 50%. This led to the conclusion that even a perfectly restrained

weld, i.e. having no butterfly distortion, would suffer from contraction and distortion

should be investigated individually for each restraint scenario.

Powerful beneficial multi-pass tempering was detected. The welds were cross-

sectioned, hardness mapped and etched to analyse the microstructure. The high

hardness of the martensitic SA508 HAZ was tempered after one heat cycle. The

resulting grain size was small, indicating good toughness. The effect on the

microstructure was found negligible after 3 tempering cycles.

The properties of the weld were uniform through the thickness. Longitudinal and

composite tensile test and impact toughness coupons were EDM cut near the root,

middle and close to the top of the weld. The results were supported with the hardness

mapping results. With the exception of the hardness of the untempered top pass

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SA508 HAZ, all mechanical properties investigated were found uniform through the

thickness of the weld indicating a uniform multi-pass tempering.

The industry-standard impact toughness requirements for the weld were met already

in as welded condition. This indicates that the PWHT is not required for the sake of

restoring the toughness of the SA508 HAZ when using NGLW.

NGLW was able to produce crack-free welds using Inconel Alloy FM52. The Filler

metal FM52 is known for generating solidification, liquation and ductility dip

cracking, especially in highly restrained welding conditions such as thick section

welding. Despite this no cracking of the two first types was observed in the analysis.

Some sporadic signs of discrete ductility dip cracking were found. This is an

indication of the benefits of NGLW considering residual stress generation in thick

section welding.

8.3 Process Characteristics

Certain factors considering the welding process were considered noteworthy. They

were found characteristical to the NGLW performed in this study.

Lack of fusion defects were found to be the principal welding flaw type in dissimilar

metal NGLW. This was judged to be the case based on radiographical analysis and

microscopy. This agrees with earlier findings in the literature of similar metal

NGLW. Three mechanisms causing LoF were found and strategies to reduce them

were developed. The following mechanisms were identified:

I. The (Ti and Al) oxides floating on top of the FM52 weld pool were observed

to flatten the melt pool surface preventing smooth wetting of sidewalls which

caused sidewall lack of fusion. This was seen in LII video monitoring

together with visual analysis of the individual weld beads. The type I LoF

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was countered by reducing and removing the oxides by improving gas

shielding and removing of the oxide layer mechanically between passes.

Change of filler material to FM52M is recommended, as it is developed

specially to reduce the floating oxide issue by different alloying.

II. Undercutting of the stainless steel was found to distort the welding groove

geometry. The phenomenon was monitored visually, severe cases being

apparent. In cases of excessive undercut, the sluggish FM52 weld pool was

not able to wet the stainless steel sidewall adequately leading to lack of

fusion. The type II mechanism was very harmful. A viable way to avoid the

undercutting was found to be avoiding excess laser power and heat input,

hence reducing the energy projected to overheat the stainless steel.

III. Using lower than optimal laser power resulted in an increase in the

probability of interpass lack of fusion. This was demonstrated by comparing

the cross-sectional macrographs against the laser power applied. The

oxidation of the weld pool was suspected to have added to this issue.

Increasing the laser power reduced the type III issues, as higher heat input

increases penetration and enhances the turbulence in the weld pool which

helps to convey the oxides to the surface.

Lack of sidewall fusion was also generated if the laser spot was too small for the

groove. Irregularities of the previous bead promoted the fusion issues via the less

than optimal geometry and by impeding the removal of oxidation in the crevices.

Very few issues with other welding flaws were encountered.

Minor irregularities during the welding process were found repairable by re-melting

the pass. In some cases, for example welding the root pass, the wetting of the

sidewalls was found inadequate and the weld bead generated was uneven. This was

proven repairable to an extent by re-melting the bead. The limitation of this method

was that each re-melt would introduce as much welding distortion as a normal filling

399

pass. To avoid excess contraction of the welding groove, the number of re-melt

passes should be limited.

Careful control of the welding parameters was required in DMW. In a test in which

lack of fusion between passes was detected due to insufficient laser power, an

increase of 500 W (17%) in the power was able to cause detrimental undercutting of

the stainless steel. Some reliability was sacrificed for targeting high efficiency by

using a maximal filler feed rate. It was concluded that using a more flexible welding

robot capable of real-time adjustment, would ease many of the requirements.

Accumulation of oxides was concluded to be the limiting factor for thickness with

the current welding setup and materials. The quality of fusion was found good until a

certain amount of oxides had accumulated on the weld bead. Fusion defects were

typically only found after several (10 to 13) welding passes had been deposited and a

thickness of 25 to 30 mm had been filled. This was based on macrography of weld

cross-sections. This suggested a limit to the number of individual passes applicable

unless the oxide formation was somehow reduced. Preventing the formation of

oxides by improving the gas shielding performance and removing the accumulated

oxide layer mechanically were found to be efficient in improving fusion quality and

flawless welding up to 40 mm was achieved.

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401

9 Future Work

This exploratory study is not by any means comprehensive. Several aspects in the

process can be improved and during the work, multiple new areas of interest were

found. The following future work is suggested.

Broadening the available weld applications by increasing the thickness

by improvements in weld design and materials

The 40 mm weld thickness achieved in this study is suitable for applications such as

pressuriser nozzles in PWR reactors. To allow the application of NGLW for all PWR

PCC dissimilar welds, the thickness needs to be increased. The maximum thickness

of the welds was limited due to weld oxidation. To address the oxide accumulation,

and allow for higher thickness welds to be made, the weld design could be

reconsidered. Welding only some of the passes with high-chromium nickel filler and

filling the rest with FM82 is suggested. This would create a sufficient corrosion

barrier, but reduce the use of the challenging filler material to a minimum. To further

reduce the oxidation propensity of the weld filler metal options FM52M and

FM52MSS could be considered. They also would address the susceptibility to DDC.

Introducing post weld heat treatment

Post weld heat treatment is required by different nuclear codes for thick section

welds in coolant boundary applications. The objective of the treatment is to temper

the potentially brittle microstructures to restore toughness and reduce susceptibility

to environmentally assisted cracking by reducing residual stresses. The welds in this

study were not post weld heat treated. The microstructures produced in this study

resulted in sound toughness properties due to the multi-pass tempering, but the

402

residual stress levels were still high and could still be reduced by tempering.

Exploring the resulting residual stresses, grain structure, hardness and impact

toughness properties in post weld heat treated condition would allow the properties

of the welds to be investigated in a condition where they are used in the target

applications.

Comparison with conventional welding processes using established

industry practices

Comparable welds with actual application designs and dimensions could be

produced using current standard procedures and welding processes. The welds would

be manufactured according to the industry-standard processes. Resulting welds can

then be analysed in direct comparison with each other.

Comparison of the contour method results with other residual stress

measurement methods

Other residual stress measurement methods, such as neutron diffraction or deep hole

drilling, could be done in order to validate the contour method data.

Investigation of corrosion properties

This study did not address the corrosion properties of the produced welds. A detailed

corrosion study of the welds would increase understanding of the behaviour and

longevity of the welds.

Investigations into laser hot wire welding

Hot wire laser welding has shown good potential in stainless steel similar metal

welding. HWLW has been reported to have a beneficial effect on fusion as part of

the heat is introduced by the hot wire. The additional heat input by the wire has been

reported to allow reduction of the laser power. This would be especially beneficial in

countering the undercutting of the stainless steel found in this study.

403

10 Publications

Journal papers in preparation

o “Thick Section Dissimilar Metal Narrow Gap Laser Welding of S275

to 316L Using Alloy 52 Filler”

o “Microstructural and Process Characterization of Dissimilar Metal

Narrow Gap Laser Welding of SA508 and 316L Using FM52”

o “Mechanical Properties and Residual Stresses of Thick Section

Dissimilar Metal Narrow Gap Laser Welds of SA508 Gr3 Cl2 and

316L”

Podium presentations

o Presentation

The 5th Industrial Laser Applications Symposium ILAS 2017.

“Thick Section Narrow Gap Laser Welding of Dissimilar Metals,

Low Alloyed Steel to Austenitic Stainless Steel”.

Grantham, Belton Woods Hotel 22. - 23.3.2017

o Presentation

39th International MATADOR Conference on Advanced

Manufacturing.

“Dissimilar Metal Narrow Gap Laser Welding of Thick Section Steels

316L and S275”. Manchester, The University of Manchester 5. -

7.7.2017.

404

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12 Appendices

Appendix I

Dissimilar welding experiments

417

Appendix II (1/8)

Radiographical examination report, sample DS3 (1/2)

418

Appendix II (2/8)

Radiographical examination report, sample DS3 (2/2)

419

Appendix II (3/8)

Radiographical examination report, samples DS4 and DS5 (1/2)

420

Appendix II (4/8)

Radiographical examination report, samples DS4 and DS5 (2/2)

421

Appendix II (5/8)

Radiographical examination report, sample SA2 (1/2)

422

Appendix II (6/8)

Radiographical examination report, sample SA2 (2/2)

423

Appendix II (7/8)

Radiographical examination report, samples SA5 and SA6 (1/2)

424

Appendix II (8/8)

Radiographical examination report, samples SA5 and SA6 (2/2)

425

Appendix III

AISI 316L material data sheet [187]

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Appendix IV

SA508 Gr3 Cl2 material certificate

427

Appendix V

Inconel Alloy FM52 material certificate

Melt no: NX5611JK, [Areva SA, 2016].

428

Appendix VI

Root thermocouple data SA1

429

Appendix VII

Top thermocouple data SA1

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