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Aspects of the electroslag welding process 1979

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ASPECTS OF THE ELECTROSLAG WELDING PROCESS by WILLIAM GORDON BACON B.A.Sc., U n i v e r s i t y of B r i t i s h Columbia, 1967 A THESIS SUBMITTED IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF DOCTOR OF PHILOSOPHY i n THE FACULTY OF GRADUATE STUDIES department of M e t a l l u r g i c a l Engineering We accept t h i s thesis as conforming to the required standard THE UNIVERSITY OF BRITISH COLUMBIA February, 1979. William Gordon Bacon I n p r e s e n t i n g t h i s t h e s i s i n p a r t i a l f u l f i l m e n t o f t h e r e q u i r e m e n t s f o r an a d v a n c e d d e g r e e a t t h e U n i v e r s i t y o f B r i t i s h C o l u m b i a , I a g r e e t h a t t h e L i b r a r y s h a l l m a k e i t f r e e l y a v a i l a b l e f o r r e f e r e n c e a n d s t u d y . I f u r t h e r a g r e e t h a t p e r m i s s i o n f o r e x t e n s i v e c o p y i n g o f t h i s t h e s i s f o r s c h o l a r l y p u r p o s e s may b e g r a n t e d b y t h e H e a d o f my D e p a r t m e n t o r by h i s r e p r e s e n t a t i v e s . I t i s u n d e r s t o o d t h a t c o p y i n g o r p u b l i c a t i o n o f t h i s t h e s i s f o r f i n a n c i a l g a i n s h a l l n o t b e a l l o w e d w i t h o u t my w r i t t e n p e r m i s s i o n . D e p a r t m e n t o f METALLURGY T h e U n i v e r s i t y o f B r i t i s h C o l u m b i a 2075 Wesbrook Place Vancouver, Canada V6T 1W5 D a t e MARCH 15, 1979 ABSTRACT The e l e c t r o s l a g welding (ESW) process has been inve s t i g a t e d u t i l i z i n g consumable guide welding procedures. E l e c t r o s l a g welding employing large square c r o s s - s e c t i o n electrodes was also studied. The study of the process parameters i s of importance i n understanding the r o l e which they play i n determining the grain s i z e attained i n the heat a f f e c t e d zone of the resultant weld. The large grain s i z e adversely e f f e c t s the impact strength: the gr a i n s i z e i s a d i r e c t consequence of the thermal h i s t o r y . The thermal h i s t o r y of the base metal was determined during welding f o r various p h y s i c a l configurations and welding schedules. The welding parameters having the major e f f e c t on the thermal h i s t o r y are: mode of welding, electrode immersion depth, s l a g depth, welding v e l o c i t y and s p e c i f i c power input. The penetration of the parent metal and s i z e of the heat-affected zone as w e l l as the grain s i z e have been shown to be a consequence of the thermal h i s t o r y . The quazi-steady state e l e c t r i c a l and temperature d i s t r i b u t i o n s i n the slag are seen to determine the temperature d i s t r i b u t i o n i n the metal. The preferred modes of welding are a l t e r n a t i n g current (AC) and d i r e c t current reverse p o l a r i t y (DCRP) with d i r e c t current s t r a i g h t p o l a r i t y (DCSP) being unacceptable due to the high i n c l u s i o n content of the weld when DCSP i s employed. The electrode immersion depth, or i n t e r e l e c t r o d e gap, has been found to determine the volume of the e f f e c t i v e heat source - i i - and thus the temperature and e l e c t r i c a l d i s t r i b u t i o n s . The s l a g depth and electrode v e l o c i t y are the p r i n c i p a l f a c t o r s which determine the i n t e r e l e c t r o d e gap when the s l a g chemistry and p h y s i c a l configuration are constant. There e x i s t ranges of s l a g depth and electrode v e l o c i t y values outside of which welding i s not achieved. An a n a l y t i c a l model has been employed which adequately describes t h i c k p l a t e welding p r a c t i c e , but which only applies to t h i n p l a t e welds i n a q u a l i t a t i v e manner. In t h i s model, an empirical equation has been used to c a l c u l a t e electrode immersion values which compare favourably with the measured values. In addition,the Bastein method of grain s i z e c a l c u l a t i o n has been applied to the thermal h i s t o r i e s obtained experimentally and has been found to be a p p l i c a b l e . In welds where impact strength i s an important design c r i t e r i a the weld/HAZ grain s i z e i s an important v a r i a b l e . We conclude that no combination of ESW process parameters x v r i l l lead to an acceptable HAZ grain s i z e g i v i n g HAZ impact values comparable with the parent p l a t e . In such cases f u l l heat treatment of the HAZ would be mandatory. - i i i - TABLE OF CONTENTS Chapter Page 1 INTRODUCTION ... . . . . . . 1 1.1 Introduction .. . . . . 1 1.2 Nature of the Process . . .. .. 2 1.2.1 Current Type and P o l a r i t y .. .. 3 1.2.2 ESW slags .. ... . .. .. ...... 4 1.3 Properties of E l e c t r o s l a g Welding ... .. 7 1.4 Nature of the Problem . . .. .. 9 1.5 Previous Work • .. . . . . . . . . 9 1.5.1 Temperature Measurements .. .. .. 10 1.5.2 C a l c u l a t i o n of Penetration . . . . .. 12 1.5.3 Solutions to Low HAZ Toughness .. 12 1.5.4 Process Parameter Investigations . . . . 14 1.5.5 The Present Research Program .. .. ,. 20 2 MODELS AND CALCULATIONS .. .. .. . . . . . . .. 21 2.1 Introduction .. .. .. . . . . . . . . . . 21 2.2 An A n a l y t i c a l Model .. . . . . . . .. 23 2.2.1 Assumptions .. 23 2.2.2 C a l c u l a t i o n of the Interelectrode Gap 24 2.3 A p p l i c a t i o n of the A n a l y t i c a l Model . . . . .. 25 2.3.1 Data Generated by the Model .. .. 25 - i v - Chapter Page 2 . 3 . 2 Manipulation of the Data .. .. .. .. .. 25 2 . 4 Summary .. . . . . . . 27 3 EXPERIMENTAL PROGRAM . . . . . . . . . . . . 28 3 . 1 ESW Devices at U.B.C . . . . . . . . 28 3 . 1 . 1 Wire or Consumable Guide Welders .. .. .. 28 3 . 1 . 2 Bar or Plate Electrodes .. .. . . . . .. 29 3 . 2 I n i t i a l Experimental Procedures . . 29 3 " . 2 . 1 Stage One 29 3 . 2 . 2 Stage Two, Consumable Guide to ESW .. .. 31 3 . 2 . 3 Slag Temperature .. .. . . . . . . . . . . 33 3 . 2 . 4 Electrode Immersion . . . . .. 33 3 . 2 . 5 Other Experimental Procedures .. .. .. 34 3 . 3 Butt Welds using Consumable Guide ESW 34 3 . 3 . 1 The Welding J i g .. .. 35 3 . 3 . 2 The Water Cooled Copper Shoes 36 3 . 3 . 3 The Consumable Guide Fixed Head 38 3 . 3 . 4 Run-in Sump and Run-out Blocks .. .. .. 39 3 . 3 . 5 Thermocouple P o s i t i o n s .. 39 3 . 3 . 6 Slag Additions and Other Procedures . . . . .. .. . . . . .. .. 40 - V - Chapter Page 3.3.7 E l e c t r i c a l Current D i s t r i b u t i o n Measurements . . . . . . 40 3.4 Butt Welds using Bar Electrodes . . . . . . .. 41 3.4.1 The Fixed Head '.. . . . . .. .. 42 3.5 Evaluation of Weld Properties .. .. "... .. 42 .3.5.1 Hardness Values .. .. 42 3.5.2 Toughness Values .. 43 3.5.3 Inclusion D i s t r i b u t i o n s .. .. 43 3.6 Experimental Records .. . . . . . . . . .. 43 4 RESULTS .. ... ...... .. .. 45 4.1 Power Supply C h a r a c t e r i s t i c s 45 4.1.1 Published Power C h a r a c t e r i s t i c s .. .. 45 4.1.2 Measured Power C h a r a c t e r i s t i c s . . . . 45 4.2 I n d u s t r i a l Experiments .. .. • 46 4.2.1 Experimental Conditions . . . . .. .. 46 4.2.2 Experimental Results .. 46 4.3 I n i t i a l Large Electrode Experiments .. . . . . 46 4.3.1 Temperature Time Cycles . . . . . . .. 46 4.3.2 Power C h a r a c t e r i s t i c s and Apparent Resistance of the Slag . . . . . . . . 49 4.3.3 Deposition Rate or Welding V e l o c i t y .. 49 4.3.4 Penetration and Inclusion D i s t r i b u t i o n . . . . 50 4.3.5 Hardness Traverses .. 50 4.3.6 Oxygen Analyses . . . . 50 1 - v i - Chapter Page 4.4 Consumable Guide ESW i n Cylinders 50 4.4.1 Introduction .. 50 4.4.2 Slag Temperature Measurements . . . . 50 4.4.3 Slag Depth, Immersion and Inter- electrode Gap Measurements . . . . .. 52 4.4.4 V e l o c i t y Measurements 52 4.4.5 Temperature Time Cycles.. . . . . .. 52 4.4.6 Penetration . . . . .. 55 4.4.7 Oxygen Ana l y s i s 55 4.4.8 Use of a 304 S.S. Consumable Guide .. 55 4.5 Consumable Guide ESW with A-36 Plates . . . . 55 4.5.1 Introduction 55 4.5.2 Temperature Time Cycles .. . . . . 57 4.5.3 Penetration and HAZ Values 57 4.5.4 Hardness Traverses .. .. 57 4.5.5 Inclusions . . . . 57 4.5.6 Grain Size Analysis .. 61 4.5.7 Charpy Traverses and Fracture Surfaces .. .. 61 4.6 ESW U t i l i z i n g Bar Electrodes .. .. . . . . 61 4.6.1 Introduction 61 4.6.2 Temperature Time Cycles .. •• .. 61 4.6.3 Penetration . . . . . . 64 4.6.4 Hardness Traverses . . . . . . . . .. 64 4.6.5 Charpy Traverses .. .. 64 4.7 E l e c t r i c a l D i s t r i b u t i o n .. .. 64 4.8 General Observations . . . . . . 66 - v i i - Chapter Page 5 DISCUSSION OF RESULTS . . .. .. 67 5.1 Discussion of POwer Supplies ..' .. .. .. .. 67 5.2 I n d u s t r i a l Experiments at Canron .. ' 68 5.2.1 The Apparent Resistance of the Slag Pool . . .. .. . . . . . . 6 8 5.2.2 Energy Requirements 69 5.3 Relating Published Experimental and Production Welding Energy Requirements .... ... .. 69 5.4 I n i t i a l Large Electrode Experiments .. 74 5.4.1 Comparison of the Temperature Time Cycles .. 74 5.4.2 Apparent Resistance of the Slag 75 5.4.3 Energy Requirements .. .. 75 5.4.4 Use of Inclusion D i s t r i b u t i o n Measurements. .. .. 76 5.4.5 Measuring Penetration and the Heat Affec t e d Zone using Hardness Traverses .. .. .. . . . . .. .. .. 76 5.4.6 Oxygen Analysis and the I n t e r p r e t a t i o n of the Results .. .. 77 5.5 Consumable Guide ESW i n Cylinders .. 77 5.1.1 An Overview of the Experimental Operating Conditions 77 5.2.2 I n t e r p r e t a t i o n of the Slag Temperature D i s t r i b u t i o n 77 - v i i i - Chapter Page 5.5.3 The E f f e c t s of Slag Depth, P o l a r i t y and the Resultant Immersion and Interelectrode Gap 78 5.5.4 V e l o c i t y Measurements and R e l i a b i l i t y . . 79 5.5.5 Explanation of the Temperature Time Cycles 79 5.5.6 Penetration and Temperature Time Cycles .. 80 5.5.7 Hardness Traverses . . . . . . . . .. 80 5.5.8 Oxygen Analysis Results .. .. 81 5.5.9 S t a i n l e s s Steel Consumable Guide and Mixing i n the L i q u i d S teel Pool 81 5.6 Consumable Guide ESW with A-36 P l a t e s .. .. 82 5.6.1 An Overview of the Experimental Operating Conditions .. .. .. .. 82 5.6.2- Explanation of the Temperature Time Cycles . . . . 82 5.6.3 Penetration and the Heat A f f e c t e d Zone as a Function of the Thermal P r o f i l e . . . . .. 83 5.6.4 Hardness Traverses . . . . 84 5.6.5 Measured Grain Size and Austenite Grain Size .. .. .. .. .. . . . . 84 5.6.6 Inc l u s i o n Analysis 85 5.6.7 Charpy Impact Values .. 86 5.7 E l e c t r i c a l D i s t r i b u t i o n i n Consumable Guide ESW . . . . 86 5.7.1 Possible Interpretation of the Current D i s t r i b u t i o n 87 5.7.2 C o r r e l a t i o n with the Slag Temp- erature D i s t r i b u t i o n . . . . . . .. 87 - i x - Chapter Page 5.8 ESW U t i l i z i n g Bar Electrodes 88 5.8.1 Overview of the Experimental Operating Conditions 88 5.8.2 Explanation of the Temperature Time Cycles.. . . . . . . . . . . . . . . . . . . . . . . . . 90 5.8.3 Hardness and Charpy Values 92 5.9 Large Electrode Technology and Implication .. .... 92 5.9.1 Potential Advantages and Dis- advantages . . . . . . .... .. 92 5.9.2 Potential Uses . . . . . . .... .. .-. . . . . . . 93 6 DISCUSSION OF MODELS .. .. .. .. .... .. .. .... ..... 94 6.1 Introduction - . • 94 6.1.1 The Analytical Solution Equation . . . . . . . . .94 6.1.2 The Electrode Immersion Equation . .. 97 6.1.3 Grain Size Determination 97 6.2 Application of the Analytical Model .97 6.2.1 Model Parameters 97 6.2.2 Relative Position of the Heat Sources .... 98 6.3 Comparison of the Model Generated and Measured Thermal Profiles for CGESW .. .... 99 6.3.1 Computer Calculations, D i g i t a l and Graphical Output .99 6.3.2 Analysis of Graphical Results 99 6 . 4 Comparison of the Model Generated and Measured Thermal Profiles for Published Results .. .... .. 100 Chapter Page 6.4.1 Computer C a l c u l a t i o n s , D i g i t a l and Graphical Output .. . . . ' 100 6.4.2 Analysis of the Predicted Temperature/ Time Relationships .. 103 6.5 A p p l i c a t i o n of the A n a l y t i c a l Solution to a Large E l e c t r o s l a g Weld . 103 6.6 A p p l i c a t i o n of the Electrode Immersion Equation .. .. 106 6.7 A p p l i c a t i o n of the Bastein et a l (67) Model to CGESW of A36 Plate .. .... . .. 108 7 CONCLUSIONS . .. - .. -. 10? 8 SUGGESTIONS FOR ADDITIONAL RESEARCH ,. .'. .. .. 112 REFERENCES . . . . . . . . . . . . . .. 114 APPENDIX A . . . . . . .. . . . . 119 APPENDIX B .. . . . . . 122 APPENDIX C . . . . 125 - x i - LIST OF TABLES Table Page 1 Common E l e c t r o s l a g Welding Fluxes . . . . . . . . 6 2 Welding Conditions f o r Thermal Cycles i n Figure 5 (from Patton, Ref. 1) .. 10 3 R e l a t i o n Between Welding Schedule and Weld Shape .. . . . . .. . . . . . . 15 4 . Canron Limited Welding Conditions .. 47 5 Canron Limited Welding Analysis .. .. 48 6 Experimental Conditions f o r Consumable Guide E l e c t r o s l a g Welding (CGESW) i n Cylinders 51 7 Slag Depth, Immersion and Interelectrode Gap Measurements .. . . . . .. .. 53 8 Welding V e l o c i t y Measurements 54 9 Experimental Conditions f o r Consumable Guide E l e c t r o s l a g Welding (CGESW) with A 36 Pla t e s .. .. 56 10 Penetration and Heat Affected Zone (HAZ) Values for the Experiments i n Table 9 . . 58 11 Grain Sizes i n ASTM Numbers at Various P o s i t i o n s . . 5 9 12 I n c l u s i o n D i s t r i b u t i o n s i n CGESW of A 36 Plates . . 6 0 13 Charpy Values at Various P o s i t i o n s i n Welded A 36 Plate . . . . .. .: . . . . . . 62 14 Experimental Conditions f o r ESW U t i l i z i n g Bar Electrodes . . . . .. .. 63 15 Charpy Values at Various P o s i t i o n s i n ESW Welded T - l Equivalent Plate 65 16 Experiment and I n d u s t r i a l Welding Conditions from the L i t e r a t u r e .. . 70 - x i l - Table Page 17 Input Data used for Solving the A n a l y t i c a l Solution f o r Consumable Guide E l e c t r o s l a g Welding (CGESW) with A 36 Plates .. .. . . . . . . . . .. . . 9 5 18 Input Data Used for the A n a l y t i c a l Solution f o r Published Temperature/Time Relationships .. 101 19 Electrode Immersion Values 105 20 Grain Size Determined f o r CGESW A 36 Pl a t e Experiments . . . . 107 - x i i i - LIST OF FIGURES Figure Number Page 1. Schematic representations of Consumable Guide E l e c t r o s l a g Welding (CGESW) 144 2. Schematic f l a t welding c h a r a c t e r i s t i c 145 3. Schematic comparison of thermal c y c l e s ; a, covered electrodes; b, submerged arc; c, ESW (Ref. 6) 146 4. Rela t i v e deposition rates f or various welding processes (Ref. 13) 146 5. Thermal cycles f o r welding conditions i n Table 2 (Ref. 1) .. 147 6. Thermal cycles from Mel'bard (Ref. 15) .. . . . . . . .. .. ... 148 7. Slag temperatures from-Mel'bard (Ref. 15) 14R 8. Charpy impact traverses values f o r ESW from Woodley et a l (Ref. 17) . . . . . . . . . . 149 9. DCRP and DCSP immersion and e l e c t r i c a l d i s t r i b u t i o n (Ref. 53) • • 150 10. Penetration versus amperage for various voltages (Ref. 54).. 151 11. Penetration versus amperage for two sl a g depths (Ref. 54) .. 151 12. Penetration versus amperage for slags of d i f f e r e n t c onductivity (Ref. 54) .. 152 13. Schematic representation of the a n a l y t i c a l model .. . . . . . . 153 14. Nomograph f o r determining the austenite grain s i z e 154 15. Thermocouple placement for i n i t i a l block welds .. 155 16. S i l i c a guides for Consumable Guide E l e c t r o s l a g Welding i n c y l i n d e r s 156 17. CGESW guided head for welding i n cy l i n d e r 157 18. I n i t i a l thermocouple set-up for CGESW of c y l i n d e r s .. . . . . 158 19. F i n a l thermocouple set-up for CGESW of c y l i n d e r s 159 20. Slag temperature measuring thermocouple .. .. 160 21. Immersion measuring loops 161 22. The welding j i g for butt welding . .. , , , ,, .. 162 - x i v - Figure Number Page 23. The cooling shoes . . . . . 163 24. The consumable guide f i x e d head f o r butt welding .. . . . . 164 25. Thermocouple p o s i t i o n i n g i n plate butt welding 165 26. Shunt set-up for measuring the current d i s t r i b u t i o n .. . . 166 27. Hobart 750 d i r e c t current power c h a r a c t e r i s t i c s .. .. . . . . 167 28. Hobart T-500 a l t e r n a t i n g current power c h a r a c t e r i s t i c s .. 168 29. Variable resistance current apparatus 169 30. T y p i c a l Canron weld i n c l u s i o n d i s t r i b u t i o n .. 169 31. Canron Oxygen d i s t r i b u t i o n and penetration .. 170 32. Canron Oxygen d i s t r i b u t i o n and penetration 171 33. Canron weld, m i c r o s t r u c t u r a l l y determined ( s o l i d l i n e ) and hardness traverses determined (broken l i n e ) HAZ . . . . 172 34. Electrode v e l o c i t y and weld v e l o c i t y f or d i f f e r e n t weld s i z e s . . . . . 173 35 Thermal p r o f i l e f o r DCRP .. .. .. . . .. 174 36. Thermal p r o f i l e f o r DCSP 175 37. Thermal p r o f i l e f o r AC .. 176 38. Thermal p r o f i l e comparisons for large electrode experiments .. .. . . . . .. • .. .. . 177 39. Inclusion d i s t r i b u t i o n f or DCRP 178 40. Inclusion d i s t r i b u t i o n f o r DCSP .. .. 179 41. Inclusion d i s t r i b u t i o n f o r AC 180 42. Hardness traverses for large electrode welds .. .. 181 43. Oxygen content traverses f o r large electrode welds 182 44. Slag temperatures for DCRP (experiment 10) 183 45. Experiment 10, v e r t i c a l s ection macrograph .. 183 46. Slag temperatures f o r DCSP (experiment 3) 184 47. Experiment 3, v e r t i c a l section macrograph 184 - XV - Figure Number Page 48. Slag temperatures f o r AC (experiment 14) 185 49. Experiment 14, v e r t i c a l section macrograph .. .. . . . . .. 185 50. Thermal cycle f o r DCRP experiment 10 .. j.86 51. Thermal cycle f o r DCSP experiment 8 .. 187 52. Thermal cycle f o r AC experiment 14 .. .. .. 188 53. Example Vickers hardness traverses for CGESW c y l i n d e r s .. 189 54. Representative oxygen values for CGESW cy l i n d e r s .. .. .. 190 55. Macrograph of etched weld produced using a 304 S.S. consumable guide 191 56. C a l i b r a t i o n curve f o r s l a g feeder and method f o r determining s l a g a d d i t i o n rate .. .. — 192 57. Thermal p r o f i l e s f o r a l l thermocouples \ inch from the o r i g i n a l surface f o r experiment P15 193 58. Thermal p r o f i l e s f o r a l l thermocouples % inch from the o r i g i n a l surface f o r experiment P15 .. 194 59. Thermal p r o f i l e s f o r a l l thermocouples 3/4 inch from the o r i g i n a l surface f o r experiment P15 .. 195 60. Thermal p r o f i l e s f o r a l l thermocouples 1 inch from the o r i g i n a l surface f o r experiment P15 196 61. Thermal p r o f i l e s f o r experiment P15 .... 197 62 Thermal p r o f i l e s f o r experiment P04 .. , . 198 63. Thermal p r o f i l e s f o r experiment P05 . . 199 64. Thermal p r o f i l e s f o r experiment P08 200 65. Thermal p r o f i l e s f o r experiment P10 201 66. Thermal p r o f i l e s f o r experiment P l l .. .. 202 67. Thermal p r o f i l e s f o r experiment P13 203 68. Thermal p r o f i l e s f o r experiment P14 .. 204 69. Thermal p r o f i l e s f o r experiment P16 .. 205 70. Thermal p r o f i l e s f o r experiment P18 .. .. 206 71. Thermal p r o f i l e s f o r experiment P21S .. . . . . . . 207 72. Thermal p r o f i l e s for experiment P22S 208 • .- x v i - Figure Number . . . . " Page 73. Thermal p r o f i l e s f o r experiment P23S . . . . . . . . . . . . . . 209 74. T y p i c a l hardness traverses f o r various degrees of fusion and models of operation , . . 210 75. Oxygen values corresponding to the same traverses as Figure 74 .. 211 76. Inclusions i n DCRP .. ........ . . . . . . 212 77. Inclusions i n DCSP .. .. ,. 212 78. Micrographs of grain s i z e v a r i a t i o n with distance from fusion boundary i n DCRP (P21). . . . . . . . . . . . . . . 213 79. Micrographs of grain s i z e v a r i a t i o n with distance from fusion boundary i n DCSP (P22). .. 214 80. Fractograph of Charpy fra c t u r e surface DCRP (P21). . . . . 215 81. Fractograph of Charpy f a i l u r e surface DCRP (P21)'., . . . . . 216 82. Fractograph of Charpy f a i l u r e surface i n weld zone showing d u c t i l i t y and i n c l u s i o n s (P22)_ , 216 83. Fractographs of Charpy f a i l u r e surface (P22) . . 217 84. Fractograph of Charpy fracture surface i n spheroidized zone (P22) . . . . . . 218 85. Fractograph of Charpy f a i l u r e surface (P23) . . . . . . . . 218 86. Fractograph of Charpy f a i l u r e surface 1.5 mm to HAZ from FB . 220 87. Thermal p r o f i l e f o r PP1 .. 221 88. Thermal p r o f i l e f o r PP3 .. . . . . . . . . .. .. . . . . .. .. 222 89. Thermal p r o f i l e f o r PP4 . 223 90. Thermal p r o f i l e f o r PP7 .. .. 224 91. Thermal p r o f i l e f o r PP8 225 92. Thermal p r o f i l e for E2 .. .. 226 93. Thermal p r o f i l e f o r E3 227 94. Thermal p r o f i l e f o r E4 228 95. Comparison of thermal p r o f i l e s f o r some s p e c i f i c power inputs . . 229 96. Hardness values f o r ESW u t i l i z i n g bar electrodes 230 - x v i i - Figure Number Page 97. Penetration versus s p e c i f i c power input, bar electrodes .. 231 98. Hardness values along the c e n t e r l i n e and near the surface of h o r i z o n t a l sections of T - l equivalent cast plates . ... 232 99. E l e c t i r c a l d i s t r i b u t i o n i n DCRP'.. . . . . . . . . . . . . . . . . 233. 100. E l e c t r i c a l d i s t r i b u t i o n i n DCSP . . . . . . . . .. 234 101. E l e c t i r c a l d i s t r i b u t i o n i n AC .. 235 102. Shunt set-up f o r P21S, DCRP . 236 103. Shunt set-up for P22S, DCSP 236 104. Shunt set-up f o r P23S, AC 236 105. T y p i c a l h o r i z o n t a l and v e r t i c a l sections of a CGESW weld.. 237 106. Penetration shape with bar electrodes i n the DCRP mode .. 238 107. Macroetched l o n g i t u d i n a l section of a weld produced during unstable welding conditions 239 108. Macroetched l o n g i t u d i n a l section of a weld produced during s t a b l e welding conditions .. 240 109. T y p i c a l e l e c t r o s l a g weld surface (DCSP) 241 110. E l e c t r o s l a g surface f or DCRP and AC . '.. ' 242 111. Weld run-out .. .. .. .. .. . . . . .. .. 243 112. Hor i z o n t a l and v e r t i c a l sectons of two t y p i c a l welds, used f o r oxygen a n a l y s i s , hardness traverses, grain s i z e measurements and Charpy impact 244 113. Casting produced when welding with too deep a s l a g . . . . 245 114. Voltage-amperage c h a r a c t e r i s t i c s f o r Canron x-^elds .. .. 246 115. Amperage versus apparent resistance f o r Canron welds . . . . 246 116. Amperage as a function of wire v e l o c i t y 247 117. Apparent resistance as a function of wire v e l o c i t y for Canron welds ' . .' 247 118. Energy per gram of metal deposited f o r Canron welds . . . . 248 119. .. L i t e r a t u r e c i t e d voltage-amperage c h a r a c t e r i s t i c s 248 120. L i t e r a t u r e c i t e d amperage-welding rate r e l a t i o n s h i p s .. .. 249 - x v i i i - Figure Number ; ' Page 121. L i t e r a t u r e c i t e d apparent resistance dependence on welding v e l o c i t y . .. 250 122. L i t e r a t u r e c i t e d energy requirements . . . . . . . . . . . . . . 251 123. Electrode immersion versus apparent s l a g resistance . . . . 252 124. Electrode immersion versus energy input .. 252 125. Penetration versus energy input .. .. .. .. 253 126. Penetration versus s l a g depth . 254 127. Temperature p r o f i l e across weld sections 255 128. Immersion as a function of apparent r e s i s t a n c e .. . . . . . . 256 129. Amperage as a function of apparent res i s t a n c e .. . . . . .. 256 130. T h e o r e t i c a l s l a g flow pattern from D i l a w a r i (75) 257 131. T y p i c a l bar electrode t i p s 258 132. Model predicted thermal p r o f i l e f o r P05 ... . . 259 133. Model predicted thermal p r o f i l e for P06 .. 260 134. Model predicted thermal p r o f i l e f o r P08 .. 261 135. Model predicted thermal p r o f i l e f o r P09 . . . . . . . . . . .. 262 136. Model predicted thermal p r o f i l e f o r P10 263 137. Model predicted thermal p r o f i l e for P l l .. . . . . 264 138. Model predicted thermal p r o f i l e for P12 . . . . . . 265 139. Model predicted.thermal p r o f i l e f o r P13 . . . . 266 140. Model predicted thermal p r o f i l e f o r P14 .. .. 267 141. Model predicted thermal p r o f i l e f o r P18 .. . . . . . . .. .. 268 142. Model predicted thermal p r o f i l e f o r P20 . . . . . . . . . . . . 269 143. Predicted thermal p r o f i l e for experiement 1(a), Pugin (16) . 270 144. Predicted thermal p r o f i l e f o r experiement 1(b), Pugin (16) 271 145. Predicted thermal p r o f i l e f o r experiments, Pugin (16) . . . . . . . . 272 - x i x - Figure Number Page 146. Predicted thermal p r o f i l e f o r experiment 3, Pugin (16) 273 147. Predicted thermal p r o f i l e f o r experiment 4, '. Pugin (16) .. .. . . . . . 274 148. Predicted thermal p r o f i l e f o r experiment 1, Sharapov (17) 275 149. Predicted thermal p r o f i l e f o r experiment 1(b) . . Sharapov (17) .. . . . . . . . . . . . . 276 150. Predicted thermal p r o f i l e f o r experiment 3, Sharapov (17) 277 151. Predicted thermal p r o f i l e for experiment 1, Sharapov (18) .... .. .. .. .. 278 152. Predicted thermal p r o f i l e f o r experiment 1, Trepov (19) . . . . 279 153. Predicted thermal p r o f i l e f o r experiment 2, Trepov (19) . . 280 154. Predicted Isotherms i n a large e l e c t r o s l a g weld 281 - xx - LIST OF SYMBOLS Symbols T temperature, °K or °C t time, seconds To preheat temperature, °C -1 -1 q heat source,cal.sec cm v weld v e l o c i t y , cm.sec ^ ' x distance, cm. y distance, cm. 2 a c o e f f i c i e n t of thermal d i f f u s i v i t y cm sec. b c o e f f i c i e n t of heat t r a n s f e r sec. ̂ -1 -1 -1 k thermal conductivity, cal.cm sec °C - 1 - 1 c s p e c i f i c heat, cal.gm. °C p density, gm.cm 6 thickness of plat e s being welded, cm. V voltage, v o l t s I amperage, amps. n e f f i c i e n c y , u n i t l e s s . R resistance, ohms. p s p e c i f i c r e sistance, ohm-cm. L immersion, cm. r radius of electrode, cm. - x x i - ACKNOWLEDGEMENTS The author s i n c e r e l y appreciates the advice and d i r e c t i o n provided by his research d i r e c t o r , Dr. Alec M i t c h e l l , throughout the course of t h i s p r o j e c t . The author i s indebted to Mr. Jim Brezden and Mr. Rick Palylyk f o r t h e i r assistance throughout the program. The author also wishes to thank the e n t i r e t e c h n i c a l s t a f f f o r t h e i r h e l p f u l suggestions. He thanks the departmental f a c u l t y members and fellow grad- uate students f o r t h e i r f r i endship and endless encouraging d i s c u s s i o n . F i n a n c i a l assistance i n the form of the National Research Council of Canada Postgraduate Scholarship (.1969-1972) and a Stelco Graduate Research Fellowship (1972-1974) i s greatly appreciated. A s p e c i a l thanks i s given to Dr. Ed Teghtsoonian, Dr. N e i l Riseborough and Dr. K. C. Donaldson for t h e i r moral support and under- standing. - 1 - CHAPTER 1 INTRODUCTION 1.1 Introduction Since the end of World War II there has been a marked increase i n the s i z e of fabricated equipment which i n turn has caused a marked increase i n the demand for e f f i c i e n t methods of j o i n i n g or welding of large component parts. The increase i n demand for welding heavier sections has lead to the development of lower cost, f a s t e r methods, one of which i s e l e c t r o s l a g weldin (ESW). ESW was developed by Prof. B.E. Paton at the E.O. Paton E l e c t r i c Welding I n s t i t u t e of the Academy of Sciences of the U.S.S.R. and h i s book (1) i s the only a u t h o r i t a t i v e book a v a i l a b l e on ESW. There has been a s t e a d i l y i n c r e a s i n g demand f o r the use of ESW, e s p e c i a l l y i n the following areas: 1. Long, heavy sec t i o n , v e r t i c a l or n e a r - v e r t i c a l butt and T-type welds. 2. Large, heavy s e c t i o n c i r c u m f e r e n t i a l welds. 3. J o i n i n g of large ingots f o r subsequent forging. 4. Surfacing of both f l a t and c y l i n d r i c a l surfaces. 5. As a replacement f o r automatic submerged arc. 6. J o i n i n g of large castings, - 2 - 1.2 Nature of the Process ESW i s a welding process u t i l i z i n g the heat produced by an e l e c t r i c current passing through a molten slag to produce a f i n i s h e d weld i n a s i n g l e pass. The parent metal components are placed i n a v e r t i c a l p o s i t i o n and secured with a gap between them ( F i g . 1). The bottom of the weld i s sta r t e d i n a constructed sump and the top of the weld ended i n a set of run-out blocks. A set of water cooled copper shoes (or molds) are a f f i x e d against the side of the parent metal, overlapping the gap completely. The shoes may be e i t h e r the continuous length of the weld or moved up as welding proceeds. The electrode i s lowered i n t o the sump u n t i l a r c i n g s t a r t s at which time the granulated slag i s added slowly to extinguish the arc and thus provide the s l a g pool. Enough sl a g i s added to prevent arcing,and keep the process sta b l e . The electrode i s lowered at a constant r a t e , with slag being added continuously u n t i l the weld has progressed into the run-out blocks and the electrode i s stopped. There are various types of ESW that have been patented and are i n use: 1. Standard ESW employing from 1 to 15 bare electrode wires from 2 to 5 mm. i n diameter. This i s b a s i c a l l y the process developed by Prof. B.E. Paton. 2. Consumable Guide ESW. The guide may be f l u x coated or bare and cons i s t s of a non-moving nozzle throughout the length of the weld. It i s e s s e n t i a l l y a s t a t i o n a r y electrode with a moving electrode i n i t s centre. Both the wire and the guide are thus f i l l e r metals (1,2,3,4,5). The consumable guide i s u s u a l l y coated with a slag . with a composition the same as the working: s l a g , and theiguide ".' . performs the fun c t i o n of i n s u l a t i n g the wire and providing make-up - 3 - slag. This method i s the most common when welding with wires and weld lengths that are l e s s than 3m. (Figure 1). 3. Plate Electrode ESW employs an electrode or electrodes that have the shape of the gap between the parent metal components. There may be one or more plates some of which may simply be f i l l e r metal ( i . e . e l e c t r i c a l l y neutral)(6,7,8,9,10). 4. Use of an a d d i t i o n a l f i l l e r metal,, granulated i n texture, to act as a f i l l e r or as an a l l o y i n g agent i s p r a c t i c a l with any of the above three standard processes. Various other ESW procedures have been t r i e d but few are i n common use and are not dealt with herein. 1.2.1. Current Type and P o l a r i t y There.are two electrodes,.the f i l l e r metal and the parent metal, as i n a l l welding processes. The welding can be accomplished by one of three methods: 1. D i r e c t c u r r e n t , . s t r a i g h t p o l a r i t y ( i . e . electrode or f i l l e r metal i s negative) which i s designated D.C.S.P. (This method i s not commonly used i n ESW). 2. D i r e c t current, reverse p o l a r i t y ( i . e . electrode or f i l l e r metal i s p o s i t i v e ) which i s designated D.C.R.P. 3. A l t e r n a t i n g current designated A.C. There are d i s t i n c t d i f f e r e n c e s between these methods with respect to the welding conditions.and r e s u l t a n t p h y s i c a l p r o p e r t i e s . D.C.S.P. i s v i r t u a l l y unused because of dele t e r i o u s i n c l u s i o n contents when the slag/ metal pool i n t e r f a c e i s p o s i t i v e . D.C.R.P. and A.C. are common and both are used, but A.C. i s more common. When employing ESW (and other - 4 - welding processes), i t i s advantageous to have a f l a t welding c h a r a c t e r i s t i c (Fig. 2) or, as i t i s sometimes c a l l e d , a constant voltage source. ESW does not lend i t s e l f to a constant-current-type power supply. The constant- voltage-type source allows v a r i a t i o n of the current without varying the voltage s i g n i f i c a n t l y . V a r i a t i o n of the welding current i n ESW i s accomplished by varying the electrode feed r a t e which i n turn controls the welding r a t e . This explains why A.C. i s used more than D.C.R.P.; i t i s almost impossible to construct economically a D.C. welder with a completely constant voltage character- i s t i c and consequently most welding of t h i s type has a drooping character- i s t i c (with i n c r e a s i n g current, voltage drops). With A.C. i t i s p o s s i b l e to vary the current over a very wide range without a change i n Voltage. This feature i s very important because i t allows f o r a more stable process i f voltage and current can be vari e d almost independently. The mode of welding also d i s t r i b u t e s the current d i f f e r e n t l y , thus d i s t r i b u t i n g the heat d i f f e r e n t l y . Thus, i n a c t u a l p r a c t i c e , d i f f e r e n t weld- ing schedules are necessary for D.C.R.P. and A.C. i n order to achieve the same end r e s u l t s . 1.2.2 ESW slags Slags or fluxes used f o r the ESW process must have the f o l l o w i n g features: 1. The f l u x must allow the process to be s t a r t e d e a s i l y and maintained over a range of s l a g depths, welding voltages and amperages. 2. The sla g must wet the melting parent metal surfaces i n order to a t t a i n complete f u s i o n and c o n t r o l l a b l e penetration. - 5 - 3. The slag should not become too conductive at high temperatures or e f f i c i e n c y w i l l s u f f e r accordingly. 4. The s o l i d i f i e d slag should be r e a d i l y removable. 5. Chemical c o n p a t i b i l i t y with the system to be welded. The e l e c t r i c a l conductivity and the v i s c o s i t y of the s l a g are i t s two most important p r o p e r t i e s . Highly conductive slags f a c i l i t a t e the s t a r t i n g of the process and the s t a b i l i t y of the process. The e l e c t r i c a l c o nductivity also governs the amount of heat generated by the current passage and thus the welding rate a t t a i n a b l e . The v i s c o s i t y of the s l a g i s also important i n the process from both a p r a c t i c a l and operating point of view. I f the v i s c o s i t y i s too low ( i . e . a "long" slag) , the s l a g w i l l leak between the shoes and the p l a t e s ; i f the v i s c o s i t y i s too high, the shoes commonly acquire very large 'buildups' and weld undercutting can occur. Also, the v i s c o s i t y p a r t i a l l y governs the motion of the s l a g which greatly a f f e c t s thermal d i s t r i b u t i o n i n the s l a g and consequently the a b i l i t y to weld. - Table 1 i n d i c a t e s the major slags and t h e i r compositions used i n the U.S.S.R. (the most common quoted i n the l i t e r a t u r e ) . ESW slags are low i n SiO^ and high i n MnO because both allow far better s t a b i l i t y . Calcium f l u o r i d e i s added to increase the e l e c t r i c a l c o n d u c t i v i t y and lower the v i s c o s i t y . The chemical a c t i v i t y of the slag i s also very important be- cause i t i s necessary to wet the oxidized surfaces. Thus, a s l a g i s required that can reduce the oxide films which form on the weld edges when the surface preparation consists of flame c u t t i n g only. This r e s u l t s i n one of the advantages of ESW TABLE 1 Common electroslag welding fluxes Designation SiO, A1 20 3 Chemical composition in weight percentages MnO CaO . MgO CaF, Others AN-8 AN-8M AN-2 2 FTs-7 AN-25 ANF-1 ANF-7 ANF-6 33-36 35-38 18-21.5 46-48 6-9 5 max. 11-15. 5.5 max. 19-23 3 max. 21-26 28-32 7-9 24-26 4-7 4-8 12-15 3 max. 12-15 20 5-7 1.0 max. 11.5-15 16-18 2-4 35 13-19 12-16 20-24 5-6 33-40 92 min. 80 65 30-40 T10, i ON I - 7 - 1.3 Properties of Electroslag Welding The most outstanding features of the process are the high deposition rate and the slow welding rate, hence i t s many advantages and i t s few disadvantages. The so called "soft" thermal cycle. (Fig. 3) leads to the uniqueness of ESW compared to other processes. Any point i n the parent metal w i l l reach a particular temperature more slowly and also cool more slowly. This i s due to the low welding velocity and comparatively high heat input of ESW. The slow cooling rate allows for the welding of steels that are susceptible to.undesirable decomposition products at high cooling rates. Also preheating i s eliminated because the i n i t i a l heating i s accomplished by radiation from the slag surface and conduction from the slag. However, the longer time at higher temperatures causes excessive grain growth in the heat affected zone. The advantages of the process are: i . Increased efficiency compared to manual or submerged arc; once a weld has commenced, the operator efficiency i s 100% for ESW compared to 40-60% for manual or submerged arc welding. A greater current efficiency also results, 20 g. per amp-hr for ESW versus 15 g. per amp-hr (6) for submerged arc; also 1.5 kWH/kg for ESW versus 2.5 kWH/kg for submerged arc (13) (Figure 4). i i . V i r t u a lly no edge preparation i s necessary for ESW except where the faces are in contact with the cooling shoes, i i i . There i s a decrease i n the pos s i b i l i t y of porosity due to the slow moving sol i d i f i c a t i o n front and the continuous molten slag cover, iv. There is less distortion i n ESW because the weld i s made in a single pass (12). The single pass feature also decreases weld inspection costs. - 8 - v. The process i s very e f f i c i e n t thermally, with Paton quoting 1.2% of the energy input being l o s t to r a d i a t i o n and 7.8% to the cooling water (1); the percentage of t o t a l heat l o s t to the shoes decreases s u b s t a n t i a l l y with thickness (12). v i . There i s u s u a l l y no need to undertake any f i n i s h i n g procedures other than the removal of the s i n g l e s l a g l a y e r from both sides. The disadvantages i n ESW are two f o l d i n nature, i . e . operating and m e t a l l u r g i c a l . The operating disadvantages are: i . Set up and i n i t i a l c a p i t a l outlay i s expensive for shoes and electrode feeders, i i . A power f a i l u r e r e s u l t s i n a large defect that must be manually repaired. i i i . Multiwire feeding machines are complicated and require s t r i c t maintenance. As stated e a r l i e r the ESW process i s characterized by the slow heating and cooling rates which lead to the m e t a l l u r g i c a l problems associated with ESW welds. The HAZ i s above 1100°C for a long period of time, thus causing the austenite grains to grow excessively large and the cooling r a t e i s so slow that the austenite decomposition to i t s products i s slowed. The f i n a l structure on cooling i s u s u a l l y coarse grained f e r r i t i c / p e a r l i t i c or f e r r i t i c / b a i n i t i c , with some retained austenite or martensite (14) i n the HAZ immediately adjacent to the fusion boundary. The superheating of t h i s region can also have other e f f e c t s i f the s t e e l i s micro-alloyed; the c a r b o - n i t r i d e s w i l l d i s s o l v e i n t h i s zone and, again, g r a i n growth w i l l occur ( t h i s i s also true of the AlN p r e c i p i t a t e s . ) . _ 9 - The problem w i t h ESW p r o p e r t i e s i s e s s e n t i a l l y a problem of tough- ness. In the HAZ i t can be a l l e v i a t e d by adding aluminum and n i t r o g e n (CaC^) , n o r m a l i z i n g , or quenching and tempering. However, i n most cases, these s o l u t i o n s are uneconomical or i m p r a c t i c a l and the tough- ness problem i n the HAZ of s t r u c t u r a l s t e e l welds and i n the h i g h h a r d e n a b i l i t y s t e e l s remains unsolved. This i s true even when A1N i s used to s t a b i l i z e the small g r a i n s i z e at very high temperature. This i s due to the long time at high temperature. 1.4 Nature of the Problem The problem of l a r g e g r a i n s i z e i n the HAZ can only be approached by a l t e r i n g the thermal c y c l e of the process. This can only be achieved by understanding how the heat i s d i s t r i b u t e d during the process and what v a r i a b l e s can be used to change the thermal generation and d i s t r i b u t i o n . 1.5 Previous Work The thermal d i s t r i b u t i o n during ESW has been measured many times. References 1, 13 and 15 through 20 are t y p i c a l . Some i n v e s t i g a t o r s (16-19) have attempted to p r e d i c t the thermal' c y c l e w i t h good success (but have not t r i e d to a l t e r the c y c l e ) . E r e g i n has attempted to use these thermal c y c l e s to p r e d i c t p e n e t r a t i o n (21) . However these i n v e s t i g a t o r s have not v a r i e d the welding c o n d i t i o n s d e l i b e r a t e l y to determine i f the thermal c y c l e and consequent p e n e t r a t i o n could be d e l i b e r a t e l y c o n t r o l l e d and ( i . e . the welding c o n d i t i o n s ) to minimize the p e n e t r a t i o n w h i l e a t t a i n i n g complete f u s i o n . The approach that has been taken i s one of f i r s t l y producing a sound welding technique and then studying the r e s u l t a n t m a c r o s t r u c t u r e s , micro- - 10 - TABLE 2 • • Welding Conditions f o r Thermal Cycles In Figure 5 (from Patton, Ref. 1) Operating Conditions • ESW One 3 mm Electrode ESW Two 3 mm Electrodes ESW \ 12 mm x 110 mm Plate Multx-Pass- Submerged Arc 1 2 Mode DC AC AC Welding V e l o c i t y (cm/sec) .019 .028 .022 1.111 1.111 Voltage (Volts) 38-40 44-46 30-32 32-34 32-34 Current (Amps) 450 450 950 500 500 Slag Depth (mm) 40 55 25-30 - 11 - structures and phy s i c a l properties (22-38), and then developing a heat treatment schedule or a l l o y i n g treatment to a l l e v i a t e any problems caused by the welding technique. Many in v e s t i g a t o r s have t r i e d varying the welding parameters to vary the weldment and/or HAZ (28, 39-58) structures with varying degrees of success and disagreement. Dilawari et a l (76) have developed a model of the ESR process which i n d i c a t e s how the metal flows during the process. The model i s useful i n describing the observed metal flow i n ESW. 1.5.1 Thermal Temperature Measurements Paton (1) reports the thermal cycle ( F i g . 5) for various welding con- d i t i o n s (Table 2) and makes comments on the slow thermal cycle f or ESW versus submerged arc but does not comment on the d i f f e r e n c e between wire electrodes and plate electrodes. Lefevre (12) measures the thermal cycle and p l o t s a contour map of isotherms around the electrode p o s i t i o n but makes no comment on the extremely high temperatures a t t a i n e d and how t h i s a f f e c t e d the HAZ (note: the welding conditions for the graphs were not provided). Mel'bard (15) p l o t s a type of thermal cycle f o r the HAZ, the p l o t ( F i g . 6) showing that there i s very s i g n i f i c a n t "preheating" of the parent metal above the slag surface. He also reports the mean temperature of the s l a g pool (about 1700° C) and the temperature p r o f i l e from the slag pool surface to the molten metal pool bottom ( F i g . 7). These data are used to postulate that the current passes i n the form of a cone from the electrode t i p to the slag/molten metal pool i n t e r f a c e , but no measurement of the cur- rent d i s t r i b u t i o n was undertaken. Voloshkevich (39) states that "nearly a l l the heat i s transferred to the parent metal through the weld puddle surface". Pugin and P e r t s o v s k i i (16) have developed a model for the heat d i s - t r i b u t i o n by assuming a s e r i e s of three moving l i n e a r heat sources and found - 12 very good agreement with t h e i r measured thermal c y c l e . However they used an act u a l e f f i c i e n c y of 58%, 32.5% being l o s t to the cooling water and 9.5% by r a d i a t i o n ; these are u n r e a l i s t i c a l l y high f i g u r e s (1,12,17,18). The only reported heat balance c a l c u l a t i o n s are those of Paton (1) and they are not referenced i n h i s book. 1.5.2 C a l c u l a t i o n of Penetration One model, by Eregin (21), has been derived to c a l c u l a t e penetration and was reported to be i n good agreement with the ac t u a l penetration. However most of the welding parameters were not presented and thus the model could not be evaluated c o r r e c t l y . No other references have attempted to c a l c u l a t e penetration as a function of changing welding parameters. 1.5.3 Solutions to Low HAZ Toughness Rote (23) has investigated the ESW of various grade s t e e l s from A 285 to a high carbon, a i r hardening s t e e l (Lukens D6AC). Only the A 212B and HY-80 were tested i n the as welded condition, a l l other weldments were tested a f t e r various complicated heat treatments. T y p i c a l of many i n v e s t i g a t i o n s , the r e s u l t s were inconclusive; i n some welds the weldment had be t t e r d u c t i l i t y and impact strength than the parent metal and i n other welds v i c e versa. Sometimes the HAZ had better impact strength than e i t h e r the weldment or parent metal. Zeke (24,32,56,57,59) has done exhaustive studies of ESW and written a review paper (59) f o r i t s use i n the production of s t e e l pressure v e s s e l s . He has determined that notch d u c t i l i t y i s too low i n the HAZ and has undertaken various heat treatment procedures on welds and specimens of parent metal to determine i f the p h y s i c a l properties could be improved. I t was determined that where the s t e e l was subjected to a very complicated heat- - 13 - treatment (740°C/15 h r . / a i r + 650°C/15 h r . / a i r ) , the p r o p e r t i e s of the parent metal could be achieved. Makara (34) studied the e f f e c t of heat t r e a t - ments between the AC^ and AC^ temperatures f o r ESW welded 1010 and 1015 and found e s s e n t i a l l y the same r e s u l t s can be achieved by 5 hr. at 780°C followed by furnace co o l i n g . Makara (35) has a l s o shown that 1015 can be heat treated s u c c e s s f u l l y by tempering at 650°C f o r 2 hrs., furnace cooled to 300°C and then a i r cooled. Eichhorn (37) studied annealing between AC^ and AC^ f o r c a r b o n i t r i d e former microalloyed s t e e l s to achieve im- proved impact strengths without normalizing (5 hours at 730°C, a i r cooled plus 4 hours at 630°C, a i r cooled) and achieved very good Izod impact values. Zeke and others (28) studied the weldmetal i t s e l f by e l e c t r o s l a g remelting the welding wire and t e s t i n g the as cast metal and the metal a f t e r various heat treatments. These studies commonly indicated that there i s some suit a b l e heat treatment schedule which w i l l give the weldment good p h y s i c a l p r o p e r t i e s , but i t i s f e l t that t h i s i s an unreasonable way to study the HAZ problem. Woodley (25) ESW welded Lloyds grade B and Lloyds. E and conducted wide p l a t e f r a c t u r e analysis and Charpy V-notch t e s t s and con- cluded that both t e s t s showed that the weldment and parent metal were at l e a s t as good as when submerged arc welded. The p l a t e s u s u a l l y f a i l e d i n the HAZ and the impact values i n the HAZ were lower than e i t h e r the parent metal or weldment. Woodley (27) also ESW welded JTA 101 (a normalized, s i l i c o n k i l l e d , carbon-manganese s t e e l of b o i l e r q u a l i t y ) and performed f u l l p l a t e slow bend t e s t s and charpy V-notch impact t e s t s ( F i g . 8). Malinovska, and Hrivnak (28) studied various heat treatments between the A^ and A^ to eliminate the large HAZ grains i n mild s t e e l and found annealing between A^ and A^ could replace normalizing. Bentley (31) welded mild s t e e l (Nb added) and performed Charpy V-notch te s t s i n the HAZ f o r the welded, normalized and - 14 - double normalized conditions. He found that the HAZ was severely embrittled and the embrittlement could be eliminated by s i n g l e or double normalizing, but double normalizing could lower the toughness of the parent metal. Tempering of an ESW welded complex chromium-manganese s t e e l (0.05% T i , 0.5% Mo, 0.25% C) embrittles the HAZ according to Braun (33), normalizing w i l l improve the toughness s u f f i c i e n t l y . Novikov (36) has shown that plates can be l o c a l l y normalized or tempered using i n d u s t r i a l frequency current f o r plates up to 200 mm t h i c k . This procedure gave very acceptable impact values and eliminated the requirements for very large furnaces, but does not appear to have been accepted commercially for plates over 50 mm. The use of c a r b o n i t r i d e formers to prevent the grain growth i n high energy input welding has been investigated very e x t e n s i v e l y by Ikino et a l (38). They found that only TiN could maintain the grain s i z e at temperatures approaching 1400°C and, i f enough TiN. p a r t i c l e s of l e s s than 0.1 p diameter were present, then the g r a i n s i z e i n the HAZ could be kept at l e s s than 100 u. This i s an advance i n s t e e l making technology and does not solve the general problem of ESW welding of s t r u c t u r a l and pressure v e s s e l s t e e l s i n e x i s t i n g s p e c i f i c a t i o n s . 1.5.4 Process Parameter Investigations Paton (1) has made a summary (Table 3) of the general e f f e c t of most of the process v a r i a b l e s i n ESW and devotes part of chapter 2 of h i s book (pp. 20-54) to t h i s subject. G o t a l ' s k i i (40) f i r s t (1954) proposed that the chemistry of ESW slags should be d i f f e r e n t from submerged arc slags based on the e l e c t r o l y t i c r e a c t - ions that take place during the process. G o t a l ' s k i i (41) showed that the b a s i - c i t y of the slag plays a major r o l e i n the reduction of s i l i c o n and increase i n manganese, and that the e l e c t r i c a l c o n d u c t i v i t y increases with b a s i c i t y . - 15 - TABLE 3 Re l a t i o n Between Welding Schedule and Weld Shape (from Patton, Ref. 1) Welding Conditions i n D i r e c t i o n of Increasing Values Weld C h a r a c t e r i s t i c From 400 Amp or Higher Welding Voltage Slag Depth Pool Depth Increases S l i g h t Increase S l i g h t Decrease Weld Width (Penetration) Decreases Increases Decreases - 16 - However, these slags contained no fl u o r s p a r and as such are not used ex- te n s i v e l y . G o t a l ' s k i i (42) then investigated the e f f e c t s of s i l i c a and f l u o r - spar i n ESW sla g s . He concluded that low s i l i c a f l u x e s would make the process' more stable. He based h i s conclusion.on the determination of the depth of the slag pool when arcing would i n i t i a t e ; when iow s i l i c a slags were used he found a r c i n g would commence only when the pool became very shallow. He also suggested that fluorspar would increase the conductivity of the slag and thus improve the s t a b i l i t y of the process. He made no conclusions on the e f f e c t of the slag depth on the welds produced or on the thermal changes that must have occurred when the slag pool depth increased. Z a i t s e v (44) has determined that f l u o r s p a r based slags have a stronger desulphurizing a c t i o n than s i l i c a t e based slags and that lowering the current and increasing the voltage helps desulphurizing. This data was produced on a small e l e c t r o s l a g r e f i n i n g u n i t . Shcherbina et a l (60) have investigated slag compositions with regard to c o n t r o l l i n g the temperature of the molten pool. They conclude that the cause of the extremely high temperatures i n the slag pool (1800-2000°C (15)) i s the r e s u l t of the high melting and b o i l i n g points of the slag components (Si02, Ca^2> 3' a n C ^ M n < ^ e t c ' ) * They blame t h i s excessive temperature for the high temperatures i n the HAZ and were unable with standard fluxes to -keep the slag pool cooler and s t i l l keep the process stable. They proposed that the slag chemistry could be al t e r e d so that the temperature of the stable process could be lowered and thus the penetration would be achieved by wetting.and not by excessive heat. They used.a f l u x with a f l u o r i d e - c h l o r i d e - boride base and s t a b i l i z e d the process with borax which has a low b o i l i n g point (1575°C). This slag was used to ESW c l a d an 18-8 type s t a i n l e s s s t e e l on a mild s t e e l backing. A completely continuous bond was produced with no - 17 - penetration of the mild s t e e l . They were also able to clad s t e e l with brass by lowering the slag temperature even furt h e r by adding a l k a l i n e metal car- bonates. They concluded that x^elding i n t h i s manner would decrease grain growth i n the HAZ, allow higher welding rates and decrease power consumption. Ostrovskaya (43) investigated the welding of SAE 1010 to 1045 grade s t e e l s using AN8 f l u x (22% S i 0 2 , 20% A l ^ , 21% C a F 2 > 32% MnO, 5% CaO) and varying the voltage, current ( i . e . electrode feed rate) and the thickness of weld per wire to determine the causes of s o l i d i f i c a t i o n cracking. I t was found that, i f the columnar grains growing i n t o the weld c e n t e r l i n e met nearly head on, c e n t e r l i n e cracks would develop. Otherwise a l l the cracks ob- served were between adjoining dendrite arms. The shape f a c t o r (weld width divided by weld depth, 40 was shown to be the c r i t i c a l parameter; the lower the shape f a c t o r the more l i k e l y s o l i d i f i c a t i o n cracking w i l l occur. I t was found that increasing carbon content was more c r i t i c a l than any other element (e.g. S, S i , P and Mn) and that welding of SAE 1040 or above would l i k e l y r e s u l t . i n c r y s t a l l i z a t i o n cracks. Increasing the current ( i . e . the electrode feed rate) decreases the penetration and decreases the shape f a c t o r , thus causing s o l i d i f i c a t i o n cracking. I t was also found that increasing the voltage has the opposite e f f e c t , because the penetration increases. Poznyak (46) welded 100 mm spheroidal (magnesium treated) cast i r o n using p l a t e electrodes of the same ma t e r i a l and determined that, at a f a i r l y high heat input (160 k c a l . cm , the slow thermal c y c l e i n the HAZ a c t u a l l y does not destroy the spherebidization structure as arc welding would using standard p r a c t i c e . Kozulin (47) and Dubovestkii (4) have in v e s t i g a t e d the ESW process using 5 mm diameter wire instead of the normal 3 mm diameter wire i n order to increase - 18 - the output of the process. They found that the welding r a t e could be increased but subsequent heat treatment was necessary before the HAZ impact strength could be r e a l i z e d , thus l i t t l e i f any gain was achieved. Ivochkin (49) researched the a d d i t i o n of f i n e i r o n powder i n t o the s l a g pool to lower the slag pool temperature, increase the e l e c t r i c a l e f f i c i e n c y and to demonstrate the c a p a b i l i t y of a l l o y i n g i n the same manner. This method doubled the wire-only welding v e l o c i t y , decreased penetration and increased the HAZ Charpy impact values. The process should work more e f f i c i e n t l y the f i n e r the powder as the superheating of the melted granules should not be excessive as i t i s f o r large d r o p l e t s . Makara (51) showed that the use of ANF-6 f l u o r i d e slag (60-70% CaF 2 and 30-40% A^O^) permitted lower voltages and shallower slag depths due to the higher conductivity of ANF versus the common AN-8. Ando et a l (52) have undertaken an extensive program to study the major parameters i n ESW that a f f e c t the r e s u l t a n t weld. They postulated that the convection currents i n the s l a g are p r i m a r i l y caused by the e l e c t r i c a l d i s - t r i b u t i o n . The e l e c t r i c a l d i s t r i b u t i o n was measured using an e l e c t r i c a l l y i nsulated l a y e r i n the weld ( i . e . a s e c t i o n of metal i n the h o r i z o n t a l plane). The e l e c t r i c a l d i s t r i b u t i o n was subsequently measured and i t was concluded that 21% of the t o t a l current was concentrated below the electrode t i p to the slag/molten metal i n t e r f a c e . Close examination of the measuring tech- nique allows c r i t i c i s m as the insulated layers can be shown to a f f e c t g r e a t l y the thermal d i s t r i b u t i o n i f the i n s u l a t e d layers are placed on both sides of the p l a t e . Also the current that was measured from below and f o r the l a y e r was not c a l c u l a t e d into current d e n s i t i e s to give a true f l u x . This current d i s t r i b u t i o n can.be used to predict the convection pattern they assume and the penetration p r o f i l e they observed. Other current d i s t r i b u t i o n s could also - 19 - produce other convection currents r e s u l t i n g i n the same penetration p r o f i l e . They found that i n the standard reverse p o l a r i t y (electrode p o s i t i v e ) the electrode melted more qui c k l y and was deeper i n the s l a g than f o r s t r a i g h t p o l a r i t y (electrode negative) f o r the same welding c o n d i t i o n s . In s t r a i g h t . p o l a r i t y the current d i s t r i b u t i o n was 62% to the metal below the ins u l a t e d layer and the electrode more deeply immersed ( F i g . 9). Ando et a l (53) i n t h e i r second report investigated the voltage and current e f f e c t s on penetration using d i r e c t current reverse p o l a r i t y . They r e l a t e d the penetration to current f o r three voltages ( F i g . 10) and two s l a g depths ( F i g . 11). I t should be noted here that Ando et a l i n r e f . 52 and 53 used a submerged arc f l u x (38 Si02> 26 MnO, 10 CaO, 17 A^O^) and not an ESW s l a g . In t h e i r t h i r d report (54).they studied s l a g composition con- d u c t i v i t y at high temperature ( F i g . 12). They found that, with a f l u x that has high c o n d u c t i v i t y at low temperature, the penetration was l e s s than for a f l u x that has low conductivity.at high temperature. This i s not s u r p r i s i n g i n view.of the Russian l i t e r a t u r e that has shown that penetration can be decreased using lower melting point slags (60). Feldmann et a l (55) investigated the use of manual arc welding, e l e c - tron beam welding and e l e c t r o s l a g welding i n the construction of a l a r g e 316 L bubble chamber. They were able to u t i l i z e ESW only where the excessive g r a i n growth i n the HAZ did not cause a s i g n i f i c a n t decrease i n the impact strength or where impact strength was not important. Lanyie and Zeke researched.the e f f e c t of current, p o l a r i t y and atmosphere on the weld metal.during e l e c t r o s l a g welding. As with most studies of t h i s type, there was no actual welding, simply e l e c t r o s l a g remelting of the welding wire into a small water cooled.copper c r u c i b l e . A l l the ingots were heat treated a f t e r the remelting and impact values found to he s a t i s f a c t o r y . He - 20 determined the r e l a t i v e impact strengths i n increasing order were electrode negative, a l t e r n a t i n g current, electrode p o s i t i v e . The d i f f e r e n c e s were a t t r i b u t e d to electro-chemistry. He also concluded that nitrogen and argon are superior atmospheres to a i r , argon plus oxygen or nitrogen plus oxygen. Shaheeb ( 5 8 ) has studied the influence of process parameters on the properties of ESW welds to determine i f furnace normalizing could be eliminated. He also determined that a l t e r n a t i n g current and reverse p o l a r i t y were much superior to s t r a i g h t p o l a r i t y . He concluded that a heat t r e a t - ment producing complete or p a r t i a l transformation of the structure was necessary to achieve the desired impact strengths. Sharapov ( 5 9 ) has investigated the grain growth of austenite i n ESW of Cr-Mo-V s t e e l s and, by simulating the thermal c y c l e , has been able to show the e f f e c t of the thermal c y c l e . He reports the time above A Q^ to be c r i t i c a l to austenite grain growth and the i n i t i a l part of the cooling cycle to be c r i t i c a l to the p a r t i a l decomposition of austenite. Vinokuruv ( 7 5 ) has predicted the transient s t r e s s and deformations around an ESW weld caused by structure changes and cooling rates. 1 . 5 . 5 The Present Research Program The purpose of the present work i s to determine the e f f e c t of c e r t a i n process v a r i a b l e s on the thermal c y c l e , and thus be able to minimize the inherent problems associated with ESW welding. An a n a l y t i c a l model (63,64) w i l l be compared with experimental data to determine i f the model can be used to write welding schedules. In order to do t h i s the process par- meters w i l l have to be varied and understood before the model can be used. Since well defined ESW structures are to be produced, a u s e f u l side approach i s to determine i f chemical differences between the modes of welding lead to d i f f e r e n t i n c l u s i o n compositions and thus a f f e c t i n g weldment Charpy values. - 21 - CHAPTER 2 MODELS AND CALCULATIONS 2.1 Introduction Rosenthal (61,62) and Rykalin (63) developed a l l the b a s i c s f o r the development of a n a l y t i c a l models i n the f i e l d of heat flow i n welding. Rosenthal f i r s t introduced the "quasi-stationary" state theory of heat flow as applied to a moving heat source i n 1935 (64). Rykalin's book (63) t i t l e d " C a l c u l a t i o n of Heat Flow i n Welding" greatly expands Rosenthal's work and applie s the equations to most welding applications of the l a t e 1940s. The study of the thermal properties of welding requires a d e f i n i t i o n . of heating and cooling cycles of the metal, the heat source(s) and s i n k ( s ) . The object of the welding exercise i s to provide a heat source by an economic means such that the parent metal components are joined by a fused boundary. The quasi-steady state theory has been applied by Rosenthal (61) and Rykalin (63) to p r e d i c t the temperature d i s t r i b u t i o n around a s i n g l e moving heat source i n the following a p p l i c a t i o n s : 1. rate of fusion of the welding electrode i n arc welding 2. arc welding of t h i n plates 3. flame c u t t i n g 4. submerged arc or manual arc welding 5. TIG and MIG welding 6. p r e d i c t i n g the cooling rates a f t e r welding - 22 - Most recently Goldack et a l (65,66) have applied the theory in Electron Beam Welding to predict the thermal distribution (65) and the micro- structures (66) of EB welded eutectoid steels with good agreement. The theory has also been applied to ESW by Pugin (16), Sharapov (17,18), and Trepov (19) with some success but not applying r e a l i s t i c thermal efficiencies from actual beat balances or using r e a l i s t i c distributions of heat sources. - 23 - 2.3 An A n a l y t i c a l Model 2.3.1 Assumptions Equation [2.1] i s the equation developed (63,64) f o r the p r e d i c t i o n of the thermal c y c l e f o r a s i n g l e moving heat source. However we wish to incorporate the concepts of preheat and multiple heat sources. The equation by simple ad d i t i o n becomes: 2 v t. i = l 2 2 2v A 2 , b ( v t i + y V 7 T + « 4a [2.1] where q_̂  i s the power of the i t h moving l i n e a r heat source and t ^ incorporates the p o s i t i o n of the i t h source, To i s the i n i t i a l preheat temperature. The use of equation [ 2 . l l requires a knowledge of the p o s i t i o n of the three heat sources. I f we assume the l i t e r a t u r e i s correct and that the highest heat evolution i s between the electrode and slag/molten metal pool i n t e r f a c e , then we can assume a configuration as shown i n F i g . 13. The important decisions to be made are with regard to a p p l i c a t i o n of the three heat sources, t h e i r r e l a t i v e strengths and p o s i t i o n s . Pugin (16) has assumed that q^ i s the at the slag surface, q2 i s intermediate between the electrode and the molten pool and q^ i s positioned at h a l f the depth of the molten pool. Using t h i s assumption the p o s i t i o n of q-̂  i s always known, q^ can be measured (but not pre- dicted) and q^ i s assumed f o r a p a r t i c u l a r welding schedule as measured a f t e r sectioning. The r e l a t i v e values of heat input are assumed to be 25%/50%/25% 3 fo r q^, q^ and q^ as i s the e f f i c i e n c y n(n = ^ q./O, where Q i s the t o t a l i = l a v a i l a b l e energy) by Pugin. The present research w i l l t r y to show that the p o s i t i o n of the heat sources can be assigned so that f o r a given set of welding conditions the - 2 4 - predicted welding thermal cycle i s comparable to the a c t u a l welding thermal cycle. 2.2.2 C a l c u l a t i o n of the Interelectrode Gap The p o s i t i o n of q^ i s i n the center of the i n t e r e l e c t r o d e gap and con- sequently i f i t could be calculated then the p o s i t i o n of c°uld be predicted. P e r t s o v s k i i (67) has researched the passage of the current through the pool i n ESW using a two dimensional model c o n s i s t i n g of copper cut-outs to represent the metal components and potassium permanganate soaked paper to represent the s l a g . The power source was. a 12V battery supply. The model did confirm the observations made when ESW i s i n operation: 1. the current density i s highest around the electrode t i p 2. more current leaves the end of the electrode than the sides, when the i n t e r e l e c t r o d e gap i s small He then proposed an expression f o r the approximate c a l c u l a t i o n of the r e s i s - tance of the slag to current flow. The expression i s : where: R = t o t a l apparent r e s i s t a n c e [ohms.] p = s p e c i f i c r e s i s t a n c e of the slag [ohm-cm] L = depth of the immersed electrode [cm] r = radius of the electrode [cm] Thus i f the t o t a l apparent res i s t a n c e of the s l a g pool and the s p e c i f i c r e s i s - tance i s known, the depth of immersion of the electrode can he c a l c u l a t e d , or i f the immersion i s known and the resistance drop c a l c u l a t e d then the s p e c i f i c r e s i s t a n c e can be c a l c u l a t e d . Once the s p e c i f i c r e s i s t a n c e i s c a l - - 25 - culated the immersion can be calculated for any total resistance (i.e. voltage and amperage reading). 2.3 Application of the Analytical Equation 2.3.1 Data Generated by the Equation The application of the model i s f a i r l y simple and is accomplished using any d i g i t a l computer. Using the computer the temperature can be c a l - culated for a series of points in the parent metal as a function of the position of the weld i t s e l f (i.e. the bottom of the electrode) and the thermal cycle for each point calculated. This d i g i t a l data can be then fed into a plotting device and a plot of the thermal cycle can be arrived at for any selected point. The data can also be contoured to give a mapping of the temperature f i e l d at any desired isothermal contour interval. The thermal cycle plots can be used to predict microstructural changes at each point with the help of Continuous Cooling Transformation (C.C.T.) curves i f an appropriate C.C.T. curve i s available. The contours of isotherms allow the penetration to be read directly from the map. 2.3.2 Manipulation of the Data Bastien et al (68) have characterised the entire mechanical properties of the HAZ of welds using three independent variables: 1. The austenitising parameter 2. The rate of cooling 3. The tempering parameter The cooling rates can be calculated using Rykalin's (69) or Adam's (70) formulae both of which are derived from Rosenthal's equation. Adam's Equation, which is - 26 - V = 2irk p C ( - ) 2 (T - T o ) 3 [2.12] c a l c u l a t e s the cooling rate from temperature T, f o r two dimensional heat flow and i s u s u a l l y applied to t h i n p l a t e s . However, due to the nature of ESW already described, i t can be applied to the thermal c y c l e i n ESW. This rate can also be ar r i v e d at by taking the slope at the appropriate point from the thermal cycle p l o t t e d from the a n a l y t i c a l model generated data. The a u s t e n t i s i n g and tempering parameters allow the d i r e c t use of the temperature versus time curves and d i f f u s i o n p r i n c i p l e s and are represented by a parameter P: F • •! ~:v' h^h ' t.2-w. where: T = temperature i n °K AH = a c t i v a t i o n heat f o r the p a r t i c u l a r phenomenon t = time to = time unit and i s expressed as an equivalent temperature f o r a thermal c y c l e f o r which the holding temperature equals the time u n i t to. The time, t , at holding i s c a l c u l a t e d by in t e g r a t i n g the area under the curve above a c e r t a i n temperature, T„ (850°C for austentising) and converting t h i s to a rectangular E thermal c y c l e of the same area and maximum temperature, T M. Now T M 2 . T M - T E = R M L 2 " 1 4 ] . R T M 2 c - 27 - The a c t i v a t i o n enthalpy has been found (71) to be 110 kcal/mole f o r a u s t e n i t i s a t i o n and 92.5 kcal/mole f o r tempering over a wide range of s t e e l s . We can now c a l c u l a t e P i n degrees K e l v i n and using Figure 14 determine the average a u s t e n i t i c grain s i z e . 2.4 Summary An a n a l y t i c a l equation has been presented which w i l l be used to p r e d i c t temperature-time c y c l e s . The parameters that can be v a r i e d have been i d e n t i f i e d and incorporated into the equation. The formulae for the c a l c u l a t i o n of the i n t e r e l e c t r o d e gap and austenite grain s i z e have been presented. Ikwever, other than f o r the a n a l y t i c a l model there i s no published data r e l a t i n g these theories to actual f i e l d or experimental r e s u l t s . This has been undertaken i n t h i s experimental program. The procedures used are developed i n Chapter 3, the r e s u l t s obtained i n Chapter 4 and the comparisons with the equation and c a l c u l a t i o n s i n Chapter 6. CHAPTER 3 EXPERIMENTAL PROGRAM 3.1 ESW Devices at U.B.C. 3.1.1 Wire or Consumable Guide Welders When welding i n the D.C. mode, power was supplied by a Hobart RC 750 welder with 750 amperes and operating voltage of 10-50 v o l t s . The welder was p a r t i a l l y c o n t r o l l e d by a rheostat which v a r i e s the degree of satu r a t i o n of i t s reactor. The r e c t i f i c a t i o n was not perfect and resulted i n approxi- mately 10% RMS of the t o t a l D.C. current. When t h i s power supply was used for wire or consumable guide electrodes, i t was used i n conjunction-with an Arcos Vertomatic GSp wire feeder and welding c o n t r o l l e r . This c o n t r o l l e r was a prototype and, as such, was never marketed i n the form employed at U.B.C., The c o n t r o l l e r was wired into the Hobart RC 750 i n such a way that the saturation of the reactor was var i e d at the c o n t r o l l e r . The only con- t r o l s used i n the experiments conducted at U.B.C. were the voltage and wire feeding c o n t r o l s . The voltage c o n t r o l e s s e n t i a l l y replaced the Hobart rheostat and the wire feeding c o n t r o l adjusted the amperage. The voltage and amperage can be read d i r e c t l y at the c o n t r o l panel. A Hobart Model T-500 AC power supply was used i n conjunction with the Arcos vertomatic wire feeder when welding i n the AC mode. In t h i s case both the voltage and current were measured e x t e r n a l l y , not at the c o n t r o l l e r . The voltage was adjusted using various tap connections with a f i n e control varying the primary voltage to the step down transformers (tap changes can be made on-load). 3.1.2 Bar or Plate Electrodes The U.B.C. electroslag remelting furnace was used for large cross- section electrodes. It i s designed to undertake a range of research pro- jects and has been described in detail by Etienne (72). It was used in the preliminary and f i n a l stages of the present research. 3.2 I n i t i a l Experimental Procedures 3.2.1 Stage One To determine i f there were significant differences in the three avail- able modes of operating, and to see at the outset i f ESW welds could be made using the U.B.C. remelting furnace, three welds were made, each using a different mode (D.C.R.P., D.C.S.P. and AC). These welds were made u t i l i z i n g a colo r l i t h plate (3/4" thick) between the base plate and a steel plate. The material to be welded was 5k," x 5 V x 12" in height (from Western Canada Steel continuous cast strand)AISI 1020 steel. The electrode was a 1 3/8" diameter AISI 1018 rod. Thus, for remelting, the water cooled copper mould was replaced with a solid b i l l e t of square cross section. The slag was 75 weight percent calcium fluoride (courtesy of Eldorado Nuclear Co.) and 25 weight percent alumina (Norton Abrasive Ltd.).. Twenty four chromel/alumel thermocouples were placed in a grid. This grid consisted of four different distances from the v e r t i c a l axis at six different heights. The thermocouple wires were sheathed in 1/8", double bore, alumina tubing and the end beaded by welding. Holes 3/16" in diameter were d r i l l e d into the blocks to four depths and the thermocouples inserted with pressure. The end of the holes were 1/4", 1/2", - 3 0 - 3/4" and 1 " from the surface of the a x i a l hole i n the h o r i z o n t a l plane ( F i g . 15). It required a p o s i t i v e force to i n s e r t the sheathed thermo- couples due to the non-round sheathing and, once the thermocouple was i n con- tact with the base metal, i t did not move. A multimeter, placed on res i s t a n c e , was used to ensure the thermocouple bead made contact with the base metal. This unconventional configuration was chosen f o r the following reasons: 1. There was very l i t t l e preparation required to determine i f the equipment set up was appl i c a b l e to welding as w e l l as remelting. 2. The blocks completely removed the problems associated with water cooled copper shoes or moulds. 3. There was no slag l o s s , because there were no shoes, and con- sequently no sla g had to be added during the operation. 4 . The slag had been used s u c c e s s f u l l y by others to remelt s t e e l s with r e l a t i v e ease (72,73). 5. I t should show the main d i f f e r e n c e s between the temperature time d i s t r i b u t i o n s f o r the three modes of welding. 6. The set-up would allow t e s t i n g of the thermocouple placement technique. The weld was started cold using a compact c o n s i s t i n g of metal turnings of AISI 1018 and calcium f l u o r i d e . The compacts were approximately 40 mm i n height and contained 100 gms of calcium f l u o r i d e . The electrode i s forced down onto the compact and the granulated, pre-fluxed s l a g i s poured around the electrode and compact. The compact melts, due to i t s r e s i s t a n c e , when the power i s turned on and the process commences with a molten s l a g . The voltage and amperage were measured and recorded using Sargent Model SR 10 M i l l i v o l t recorders (voltage required a 100:1 d i v i d e r and amperage a c a l i b r a t e d shunt) when the operating mode was d i r e c t current. In order to - 31 - operate i n the A.C. mode,it was necessary to s t a r t using D.C. and then change power supplies once the process was s t a b i l i z e d . The voltage was measured using an RMS AC Voltmeter (model Eico 250) and recorded manually. The amperage was measured using an AC to DC convertor and recorded using a Sargent Model SR M i l l i v o l t recorder. The thermocouple output emf. was recorded using a Texas Instrument Multi- p o i n t Recorder set on i t s f a s t e s t speed (one reading per s e c ) . The other process parameters recorded were: 1. The t o t a l electrode t r a v e l from s t a r t 2. The time from a f i x e d point 3. The electrode d r i v e motor speed s e t t i n g 3.2.2 Stage Two, Consumable Guide ESW Stage two of the experimental setup was designed to research the ESW parameters for the Consumable Guide process. The s i n g l e wire non-consumable guide technique xjas not studied because of the complexity of t r a v e l l i n g c o o l i ng shoes on the r e l a t i v e l y short welds produced i n our laboratory. Instead i t was decided to study the process without the problems associated with the complicated heat sink created by a slag skin and water c o o l i n g . The welds were produced i n 5k" diameter, 24" long AISI 1020 and 4340 s t e e l c y l i n d e r s with a l h " diameter a x i a l l y bored hole. The electrode wire was AISI 1010 and 1040 (2.4 mm and 2.5 mm diameter r e s p e c t i v e l y ) . The slags were Arcos BV8 (an e l e c t r o s l a g f l u x : 42 CaF2» 26 A^O^, 24 CaO, 8 Si02) and Hobart PF 201 (an e l e c t r o s l a g f l u x : 10 A l ^ , 5 MgO, 35 CaO + CaF 2, 15 MnO, 30 S i 0 2 ) • The wire was fed through a nominal 1/8" mild s t e e l pipe which was used as the consumable electrode. Two methods were used to i n - sulate t h i s consumable guide from the sides of the hole and to keep the electrode c e n t r a l i n the hole: 1. S i l i c a tubes approximately 3/4" long were f l a r e d to give approxi- mately 180° included angle. A wire was wrapped around the pipe and f i x e d by 3 very small tack welds. The s i l i c a tube was then slipped over the pipe and the f l a r e d flange l a y on the wire. A 1/16" thick f l a t , six-pointed mild s t e e l d i s k was then placed on top of the f l a r e . The s i l i c a acted as an i n s u l a t o r and the di s k as a po s i t i o n e r ( F i g . 16). 2. The s i l i c a i n s u l a t o r s were replaced by an i n s u l a t e d guide f o r the top 8" of the pipe ( F i g . 17). This was found to be superior to the s i l i c a guides which tended to break. This guide also allowed the use of the v e r t i c a l t r a v e l of the wire feeder i t s e l f . Thus, i n the event of the wire fusing with the pipe, the consumable guide was lowered and became the a c t i v e electrode u n t i l the wire once again started feeding. The thermal cycles were monitored using 14 i n t e r n a l thermocouples at seven d i f f e r e n t depths at 1/4" i n t e r v a l s s t a r t i n g 1/8" from the hole surface. Six thermocouples were placed on the surface of the c y l i n d e r s by spot welding at p o s i t i o n s between the i n t e r n a l thermocouples. This thermocouple set-up i s shown i n Figure 18, and was used f o r the f i r s t 10 experiments. The remainder of the runs were monitored using 24 i n t e r n a l thermocouples as shown i n Figure 19. The weld was started by adjusting the power rheostat at the highest p o s s i b l e reading to provide the highest voltage. The wire advance rate was kept very low and as soon as the wire started to arc some sl a g was added to extinguish the arc. This procedure was continued u n t i l the wire advanced without arcing any f u r t h e r , at t h i s point the power was reduced u n t i l the required voltage was j u s t exceeded. The wire ra t e was then increased u n t i l - 3 3 - the desired current was attained. During t h i s period the desired amount of- slag was added. The e n t i r e start-up procedure was completed before the molten metal was" above the top of the run-in sump. During the remainder of the weld, the voltage and amperage were kept as close as po s s i b l e to the desired values. The weld was stopped once the molten s l a g surface was even with the top of the c y l i n d e r . 3.2.3 Slag Temperature Measurements The slag temperatures were measured during the run using a W3Re/W25Re thermocouple with a boron n i t r i d e sheathing(Fig. 20). The thermocouple was held i n two eye-loops and, to prevent thermal shock, the molten slag was allowed '. to approach the sheathed end at the welding v e l o c i t y . Once the sla g was touched the thermocouple was lowered at 10 mm i n t e r v a l s u n t i l the sheathing was destroyed and the thermocouple ceased working. The thermo- couple emf. was recorded on a Sargent Model SR4 M i l l i v o l t recorder. 3.2.4 Electrode Immersion The immersion of the electrode was measured using a rather simple mechanical means and was only possible when the f i x e d head (with insulated holes) was i n use. Two s t r a i g h t tungsten wires, each with a c i r c u l a r loop at r i g h t angles ( F i g . 21) were placed over the nozzle and moved to the top of the c y l i n d e r and held by the top f i x i n g head. When the sla g surface approached the l e v e l of the lowest thermocouple p o s i t i o n , the lower loop was lowered down the nozzle and over the extending wire with a small l a t e r a l force. As soon as.the loop moved sideways, i t was immediately withdrawn and held at the f i x e d head by bending i t over. The second loop was lowered a f t e r the uppermost thermocouple p o s i t i o n had been passed by the slag metal pool i n t e r f a c e . The same procedure was - .34 _ used for the second loop as for the f i r s t . The electrode immersion was assumed to be the same as the length of wetted (slag coated) tungsten wire. 3.2.5 Other Experimental Procedures The welding velocity was calculated by recording the time for the entire length of the f i n a l weld. However, because the exact point of the welding start was never actually.known, i t was decided to add tungsten powder before the slag/air interface reached the bottom thermocouples and after the molten metal/metal interface passed the last thermocouple. The times of addition were recorded and the position determined after sectioning of the deposited metal. In some instances 304 S.S. 1/8" pipe consumable nozzles were used instead of the mild steel pipe commonly used. These experiments were sectioned and 4% n i t a l etchs were used to determine i f the stainless components were evenly distributed during the welding process. 3.3 Butt Welds using Consumable Guide ESW Once the parameters and procedures involving the cylinders were f a i r l y well understood, i t was decided to produce f u l l scale production type welds using the Arcos Vertomatic GSP wire feeder, Hobart A.C. and D.C. welders and water.cooled.copper shoes. This decision led to many changes i n the equipment and procedures. It necessitated the design of a welding j i g , water cooled copper shoes, a fixed head for the consumable guide, run-in and run-out sumps and some method for adding slag. The essential differences between the cylinders and I V plates are: - 35 - 1. There i s heat l o s t to the moulding shoes and t h i s must be measured and accounted for i n the power input. 2. There i s a continual l o s s of slag to the faces of the cooling shoes and t h i s slag must be replenished at frequent equal i n t e r - v a l s or at a constant r a t e . 3.3.1 The Welding J i g The welding j i g was designed to perform the f o l l o w i n g f u n c t i o n s : 1. The j i g would hold the pla t e s v e r t i c a l so that there was no play or movement perpendicular to the pla t e s and at the same time allow a s l i g h t movement p a r a l l e l to the p l a t e s . 2. The j i g must have the f i x e d head attached to i t so that, by a l i g n i n g the head and j i g once, the consumable guide would remain i n the center of the gap. 3. The j i g would hold the water cooled copper moulding shoes as t i g h t l y against.the plates as f e a s i b l e without d i s t o r t i n g the shoes. The j i g was designed and constructed out of mild s t e e l as shown i n Figure 22. The twelve screws i n the v e r t i c a l members were rounded on the contact ends to provide movement of the plates i n the plane of the p l a t e s . The bottom of each p l a t e was shimmed to provide a v e r t i c a l gap. The top of the v e r t i c a l members were d r i l l e d and threaded to hold the f i x e d head. The j i g was elevated with two I-beams to accommodate the run-in sump. Two sets of h o r i - zontal clamps were used to hold the copper shoes i n alignment throughout the weld. Steel bars were sometimes used to back up the copper shoes when the pla t e s were not f l a t ; t h i s allowed the clamping screws to be torqued very high and forced the shoes to conform.to any bend i n the p l a t e s . This j i g , with no changes, was used for the remainder of the experiments. - 36 - 3.3.2 The Water Cooled Copper Shoes There are three types of water cooled copper shoes commercially used: 1. The moving shoe which i s approximately 8-12" long and 3" wide with water channels through a s o l i d block. These shoes move up both sides of the plate at the welding v e l o c i t y and r e t a i n the molten metal and slag. 2. The walking shoe i s any length from 8" to 36". The walking shoe can be designed i n a number of ways to accommodate the water flow but i s u s u a l l y formed from a s o l i d block of copper with m i l l e d out water channels and a facing p l a t e . The water enters the bottom and e x i t s from the top of the s h o e The walking shoe method requires four shoes, two f o r each side. The procedure involves p l a c i n g two shoes at the s t a r t of the weld and welding u n t i l the s l a g / a i r face i s a few minutes from the top of the shoe. The second set of shoes i s then attached above the f i r s t set and, when the weld i s i n the second set, the f i r s t set i s removed and placed above the second set, thus the s o - c a l l e d walking shoe method. 3. F u l l length shoes are also used i n circumstances which warrant a large c a p i t a l investment. They are u s u a l l y used when the weld i s not v e r t i c a l , not s t r a i g h t or has s e c t i o n changes throughout i t s length. I t can also be used when many welds of exactly the same con f i g u r a t i o n are being made. The f u l l length shoe (24") was used for the f o l l o w i n g reasons: 1. The problems associated with the moving shoe and wire feeder were immediately eliminated. 2. The procedure of walking shoes i s time consuming, requires more materials and cooling connections than the s i n g l e shoe. - 37 - 3. The length of the welds i n the present program was constant (22 i n . ) and short enough to use one set of shoes. (The set of shoes could be considered the f i r s t set of shoes i n the walking shoe method). The shoes were designed as shown i n Figure 23. A s o l i d t h i c k , 1 3/4" wide and 22" long was m i l l e d 3/8" deep making a channel l h " wide, 21h" long. A water i n l e t hole was d r i l l e d at the bottom and an o u t l e t hole at the top and 1/4" nipples welded onto the back of the shoe. The channel was then closed by h e l i a r c welding a 1/8" t h i c k p l a t e to the channel and the shoe pressure tested. The 1/8" t h i c k . p l a t e i s the cooling surface and had to be m i l l e d to contain the frozen slag. Therefore the m i l l e d channel was 1/16" (1.6 mm) deep, 1 3/8" wide and 21 3/4" long. This allowed f o r a slag skin 1/16" t h i c k and 1/16" wider than the 1 1/4" wide root gap and ensured the weld deposit would be at l e a s t as t h i c k and wide as the parent metal. I t also allowed for 3/16" contact width with the p l a t e s on each side of the weld. It would have been preferable.to have the l/2"-5/8" contact width that i s used commercially but t h i s would have i n t e r f e r e d with the placement of the thermocouples. I t should be noted that the shoes did wear out and had to be replaced with a second set. This set was made using 1 3/4" x 1/2" s t r u c t u r a l grade channel instead of a copper channel. The f a c i n g p l a t e was the same as before because the thermal conductivity of copper i s essen- t i a l f o r the required heat t r a n s f e r . The new set was used f o r the remainder of the experiments and performed as w e l l as the e n t i r e l y copper shoes. The shoes were cooled using a.mixture of hot and col d water to provide 25°C water, which would not condense water vapour from the a i r . The flow of t h i s water was controlled.using a rotameter and the temperature measured using thermistors on the i n l e t and o u t l e t water flows. - 38 - 3.3.3 The Consumable Guide Fixed Head The f i x e d head and j i g were r e a l l y one s o l i d component, once the head was attached to the j i g . I t performed the following functions: 1. I t held the guide so that there was v i r t u a l l y no movement of the guide, and e l e c t r i c a l l y i n s u l a t e d the guide from the parent metal as shown i n F i g . 24. 2. I t contained other guides for slag temperature measurement and electrode immersion measurements. 3. I t held the tube f o r the a d d i t i o n of tungsten powder and make-up sl a g . 4. I t also allowed for the v e r t i c a l movement of the consumable guide i f the welding x^ire became fused to the guide. The head was e l e c t r i c a l l y i n s u l a t e d from the j i g and parent metal by a 1/2" p l a t e of c o l o r l i t h , c o l o r l i t h sleeves f o r the b o l t holes, and rubber washers under the s t e e l washers. The four v e r t i c a l posts were machined with two steps to hold the h o r i z o n t a l guide p l a t e s . The bottom guide p l a t e had four holes with exactly the same diameter of the middle section of the v e r t i c a l posts. This provided the pivot point f or the a l i g n i n g of the con- sumable guide with the rectangular section to be welded. The top guide p l a t e had four holes 1/16" larger i n diameter than the diameter of the top s e c t i o n of the v e r t i c a l posts. This play allowed movement of the top guide p l a t e p a r a l l e l and perpendicular to the plates and thus the a b i l i t y to center the electrode and guide with respect to the bottom of the p l a t e s . T h i s , i n conjunction with the j i g screws,provided for alignment of the guide along the c e n t e r l i n e of the e n t i r e gap between the shoes and p l a t e s . - 3 9 - 3.3.4 Run-in Sump and Run-out Blocks The run-in sump was constructed out of 1 1/4" square bars providing a sump 2" deep. The sump was constructed t h i s way to ensure the penetration of the-parent metal before the slag metal i n t e r f a c e reached the ac t u a l blocks. The 1 1/4" thickness was used because.it had adequate strength to hold the plates and yet was t h i n enough to provide a. poor heat sink and allow over- heating. The overheating ensured penetration of the base metal i n the sump and at the s t a r t of the weld. The sump was f a b r i c a t e d by p a r t i a l pene- t r a t i o n , manual arc welding; i t was ground o f f so the shoe faces would f i t the sump faces and prevent sl a g run-outs. The run-out blocks were 1 1/4" square and simply tack welded to the top of.the plates to provide a sump to contain the sla g a f t e r welding was com- pl e t e . They were only 2" high to ensure the weld ended i n the p l a t e and to permit observation of the metal/slag i n t e r f a c e a f t e r welding. In actu a l p r a c t i c e they are 3-4" high to ensure the weld i s complete above the p l a t e s . 3.3.5 Thermocouple P o s i t i o n s . Chromel/alumel thermocouples were inserted i n t o the plate s as shown i n Figure 25. There were two groups of thermocouples used, with 20 thermo- couples i n each group (sometimes only one b l o c k ) . Each group consisted of f i v e rows ...of . thermocouples at. four depths. The four depths were 1/4", 1/2", 3/4" and 1" from the o r i g i n a l surface of the p l a t e and 5/8" deep in t o the p l a t e . F i v e thermocouples at each depth were used to e s t a b l i s h i f the welding was i n steady state ( i . e . i f a l l f i v e reached the same m i l l i v o l t readings, then steady state was achieved). The m i l l i v o l t output was r e - corded .on a p r i n t i n g HP Model 2070 A Datalogger with a v i s u a l d i g i t a l d i s p l a y . This device allowed quick, accurate c o l l e c t i o n of data. The - 40 - lowest, c l o s e s t thermocouple i n the group was recorded continuously (using a Sargent Model SR 4) as a reference f o r a l l the other thermocouples. The second block of thermocouples was used f o r two a p p l i c a t i o n s during the t e s t - work. . F i r s t l y , i f the f i r s t set of thermocouples was found to be i n steady state .throughout the range, the second group could be used to see how "steady" the process r e a l l y was over a long period of time. Secondly, the second set could.be used for a second experiment i f the f i r s t group was i n steady state. Therefore, the welding parameters could be changed between thermocouple groups and a second weld achieved from the same set-up. This second method was used as often as po s s i b l e to provide the most data p o s s i b l e from one experiment. It should be noted that the change i n the groups meant that the cooling c y c l e was not completely recorded f o r many of the uppermost thermocouples, and t h i s . n e c e s s i t a t e d using the cooling cycle f o r the f i r s t one, two or three thermocouples at each depth to obtain the e n t i r e thermal c y c l e . A l s o , the f i r s t thermocouple was sometimes found to be not i n the steady s t a t e condition when the other four were (the f i r s t always read low i f t h i s happened) and then the readings from t h i s f i r s t thermocouple were disregarded. 3 . 3 . 6 Slag Additions and Other Procedures Appropriate slag additions were made i n 5 gm amounts at i n t e r v a l s de- pending upon the welding rate. Tungsten powder ad d i t i o n s were made at the s t a r t and f i n i s h of the thermocouple readings f o r each block. Both additions were made using the slag a d d i t i o n tube on the f i x e d head. 3 . 3 . 7 E l e c t r i c a l Current D i s t r i b u t i o n Measurements E l e c t r i c a l current measurements were made by p l a c i n g p l a t e s i n the parent metal and e l e c t r i c a l l y - i n s u l a t i n g these pla t e s from each other and the shoes by the use of asbestos sheets. These pla t e s were 1/4" thick, - 41 - 1 1/8" wide and 2 inches longer than the parent metal p l a t e s Csee F i g u r e 26). A shunt was bolted to the end of each p l a t e and the other end of the shunt to the same side of the welder.as the parent metal. The p r e c i s e l o c a t i o n of each shunt was recorded with respect to a thermocouple i n the block opposite the shunt. The m i l l i v o l t drop across each shunt was continuously monitored using-Sargent Model SR 4 and 10 recorders as the weld passed the shunt. Thus the current d i s t r i b u t i o n was recorded as the weld passed the shunt. This procedure was only possible while welding i n the D.C. mode. When welding i n the A.C. mode, only one pl a t e was used with one shunt and the. A.C. m i l l i - v o l t drop was converted to D.C. using a HP Model 8807 A A.C. to D.C. con- v e r t e r - a m p l i f i e r . Only one plate was used because only one converter was a v a i l a b l e . However, the moving current d i s t r i b u t i o n was measured and the r e l a t i v e p o s i t i o n of the shunt and i n t e r f a c e s c a l c u l a t e d . 3.4 Butt Welds Using Bar Electrodes Once.consumable guide electrode welding was completed, i t was decided to study the large or p l a t e electrode method of ESW. This n e c e s s i t a t e d few changes-from the consumable guide process and, with the exception of the electrode feed mechanism, f i x e d electrode guide and s l a g a d d i t i o n method, no changes were made. The electrodes f o r the majority of the welds were 3/4" square leaving a gap of 1/4" on a l l four sides of the electrode f o r m i s a l i g n - ment or d e v i a t i o n during electrode t r a v e l . The electrodes were f a b r i c a t e d by c u t t i n g 3/4" slabs o f f the parent metal sections and c u t t i n g the slabs to 3/4" x 3/4" x 18". The ends of these bars were bored and tapped to receive a 1/2"^N.C. b o l t and.joined together. Three and one-half 18" electrodes were joined using .1/2" b o l t s to provide the electrode f o r one weld. The welds were started using calcium f l u o r i d e Annco i r o n compacts that had been _ 42 _ shaped to f i t the 1 1/4" square gap. The electrode feed mechanism was the mechanism of the e l e c t r o s l a g r e f i n i n g u n i t described by Etienne (72).. The slag was continuously added using the e l e c t r o s l a g r e f i n i n g slag; feeder described by J o s h i (74) and r e c a l i b r a t e d f o r the appropriate e l e c t r o s l a g sl a g . 3.4.1 The Fixed Head The most extensive change required to weld using large electrodes was the welding head. Both h o r i z o n t a l p l a t e s had to be modified to accommodate the l a r g e r electrode. The head was r e b u i l t so that the upper p l a t e became the pivot point and the electrode d r i v e the guide point. Alignment was most important during large electrode welding because the gap was very narrow between the electrode and parent metal. The a c t u a l guides were made out of c o l o r l i t h to allow f o r the electrodes not being p e r f e c t l y square and not the same s i z e f o r each section of the electrode. E s s e n t i a l l y the f i x e d guide serves the same purpose as i n consumable guide ESW. 3.5 Evaluation of Weld Properties 3.5.1 Hardness Values Hardness traverses were performed using a Vickers Pyramid Hardness Testing Machine. Transverse sections were cut and surface ground so that the transverse sides were p a r a l l e l . One side was polished to a 1 u f i n i s h and l i g h t l y etched using 2% n i t a l to d i s t i n g u i s h the f u s i o n boundary and HAZ. The hardness readings were taken at 1/8" i n t e r v a l s from the c e n t e r l i n e of the weld i n t o the parent metal. The majority of the traverses were per- formed using a 10 kilogram load and a 2/3 o b j e c t i v e l e n s . - 43 - 3.5.2 Toughness Values A l i m i t e d s e r i e s of Charpy Impact te s t s were performed on p l a t e welded cast T - l equivalent s t e e l , (ASTM 487 (Q7)) provided by Esco L t d . These t e s t s were made using Warnock Hersey Internationals Charpy t e s t i n g machine and i n accordance with ASTM-370 at 40°F. The specimens were taken so that the 2 mm V-notch was. located at 1/4" In t e r v a l s from the weld c e n t e r l i n e i n t o the parent metal. Only AC and DCRP were toughness tested. 3.5.3 In c l u s i o n D i s t r i b u t i o n s Transverse samples were taken across the weld i n 1/4" cubes f o r oxygen a n a l y s i s using a d i g i t a l readout Leco oxygen analyser. Polished sections were taken adjacent to the samples for oxygen a n a l y s i s and the i n c l u s i o n d i s t r i b u t i o n determined using the Ouantimet with the microscope attachment. This procedure was used on two welds f o r each mode of welding. 3.6 Experimental Records The f o l l o w i n g parameters were recorded f o r each run when a p p l i c a b l e : 1. Welding mode; AC, DCRP, DCSP 2. Weld type; block, c y l i n d e r , p l a t e 3. Electrode type; consumable guide, p l a t e 4. M a t e r i a l s ; s l a g , electrode, parent metal 5. Voltage, recorder 6. Amperage, recorder 7. Power s e t t i n g 8. Slag a d d i t i o n rates 9. Temperatures, slag and parent metal - 44 - 10. Time, recorder 11. Tungsten a d d i t i o n times 12. Immersion times and depths 13. Time slag temperatures taken 14. Cooling water flow rate and temperature change 15. Slag skin thickness 16. F i n a l Slag depth 17. Length of weld 18. Operational problems CHAPTER 4 . RESULTS 4.1 Power C h a r a c t e r i s t i c s 4.1.1 Published Power C h a r a c t e r i s t i c s Figures 27 and 28 r e s p e c t i v e l y show the published voltage amperage c h a r a c t e r i s t i c s for D.C. and A.C. welders used i n t h i s research. Figure 29 shows the v a r i a b l e resistance device constructed to measure the voltage amperage c h a r a c t e r i s t i c s of the welders; these appear as the l i n e s through the dots on Figure 27 and the lower voltage curves on. Figure 28. 4.1.2 Measured Power C h a r a c t e r i s t i c s , The voltage amperage c h a r a c t e r i s t i c s i n d i c a t e the working range within-which the welders would have to be operated to a t t a i n the required input energy e s s e n t i a l f o r a sound weld. In both cases, A.C. and D.C, the welders required much higher settings than the published data by the manufacturer indicated. _ 46 _ 4.2 I n d u s t r i a l Experiments 4.2.1 Experimental Conditions . A s e r i e s of welds were made at Canron Limited using the voltage amperage settings shown i n Table 4. The other f o l l o w i n g conditions used were: i ) D.C.R.P. i i ) 3.18 cm root gap I i i ) A 441 Grade C hot r o l l e d p l a t e (see Table 5) i v ) Hobart'PF 201 f l u x (3.81 cm operating depth) v) HB 25P wire (see Table 5) v i ) 1.27 cm diameter consumable guide 4.2.2 Experimental Results The weld regions were examined f o r the i n c l u s i o n d i s t r i b u t i o n ( F i g . 30), oxygen d i s t r i b u t i o n and penetration (Figs. 31 and 32), and HAZ (Figure 33). The r e s u l t s were used to p r e d i c t the actu a l weld v e l o c i t y required to pro- duce an acceptable weld i n the laboratory, using d i f f e r e n t weld c r o s s e c t i a n a l areas (Figure 34). These data were used i n l a t e r wire welding techniques but were a l s o used to determine the melt rate-energy requirements f o r the large electrode experiments. 4.3 I n i t i a l Large Electrode Experiments 4.3.1 Heat Balance and Thermal P r o f i l e s The thermal p r o f i l e (verus p o s i t i o n ) for the three modes of operation are shown-in Figures.35-37. TABLE 4 Canron Limited Welding Conditions Thickness . of P l a t e i n . Root Opening (in.) . Electrode Type ^ Arc Voltage O s c i l l a t i o n Distance in/min Wire Feed Speed Inch per min Approx. Amperage 3/4 1 1/4 E70S3 3/32 38-39 - . - 100 400 1 1 1/4 E70S3 3/32 38-39 - - 110 450 11/4 1 1/4 E70S3 3/32 38-39 - 120 500 1 1/2 1 1/4 E70S3 3/32 39-41 - - 130 500 1 3/4 1 1/4 E70S3 3/32 39-41 . 1 1/2 35 140 525 2 1 1/4 E70S3 3/32 39-41 1 1/2 35 150 550 2 1/4 1 1/4 E70S3 3/32 41-44 1 1/2 35 . 1 6 0 600 2 3/8 1 1/4 E70S3 3/32 41-44 1 1/2 35 165 650 2 1/2 1 1/4 E70S3 3/32 41-44 1 1/2 35 170 700 2 3/4 1 1/4 . E70S3 3/32 41-44 1 1/2 35 .180 700 3 1 1/4 E70S3. 3/32 41-44 1 1/2 35 190 750 - 45 - TABLE 5 Canron Limited Welding Analysis BASE METAL SPECIFICATION AND COMPOSITION A441 to A441 Base Metal Carbon Manganese Silicon Cu V P. S. A441 0.20 1.10 0.023 0.26 0.06 0.011 0.019 FILLER METAL COMPOSITION E70S-3 (Hobart HB-25P) Typical Type Carbon Manganese Silicon P.Max. S.Max. HB-25P 0.11 1.20 .50 0.020 0.019 DEPOSITED METAL ANALYSIS Hobart PF201 Flux Type Carbon Manganese Silicon Cu V P. S. HB-25P 0.14 1.11 0.45 0.20 0.03 .008 .019 Figure 38 compares the thermal p r o f i l e s f o r DCRP, DCSP and AC f o r a thermo- couple positioned 1/4 inch from the o r i g i n a l surface of the weld. 4.3.2 Power C h a r a c t e r i s t i c s and Apparent Resistance of the Slag The apparent res i s t a n c e of the s l a g was found to be: Mode Vo l t s Amps R(ft) DCRP DCSP AC 23 ± .4 23 +•• .9 26.5 + .1 1135 ± 152 1107 ± 100 1387 ± 45 .0229 .0169 .0285 It should be noted that the A.C. experiments using the U.B.C. E l e c t r o s l a g Remelting furnace must be started using D.C.S.P. u n t i l the sl a g i s e n t i r e l y molten and thus the apparent res i s t a n c e f o r the A.C. experiment could be aff e c t e d by the e l e c t r o l y t i c reactions during the D.C. p o r t i o n of the ex- periment. The s p e c i f i c slag r e s i s t a n c e can not be c a l c u l a t e d because of the r e l a t i v e l y l a r ge electrode diameter, low current density and a c o n i c a l shaped melting electrode. 4.3.3 Deposition Rate or Welding V e l o c i t y The deposition rates and welding v e l o c i t i e s were: melt rate (g.s.~l) welding v e l o c i t y (cm.sec DCRP 4.28 .018 DCSP 4.51 .019 AC 7.13 .030 with corresponding power inputs i n terms of Kcal.gm of: DCRP 1.49 Kcal gm DCSP 1.24 Kcal gm".. AC 0.81 Kcal gm and Kcal.cm of weld: 353 Kcal cm_1 r e s p e c t i v e l y 293 Kcal cm 193 Kcal cm - 50 - 4.3.4 P e n e t r a t i o n and I n c l u s i o n D i s t r i b u t i o n . The p e n e t r a t i o n s , measured as the maximum p l u s the minimum d i v i d e d by two, minus the o r i g i n a l h o l e diameter were found t o be: mode penetration, (cm) DCRP 4.53 cm DCSP 1.06 cm AC 0.91 cm F i g u r e s 39-41 show the i n c l u s i o n d i s t r i b u t i o n f o r c r o s s e c t i o n s of the above three welds. 4.3.5 Hardness Traverses F i g u r e 42 presents the hardness t r a v e r s e s through the same s e c t i o n s from which the oxygen and i n c l u s i o n analyses were obtained. The measured p e n e t r a t i o n i s a l s o shown. (R.G. i s the o r i g i n a l r o o t gap.) 4.3.6 Oxygen Analyses F i g u r e 43 represents the oxygen g r a d i e n t along the same t r a v e r s e s used f o r the hardness surveys. 4.4 Consumable Guide ESW i n C y l i n d e r s 4.4.1 I n t r o d u c t i o n A t o t a l of t h i r t e e n s u c c e s s f u l experiments were performed u s i n g the consumable guide method des c r i b e d i n Chapter 3. Table 6 presents the data obtained or c a l c u l a t e d f o r the experimental c o n d i t i o n s . 4.4.2 Slag Temperature Measurements The s l a g temperature was measured twice f o r each D.C. mode and once f o r the A.C. mode. The temperatures are p l o t t e d w i t h the corresponding weld p r o f i l e , e l e c t r o d e immersion and i n t e r e l e c t r o d e gap i n F i g u r e s 44-49. TABLE 6 Experimental Conditions f o r Consumable Guide E l e c t r o s l a g Welding (CGESW) i n Cylinders Experiment Number Operating Mode Volts Amps Penetration cm Weld V e l o c i t y cm s e c - l Slag Depth cm Immersion cm Apparent Resistance ohms S p e c i f i c Power kcal/cm 1 DCRP 40 370 .159 .0351 3.18 1.74 .108 100 2 DCRP 41 510 .079 .0361 8.26 2.63 .080 140 5 DCRP 40 460 .318 .0303 5.84 2.35 .087 147 10 DCRP 52 435 ,079 ,0330 6.60 1,49 .120 • . 165 11 DCRP 40 460 .300 ,0301 3.81 2.35 .087 150 12 DCRP 38 400 ,406 .0357 4.45 2.08 ,095 115- ". 3 DCSP 41 460 ,318 .0328 5.08 2.01 .089 140 4 DCSP 40 390 .000 .0305 5.08 . 1.64 .103 125 6 DCSP 41 440 .238 .0252 6.35 1.89 .093 170 7 DCSP 40 540 ,040 .0254 6.10 2.59 .074 205 8 DCSP 38 600 .000 ,0339 6.60 3,20 ,063 160 9 DCSP 40 500 .000 .0328 7.10 2,33 ,080 145 14 AC 36 843 .238 .0357 7.00 4.89 .043 205 - 52 - 4.4.3 Slag Depth, Immersion and Interelectrode Gap Measurements Table 7 shows the slag depth and immersion as measured p h y s i c a l l y during and a f t e r the experiment (above and below the thermocouple region) using the tungsten wire method. The same ta b l e gives the measured electrode immersion and c a l c u l a t e d i n t e r e l e c t r o d e gap distance. 4.4.4 V e l o c i t y Measurements The welding v e l o c i t y was measured using three d i f f e r e n t methods: i ) the t o t a l length of the weld i n cm d i v i d e d by the t o t a l time of welding i n sec. i i ) the tungsten marker bed above and below the thermocouple region i i i ) the time between temperature maxima on thermocouples as recorded by the Hewlett Packard DVM output. These r e s u l t s are presented i n Table 8. 4.4.5 Temnerature Time Cycles The thermal p r o f i l e s f o r each mode of operation are presented i n Figures 50-52. The thermocouple arrangement allowed f o r averaging the pro- f i l e s because four thermocouples were a v a i l a b l e f o r each p o s i t i o n only i n the l a t e r experiments. A new thermocouple set-up described i n Chapter3 measured four l e v e l s at s i x depths. Only the average thermal p r o f i l e i s p l o t t e d i n Figures 50-52. There i s no comparison of the a c t u a l p r o f i l e s with any t h e o r e t i c a l model because t h i s experimentation was designed to determine the d i f f e r e n c e s between the modes f o r a c t u a l welding during the l a t e r stages of the research program. TABLE 7 Slag Depth, Immersion and Interelectrode Gap Measurement Experiment Number A f t e r Run Slag Depth (cm) Below Above Thermocouple Ave Immersion (cm) Measured Interelectrode (cm) . Gap Calc. 1 3.00 3.30 3.05 3.18 1.80 1.77 2 8.10 8.40 8.15 8.28 2.80 5.42 5 5.70 5.90 5.75 5.84 2.50 3.34 10 — 6.65 6:. 55 6.60 1.65 4.95 11 3.40 3.80 3.82 3,81 2,55 1.26 12 4.20 4.60 4>30 4.45 2.20 2.25 3 4.80 5.15 5.00 5.08 , 2.50 2.58 4 4.85 5.15 5.00 5.08 1.85 3.23 6 6.15 6.45 6.25 6.35 2.05 4.30 7 5.95 6.30 5.90 6.10 2.65 3.45 8 6.20 6.70 6.50 6.60 3.35 3.25 9 6.90 7.20 7,00 7.10 2,50 4,60 14 6.65 7.20 6.80 7,00 5,00 2,00 CO - 54 - TABLE 8 Welding V e l o c i t y Measurements . T o t a l Weld Tungsten Time Between „ , Length Marker Bed Temp. Maxima Number _ i l _ t cm. sec x cm. sec - 1- cm.sec x 1 .0351 - 2 .0361 5 .0303 - .0295 10 .0330 .0330 .0330 11 .0301 - .0288 12 .0357 .0355 .0308 3 .0328 - 4 .0305 - .0300 6 .0303 - .0298 7 .0254 - . .0256 8 .0339 .0340 .0335 9 .0328 - .0314 14 .0357 .0355 .0355 - 55 - 4.4.6 Penetration The penetration depth, or f u s i o n boundary p o s i t i o n , was determined op- t i c a l l y by d i r e c t measurement a f t e r macroetching. The penetration was also determined using changes i n VPN and microstructure. F i g u r e 53 presents t y p i c a l VPN traverses f o r each mode and d i f f e r e n c e between 4340 and m i l d s t e e l . Figure 53 also presents the f u s i o n boundary and HAZ p o s i t i o n s . 4.4.7 Oxygen Analyses Figure 54 presents the oxygen traverses from the weld c e n t e r l i n e through the fusion boundary for the three modes of welding. 4.4.8 Use of a 304 S.S. Consumable Guide Figure 55 shows the n i t a l etched surface of an experiment where the consumable guide was a 304 S.S. 1/8" pipe instead of the standard mild s t e e l 1/8" pipe. 4.5 Consumable Guide ESW with A 36 Plates 4.5.1 Introduction A rotameter flowmeter was c a l i b r a t e d to give a consistent 15 l i t e r s of water per minute to each cooling shoe i n order to produce a uniform slag thickness of approximately 1.0 mm. This was achieved i n a preliminary set of experiments designed to confirm the general value of the welding set-up. Table 9 presents the a c t u a l and c a l c u l a t e d experimental conditions for seventeen experiments. The slag feeder was also c a l i b r a t e d to d e l i v e r the necessary m a t e r i a l to replace that l o s t to the s l a g s k i n , and thus maintain a constant slag depth throughout the experiment (see Figure 56). TABLE 9 Experimental Conditions for Consumable Guide E l e c t r o s l a g Welding (CGESW) with A 36 Plates Experiment Number Operating Mode Volts Amps Penetration cm Weld V e l o c i t y cm sec Slag Depth cm Immersion cm Apparent Resistance ohms S p e c i f i c Power kcal.cm -! P13 DCRP 32 685 .381 .0159 1.59 1.42 .047 340 P14 DCRP 36 760 .669 .0232 1.91 1.11 .047 215 P15 DCRP 35 690 .191 .0205 2.06 1.59 .051 280 P18 DCRP 38 750 .254 .0191 1.50 1.27 .051 360 P20 DCRP 39 570 .320 .0174 3,65 6.79 ,068 310 P21 DCRP 39 565 ,318 ,0189 3,96 0.79 ,069 280 P16 DCSP 36 735 .064 .0223 3,18 1.62 ,049 270 P22 DCSP 40 590 .254 .0200 3.18 0.95 ,068 285 P04 AC 35 450 .325 .0166 2.86 1,10 .078 230 P05 AC 42 490 ,445 .0140 3.18 1.03 .086 350 P06 AC 40 520 . .191 .0147 2,86 2,06 ,077 340 P08 AC 35 500 1,207 .0131 3.49 0,95 ,070 320 P09 AC 40 530 .254 .0217 1,91 1.11 .075 230 P10 AC 35 450 .699 .0228 1.91 0.95 . 078 170 P l l AC 40 560 • .445 .0194 1.27 0.64 .071 275 P12 AC 44 510 ,445 ,0202 3,49 0.95 ,086 265 P23 AC 45 510 .508 .0227 4.13 0,79 .088 245 - 57 - 4.5.2 Temperature Time Cycles Figures 57-60 present the raw data from one experiment showing the f i v e thermal p r o f i l e s f o r each of the four thermocouple p o s i t i o n s . Figure 61 presents the same p r o f i l e s when s h i f t e d by the welding v e l o c i t y d i v i d e d by the distance between the thermocouples and then averaged.. A l l four thermo- couple p o s i t i o n s were then plotted on a s i n g l e graph. The thermal p r o f i l e s for the other experiments were p l o t t e d i n the same manner and a r e presented i n Figures 62-73. 4.5.3 Penetration and HAZ Values The penetration was measured using a scale and c o r r e c t i n g f o r any closure of the gap between the parent plates a f t e r welding. The HAZ was determined by m i c r o s t r u c t u r a l and VPN a n a l y s i s . The r e s u l t s are presented i n Table 10. 4.5.4 Hardness Traverses In Figure 74 three t y p i c a l hardness traverses are shown f o r the widest range of penetrations i n order to i n d i c a t e the e f f e c t of g r a i n s i z e on VPN values. 4.5.5 Inclusions i ) Table 12 presents the cumulative percentage s i z e d i s t r i b u t i o n for s i x t y p i c a l welds as determined using the Quantimet. i i ) Figure 75 presents the oxygen content traverse along the same sect i o n as the above i n c l u s i o n counting was taken. i i i ) Figures 76-77 are scanning e l e c t r o n micrographs of i n c l u s i o n s from the same specimens used f o r the s i z i n g survey. _ 58 TABLE 10 Penetration and Heat Affected Zone (HAZ) Values for the Experiments in Table 9 Experiment Penetration Heat Affected Number cm. zone cm.. P13 .381 1.32 P14 .699 2.05 P15 .191 1.46 P18 .254 1.555 P20 .320 0.92 P21S .318 1.65 P16 .064 0.54 P22S .254 1.40 P04 .325 2.23 P05 .445 2.02 P06 .191 1.86 P08 1.207 2.68 P09 .254 1.49 P10 .699 1.65 P l l .445 1.08 P12 • . '• .445 0.91 P23S .508 PI3 to P21S ahove are DCRP mode P16 to P22S above are DCSP mode P04 to P23S above are AC mode - 59 - TABLE 11 Grain Size i n the Heat Affected Zone of CGESW of A36 Plates Experiment Thermocouple ASTM ASTM Diameter Number P o s i t i o n * Grain Size Grain Size of Average Number Number Grain** Measured Determined mm. P13 1 1-2 1.0 .254 2 6-7 7.0 .032 P14 1 7 7.75 .025 2 10 10.75 .0094 PI 5 1 7 8.0 .0224 2 11 12.0 .00561 P18 1 .6.5 6.5 .038 P21S 1 6.0 6.5 .038 2 11 11.25 .0070 PI 6 1 7 7.75 .025 P22S 1 9 10.00 .0112 P04 1 5 5.5 .0534 2 9.5 10.75 .008 P05 1 6 6.5 .038 P08 2 8.5 9.0 .0159 P10 1 4 4.0 .0898 2 7.5 8.0 .0224 P l l 1 6 6.5 .038 P08 2 8.5 9.0 .0159 P10 1 4 4.0 .0898 2 7.5 8.0 .0224 P l l 1 6.5 6.75 .0350 2 10 10.50 .0094 P23S 1 6.5 6.75 .0350 * P o s i t i o n s 1 and 2 are 6.34mm ( V ) and 12.70mm Qi") from the o r i g i n a l roof gap surface. ** Average diameter of grain i s f o r c a l c u l a t e d grain s i z e numb PI3 to P21S above are DCRP mode PI6 to P22S above are DCSP mode P04 to P23S above are AC mode TABLE 12 Inclusion Distributions in CGESW of A 36 Plates Experiment Number DCRP Ind% 13 Cum% DCRP Ind% 14 Cum% DCRP Ind% 18 Cum% DCSP Ind% 16 Cum% AC Ind% P5 Cum% AC Ind% P10 Cum% Size u 0 - 0.52 34.1 100.0 34.2 100.0 33.8 100.0 33.9 100.0 34.4 100.0 32.3 100.0 0.52 - 1.04 28.1 65.9 28.8 65.7 27.1 66.2 29.6 66.1 30.4 65.6 27.6 67.7 1.04 - 1.82 22.0 37.8 23.9 37.0 22.5 39.1 22.8 36.5 2.3 35.2 23.1 40.1 1.82 - 2.60 10.1 15.8 9,1 13,1 12,2 16,6 8.7 13,7 8.5 13,9 12,4 17.0 2.60 - 3.90 14.2 5.7 2.6 4.0 3.1 4.4 2.6 5.0 2.7 5.4 3.4 4.6 > 3.90 1.5 1.5 1,4 1.4 1.3 .1.3 2.4 2.4 2.7 2.7 1.2 1.2 Total Number 1783 1549 1186 967 840 . 1201 - 61 - 4.5.6 Grain Size Analysis Grain s i z e a n a l y s i s was studied o p t i c a l l y and compared to a standard ASTM grain s i z e chart to obtain an ASTM grain s i z e number. Figures 78 and 79 show t y p i c a l g r ain s i z e s obtained at various p o s i t i o n s f o r both DCSP and DCRP. Table 11 presents the g r a i n s i z e f o r many experiments at the same p o s i t i o n s . The grain s i z e numbers were obtained by comparing a transparent ASTM overlay over the translucent viewing p l a t e of a Unitron microscope at 100 X magnification. 4.5.7 Charpy Traverses and Fracture Surfaces Standard Charpy V-notch Impact t e s t s (30°) (ASTM E23) were performed with the center of the notch i n the welds, adjacent to the F.B. i n HAZ 0.5 to 3 mm from the fusion boundary and 5 mm i n t o the HAZ. The t e s t s were performed at room temperature and the r e s u l t s tabulated i n Table 13 f o r a l l three modes of welding. SEM fractographs were taken of the f r a c t u r e surfaces, Figures 80-86. 4.6 ESW U t i l i z i n g Bar Electrodes 4.6.1 Introduction Eleven welds were made with bar electrodes and p l a t e s of the same mat e r i a l . Eight of the welds were made from ASTM A36 p l a t e and bars and four of the welds were made using cast T - l equivalent p l a t e and bars. Table 14 presents the data obtained or c a l c u l a t e d f o r the operating conditions. 4.6.2 Temperature Time Cycles Figures 87-94 are the thermal p r o f i l e s f o r the three modes of welding under the operating conditions presented i n Table 14. Figure 95 compares - 62 - TABLE 13 Charpy Values at Various Positions i n Welded A 36 Plate Experiment P o s i t i o n of Charpy Impact Number V-notch a t Room Temp. f t . l b s . DCRP P21 WELD CENTER 116 0.5 mm i n t o HAZ 27 1 mm into HAZ 75 2 mm int o HAZ Off S cale* 3 mm into HAZ Off Scale DCSP P22 WELD CENTER 70 0.5 mm int o HAZ 37 1 mm int o HAZ 41 spheroidized region (_ 2 mm).. Off Scale AC P23 WELD CENTER 95 Half way int o weld 55 Adjacent to F.B. i n parent metal 18 0.5 mm int o HAZ 29 1.5 mm into HAZ 30 Off scale denotes o f f scale above 220 f t . l b s . or the specimen did not completely break. TABLE 14 Experimental Conditions f o r ESW U t i l i z i n g Bar Electrodes Experiment Number Operating Mode Volts Amps Penetration cm Weld V e l o c i t y cm s e c ~ l Slag Depth cm Apparent Resistance ohms S p e c i f i c Power k c a l cm--'- PP3 DCRP A 36 35 690 0.417 .0767 2.69 .051 75 PP4 DCRP A 36 24 610 1.270 .0111 3.81 .039 321 PP5 DCRP A 36 33 725 0.826 .0582 2.79 .046 99 PP6 DCRP A 36 36 640 0.635 .0609 2,22 .056 92 PP7 DCRP A 36 34 710 0,640 .0512 3.68 .048 113 E l DCRP T - l 38 625 0.826 .0440 1.27 .061 128 E2 DCRP T - l 34 725 1.080 .0364 2.86 .047 160 E3 DCRP T - l 30 575 0.826 .0334 2.54 .052 124 PP1 DCSP A 36 40 775 1.27 .0685 0.953 .052 110 PP8 AC A 36 33 730 ,950 .0544 4,06 .045 106 E4 AC T - l 28 750 .064 .0440 1.27 .037 115 _ 64 _ one thermocouple p o s i t i o n f o r the three most d i f f e r e n t s p e c i f i c power inputs i n the D.C. mode of welding. 4.6.3 Penetration Figure 96 i n d i c a t e s the measured and hardness determined penetration f o r both the A36 and T - l equivalent welds. One penetration f o r each mode welding A36 and each penetration f o r a l l the T - l equivalent weld s p e c i f i c power inputs i s shown i n Figure 97. 4.6.4 Hardness Traverses Hardness traverses were only c a r r i e d out on the T - l equivalent p l a t e . These traverses were c a r r i e d out at 6.35 mm (.25 in.) i n t e r v a l s on h o r i z o n t a l sections along the center l i n e of the section and p a r a l l e l to the center- l i n e near the surface of the p l a t e (see Figure 98). The traverse along the p l a t e edge was considered necessary because of the high hardness that can be achieved i n T - l , even when only a i r quenched. The hardness traverse values were made using a Wilson Rockwell Hardness t e s t e r on the Rockwell C scale. 4.6.5 Charpy Traverses Standard Charpy V-notch Impact te s t s (30°) (ASTM E32) were performed on the base metal, weld metal and i n the HAZ adjacent to the fusion boundary at - 40°C. Table 15 presents the impacts values determined. 4.7 E l e c t r i c a l D i s t r i b u t i o n Figures 99-101 i n d i c a t e the quasi-steady s t a t e e l e c t r i c a l d i s t r i b u t i o n determined as described i n Chapter 3. In a d d i t i o n , each f i g u r e shows the quasi-steady state p h y s i c a l and thermal c h a r a c t e r i s t i c s of the weld. Figures 102-104 show the experimental arrangement. - 65 - TABLE 15 Charpy Values at Various P o s i t i o n s i n ESW Welded T - l Equivalent P l a t e Experiment Number Operating Mode Charpy Impact Value and P o s i t i o n f t . l b s . at - 40°C E l DCRP Base Metal 56 Weld Deposit 22.5 HAZ Adjacent to F.B. E2 DCRP Base Metal 52 Weld Deposit 21 HAZ Adjacent to F.B. 11.5 E3 DCRP Base Metal 52.5 Weld Deposit 22.5 HAZ Adjacent to F.B. 10 E4 AC Base Metal 56 Weld Deposit 23 HAZ Adjacent to F.B. HAZ 6.4 mm from F.B. 8 13.5 - 66 - 4.8 General Observations Figure 105 i l l u s t r a t e s the shape of the penetration i n v e r t i c a l and h o r i z o n t a l p o s i t i o n s during a t y p i c a l AC weld (PP8). Figure 106 i l l u s t r a t e s the shape of the penetration when welding i n the DCRP mode using bar electrodes. Figures 107 and 108 show the re s u l t a n t welds using Consumable Guide E l e c t r o s l a g welding f o r unstable and stable conditions r e s p e c t i v e l y . Figure 109 shows the surface obtained from a good weld. Figure 110 shows the surfaces of the welds used to determine the e l e c t r i c a l d i s t r i b u t i o n . Figure 111 ind i c a t e s the s i t u a t i o n when the penetration becomes so great as to cause the slag and molten metal to run out around the shoe. Figure 112 shows two welds i n v e r t i c a l and h o r i z o n t a l section plus the hardness test indentations. Figure 113 i s a c t u a l l y a casting and was obtained during Consumable Guide E l e c t r o s l a g welding when the sl a g depth was approximately 10 cm deep. The penetration was uniform on both sides but only about 4 cm below the s l a g / a i r i n t e r f a c e ; the molten f i l l e r metal and parent metal simply flowed downward and formed a casting. CHAPTER 5 DISCUSSION OF RESULTS 5.1 Power Supply C h a r a c t e r i s t i c s Both the Hobart welders used were constant voltage type welders i n that the slope of the voltage amperage curve was 16.6° (see F i g . 27). I t i s probable that the published data were obtained by use of a v a r i a b l e r e s i s t a n c e across the terminals with no lead wires, and readings were taken from.the panel mounted gauges f o r vol t a g e and amperage. The input voltage (published) was 230 v o l t s to the welder and, i n the experiments c a r r i e d out i n the present research, the input voltage was only 206 v o l t s . This would account f or some of the l o s s i n power as would the r e s i s t a n c e of the leads used i n the experiments. The nature of the resi s t a n c e of the s l a g pool i s not r e a l l y known at t h i s time but i t c e r t a i n l y i n volves the temperature dependence of the c o n d u c t i v i t y of the s l a g , and the p o l a r i t y and the mode of welding. For these reasons, each welding c o n f i g u r a t i o n and mode w i l l be d i s - cussed l a t e r ; the dead load curves shown i n Figures 27 and 28 are probably not a p p l i c a b l e to the resistance that the welding supply experiences during a c t u a l welding. There i s one w e l l established observation throughout t h i s research - AC always presents a greater o v e r a l l r e s i s t a n c e than e i t h e r DCRP or DCSP. This r e s u l t i s p o s s i b l y explained by the f a c t that the - 68 - p o l a r i z a t i o n f i e l d during AC i s very much smaller than during DC, because of the change i n p o l a r i t y every 60 seconds. During DC the formation of highly conducting surface layers (77) may reduce the apparent res i s t a n c e drop s i g n i f i c a n t l y as the p o l a r i z a t i o n f i e l d s are much l a r g e r . 5.2 I n d u s t r i a l Experiments at Canron 5.2.1 The Resistance of the Slag Pool The s e r i e s of welds made at Canron were made using Hobart 750 welders and thus the weld r e s u l t s should have been u s e f u l i n p r e d i c t i n g and obtaining the voltage and amperage c h a r a c t e r i s t i c s desired. Because the s l a g depth was constant, and the only desired property was f u l l pene- t r a t i o n of the welds, most of the other v a r i a b l e s were eliminated. Figure 114 presents the amperage-voltage c h a r a c t e r i s t i c s and Figure 115 presents the amperage-apparent r e s i s t a n c e c h a r a c t e r i s t i c s . The voltage amperage c h a r a c t e r i s t i c i s approximately what was expected; i t would be expected .to be a s t r a i g h t l i n e through o r i g i n i f the resistance was con- stant and not dependent on the slag temperature and weld c o n f i g u r a t i o n . The apparent s l a g r e s i s t a n c e drops dramatically with increasing current due to the increase i n temperature. Figures 116 and 117 show that the amperage i s markedly:increased when the wire feed rate i s increased, and that the res i s t a n c e drops markedly with increasing wire feed v e l o c i t y . This can be explained by the p r o b a b i l i t y that the f a s t e r wire r a t e allows f o r a greater immersion of the wire before i t melts o f f , and thus the i n t e r e l e c t r o d e gap i s smaller and the current density higher, causing a higher temperature below the electrode t i p . It i s also p o s s i b l e that the higher current density i s achieved when the wire s e t t i n g i s higher, and thus a higher temperature - 6 9 - i s achieved i n the slag pool and a lower apparent r e s i s t a n c e . 5.2.2 Energy Requirements Figure 118 presents the energy required to melt one gm of deposited metal as a function of the melt r a t e (or wire v e l o c i t y ) . This may be explained by the lower resistance and increased amperage when the melt rate i s high. This view i s s i m p l i s t i c because i t does not take i n t o account the degree of melting of the parent metal, or the increasing l o s s of heat to the water cooled copper shoes as the wire v e l o c i t y decreases. However, i t would appear that i f a higher melt r a t e i s achieved, the energy and thus the power requirement i s lower for the same weld. 5.3 Relating Published Experimental and Production Welding Parameters Table 16 presents the operating conditions for experimental and pro- duction welding from the references c i t e d . Figure 119 i n d i c a t e s that the data from the l i t e r a t u r e must be i n - terpreted c a r e f u l l y before any p r e d i c t i o n s can be made f o r the correct welding procedure. There should be some r e l a t i o n between voltage and amperage that i s nearly a s t r a i g h t l i n e , or curves as shown i n Figure 114. Figure 119 can only mean that the welding conditions of the present program must d i f f e r appreciably from the welding conditions of the published work. Figure 120 again shows the r e l a t i o n s h i p between welding r a t e and am- perage. In general an increase i n amperage causes the welding rate to increase. However, for AC, t h i s i s not the case and cannot be accounted for by a change i n voltage (see Table 16.). TABLE 16 Experimental and I n d u s t r i a l Welding Conditions from the L i t e r a t u r e Thermal Welding Resistance Energy Melt Reference Mode Vol t s Amps Energy Rate Rate k c a l sec cm s e c - l ohms k c a l gm -l gm s e c - l 4 DCRP 33 405 3.2 .115 .081 1.67 1.92 DCRP 36 415 3.6 .104 .086 1.75 2.0.6 DCRP 39 415 3.9 .092 .094 1.75 2.23 DCRP 39 425 4.0 .061 .092 2.00 2.00 DCRP 39 475 4.4 .056 .082 2.25 1.96 DCRP 40 500 4.8. .051 .080 2.50 1.92 DCRP 40 525 5,0 .057 .076 3.08 1.62 DCRP 40 550 5.3 .051 .073 3,08 1,72 DCRP 40 550 5.3 .047 .073 3.08 1.72 DCRP 40 550 5,3 .046 .073 3.08 1.72 . . . DCRP 42 575 5.8 ,040 .073 3.42 •1,70 . DCRP 46 575 6.3 .026 .080 3,42 1.84 DCRP 34 450 6.3 .089 .076 3.78 1.67 DCRP 40 600 5.8 .061 .067 DCRP 40 500 4.8 .042 .080 2.80 1.71'. . DCRP 40 700 6.7 .042 .057 DCRP 40 450 4.3 .061 .089 2.21 1.95 DCRP • 40 450 4.3 .056 .089 2.35 1.83 DCRP 40 600 5.8 .064 .067 1.88 3.09 DCRP 40 550 5.3 .047 .073 2.36 2.25 DCRP 42 450 4.5 .036 .093 2.29 1.97 DCRP 40 470 4.5 .085 DCRP 40 450 4.3 .056 .089 2.32 1.85 DCRP 40 470 4.5 TABLE 16 (cont.) Reference Mode Volts Thermal Welding Resistance Energy Melt Amps Energy Rate Rate kcal s e c - x cm s e c ~ l ohms kc a l gm~l gm s e c ~ l 23 AC 42 500 5.0 .029 .084 1.72 2.91 AC 39 650 6.1 .025 .060 1.27 4.80 AC 44 625 6.6 . 020 .070 2.10 3.14 AC . 47 500 5.6 .035 .094 1.20 4.67 AC 41 550 5.4 .032 .075 1.95 2.77 AC 44 625 6.6 .028 .070 2.01 3.28 AC 48 550 6.3 .029 .087 1.85 3,41 AC 46 475 5.2 .037 .097 1.59 3,27 16 — 34 480 3.9 .0157 .071 2.73 1.43 - 35,5 650 - 5.5 .0240 .055 2.51 2.19 — 53 390 5.0 .0153 .136 2.98 1.68 42 900 9.1 .0350 .047 1.78 5,62 52 DCRP 36 280 2.42 .0268 .129 1.65 1.47 AC 36 280 2.42 .0225 .129 1.96 1.23 DCRP 43 300 3.10 .0279 .143 2.03 1,53 DCRP 41 250 2.46 .0252 .164 1.78 1.38 DCRP 41 280 2.76 .0266 .146 1.90 1.46 DCRP 41 300 2.95 .0273 .137 1,97 1.49 : DCRP 41 350 3.44 .0327 .117 1.92 1,79 DCRP 41 380 3.74 .0343 .108 1.99 1.88 TABLE 16 (cont.) Reference Mode Thermal Welding Resistance Energy V o l t s Amps Energy Rate k c a l s e c - l cm s e c ~ l ohms k c a l gm~ 52 DCRP 41 400 3.94 .0386 .013 1.86 DCRP 41 440 4.33 .0429 .093 1.84 DCSP 41 450 4.43 .0466 .091 1.74 DCSP 41 250 2.46 .0193 .164 2.33 DCSP 41 300 2.95 .0214 .137 2.52 DCSP 41 345 3.39 .0252 .119 2.46 DCSP 41 380 3,74 .0268 .108 2.55 DCSP • 41 420 4.13 ,0322 ,098 2.34 DCSP 41 450 4.43 .0375 .091 2.16 17 DCRP 38 2600 23,7 ,0097 .015 .86 DCRP 39 2500 23.4 .0083 ,016 .99 DCRP 45 2000 21,6 '.0090 ,023 1,83 DCRP • 54 1500 19.4 .033 ,036 .84 DCRP 43 1600 16.5 ,029 ,027 .77 47 — 54 1400 18,1 ,056 ,039 ,95 — 48 825 9.5 ,075 ,058 .50 Melt Rate gm sec -1 2.11 2.35 2.55 1.06 1.17 1.38 1,47 1,76 2,05 TABLE 16 (cont.) Reference Mode Volts Amps Thermal Energy k c a l s e c - x Welding Rate cm sec -^- Resistance ohms Energy k c a l gm~l Melt Rate gm s e c ~ l 2 DC 27 575 3.7 .0635 .047 1.70 2.18 DC 29 575 4.0 .0508 .050 1.71 2.35 DC 33 575 4.6 .0339 .057 1.45 3.18 DC 37 585 5.2 .0339 .063 0.84 6.21 DC 49 585 6.9 .0169 .084 0.94 7.38 AC 51 585 7.2 .0169 .087 0.65 11.00. DC 42 585 5.9 .0203 .072 1.11 5.29 35 DCRP 42 700 7.1 .053 .060 1.31 5.42 DCRP 43 900 9.3 .081 .048 1.12 8.28 DCRP 43 750 .. 7.7 .063 .057 1.20 6.44 DCRP 43 925 9.5 .086 .046 2.01 4.72 DCRP 43 925 9.5 .086 .046 2.01 4.72 DCRP 33 925 7.3 .112 .036 1.22 5.97 DCRP 33 925 7.3 .112 .036 1.22 5,97 34 AC 54 700 9.1 .017 ,077 1,78 .5,11 DCRP 42 700 7.1 .050 .060 1,32 5,38 . - 74 - Figure 121 i n d i c a t e s that i n general the apparent resistance decreases with an increase i n welding rate. This e f f e c t can be a t t r i b u t e d to higher current causing a higher temperature and, consequently, a higher c o n d u c t i v i t y of the slag or, a l t e r n a t i v e l y , that the i n t e r e l e c t r o d e gap at higher welding rates i s smaller and the resistance path shorter. There i s a problem with both of these i n t e r p r e t a t i o n s because the higher temperature causes the electrode to melt more quickly and thus increases the i n t e r e l e c t r o d e gap, which should increase the resistance causing the temperature to r i s e . This s y n e r g i s t i c e f f e c t apparently gives r i s e to a quasi-steady state r e s i s t a n c e , temperature d i s t r i b u t i o n and i n t e r e l e c t r o d e gap. The energy required to deposit a gm of metal decreases r a p i d l y to approximately 1 k c a l per gm as the melt rate increases (See Figure 122). 5.4 I n i t i a l Large Electrode Experiments 5.4.1 Comparison of the Temperature Time Cycles Figure 38 i n d i c a t e s the d i f f e r e n c e between the three modes of welding using l a r g e electrodes and a high c o n d u c t i v i t y s l a g (75% CaF^ and 25% Al^O^). The r e s u l t a n t thermal p r o f i l e s show that the softness of the thermal c y c l e i n order of i n c r e a s i n g softness i s : DCRP, DCSP, AC. This suggests that the order of increasing penetration would be DCRP, DCSP and AC which, i n turn, suggests that i f maximum energy u t i l i z a t i o n at minimum penetration i s desired, then the weld v e l o c i t y should increase i n the following order: DCRP, DCSP, AC. However, decreased penetration can also be achieved by decreasing the electrode immersion ( i . e . the welding amperage or r e s i s t a n c e of the s l a g ) . I f the immersion i s decreased, the heat source volume w i l l be expanded to a larger area and the penetration decreased. The shape of the - 7 5 - thermal p r o f i l e s seen i n Figure 38 are of utmost importance as the ob- j e c t i v e i s to achieve the l e a s t penetration p o s s i b l e while s t i l l obtaining complete fu s i o n , and t h i s i s only p o s s i b l e with a s o f t thermal p r o f i l e . 5.4.2 Resistance of the Slag As stated i n Chapter 4, the re s i s t a n c e s are .017, .023 and .029 ohms for DCRP, DCSP and AC r e s p e c t i v e l y . Under normal e l e c t r o l y t i c r e a c t i o n s , where the i n t e r e l e c t r o d e gap remains constant, i t would be expected that the resist a n c e f o r DCSP would be l e s s than for the DCRP mode due to the higher production of FeO i n the DCSP mode. In t h i s case i t must be assumed that the apparent resistance of the DCRP i s due to a very pointed electrode t i p which a c t u a l l y r e s u l t s i n producing a small, i n t e r e l e c t r o d e gap and res u l t a n t small ..resistance path. This small electrode gap could only be due to the current d i s t r i b u t i o n i n the s l a g , and r a d i a t i o n plus convection of the s l a g . The AC re s i s t a n c e i s always higher regardless of the type of e l e c t r o s l a g welding or r e f i n i n g , and i s believed to be due to the small p o l a r i z a t i o n f i e l d s around the electrode.and parent metals' surfaces - which leads to an e s s e n t i a l l y ohmic re s i s t a n c e . With the higher r e s i s t a n c e the i n t e r e l e c t r o d e gap becomes la r g e r and, consequently, the heat source i s expanded leading to a s o f t e r thermal regime. Unfortunately, i n the AC experiment the s l a g was inadv e r t e n t l y deeper which also causes a higher r e s i s t a n c e ; however, the welding v e l o c i t y was 1.6 times f a s t e r and t h i s could r u l e out the deeper slag as a prime cause of the higher r e s i s t a n c e . 5.4.3 Energy Requirements The energy requirements per gram of deposited metal are 1.49, 1.24 and 0.81 k c a l per cm of weld height f o r DCRP, DCSP and AC r e s p e c t i v e l y . This would appear to mean that the most e f f i c i e n t energy u t i l i z a t i o n order would be AC, DCRP and DCSP, but t h i s does not account for the great d i f f e r e n c e s i n penetration for the three modes. When the c y l i n d e r of penetrated metal i s taken i n t o consideration, the energy f o r melting the metal per gram be- comes 0.88, 1.06 and 0.71 k c a l for DCRP, DCSP and AC r e s p e c t i v e l y . 5.4.4 Use of Inclusion D i s t r i b u t i o n Measurements Figures 39-41 show the i n c l u s i o n s to be uniformly small and very numerous. The i n c l u s i o n d i s t r i b u t i o n s could be construed to mean that the impact values are adversely a f f e c t e d by the i n c l u s i o n s , but probably no more than i n normal welding due to the small s i z e ( u s u a l l y l e s s 5ym). Except f o r the s i z e , shape and frequency of the i n c l u s i o n s , the usefulness of these data i s minimal without other data ( i . e . oxygen and s i l i c o n a n a l y s i s ) . 5.4.5 Measuring Penetration.and the Heat Affected Zone using Hardness Traverses The Vickers Pyramid Number.(VPN) traverse i s considered to be a u s e f u l technique f o r determining the f u s i o n boundary (FB) and heat a f f e c t e d zone (HAZ).. I f the s t e e l can be sectioned, polished and viewed m i c r o s t r u c t u a l l y , t h i s i s the best method. However, i n many cases t h i s i s not p o s s i b l e and the use of the hardness traverse i s very easy i n comparison. The hardness values drop s i g n i f i c a n t l y at the f u s i o n boundary and slowly r i s e again i n the grainrefinedyfepheroidized region. This i s considered to be due mainly to a Hall-Petch type r e l a t i o n s h i p . Figure 42 shows the m i c r o s t r u c t u r a l FB and HAZ and the VPN determined FB and HAZ; the agreement i s very good and, thus, hardness traverses have been used throughout t h i s research to confirm macroetched and microetched surfaces. (R.G, denotes the o r i g i n a l preweld root gap p o s i t i o n . ) - 77 - 5.4.6 Oxygen Analysis and the I n t e r p r e t a t i o n of Results Figure 43 represents the oxygen a n a l y s i s across the three modes of welding. I t i s obvious that the DCSP i s high' i n oxygen, approximately s i x times DCRP or AC. This i s due to production of oxygen a t the anode (metal pool surface). The oxygen analyses make the i n c l u s i o n d i s t r i b u t i o n values more meaningful i n that there are many more s i l i c a i n c l u s i o n s i n DCSP than DCRP or AC. Since they are more harmful to impact values, these l a t t e r are also lower for DCSP. 5.5 Consumable Guide ESW i n Cylinders 5.5.1 An Overview of the Experimental Operating Conditions This s e r i e s of experiments (see Table 5) was undertaken to determine the v a r i a b l e s of the process not mentioned i n the l i t e r a t u r e . There were no r e l i a b l e measurements of slag temperature d i s t r i b u t i o n s f o r e l e c t r o s l a g welding. There were no values of current d i s t r i b u t i o n when e l e c t r o s l a g welding with an e l e c t r o s l a g s l a g . There were no values f o r e l e c t r o d e immersion during welding. The present experiments were designed to determine these values and to attempt to explain.the observed thermal p r o f i l e s and pene- t r a t i o n mechanism. Figure 123 represents the electrode immersion f o r both DC modes of operating as a function of the slag r e s i s t a n c e . I t seems reasonable that the electrode immersion should increase with decreasing r e s i s t a n c e .because the j o u l e heating decreases. 5.5.2 I n t e r p r e t a t i o n of the Slag Temperature D i s t r i b u t i o n The slag temperature measurements are only i n d i c a t i v e of the r e a l temp- erature because the power had to be turned o f f and the boron n i t r i d e - 78 - sheathing d i s s o l v e d quickly. In addition,the m i l l i v o l t readings were only over a range of 20 mv to approximately 30 mv during a very b r i e f time i n t e r - v a l . Plus or minus 0.5 cm or 0.5 mv was observed and i s considered to be the error i n the Figures 44, 46 and 48. A l l three modes have the highest temperature region at the bottom of the electrode or j u s t below t h i s p o i n t . However, DCRP has the steepest temperature gradient and AC the lowest temperature gradient. The electrode immersion would appear to have the greatest e f f e c t on the temperature gradients but t h i s i s ambiguous as the slag r e s i s t a n c e w i l l drop with increasing temperature causing more current. to flow through the hot region and thus also regulate the immersion. 5.5.3 The E f f e c t of Slag Depth, P o l a r i t y and the Resultant Immersion and Electrode Gap on Penetration Figure 124 i n d i c a t e s that there i s a maximum i n the immersion versus energy per cm of weld produced. This curvature i s i n t e r p r e t e d as the i n t e r - play between the v e l o c i t y of the electrode f o r c i n g i t f u r t h e r into the slag and increasing energy a v a i l a b l e to melt the electrode. This e f f e c t appears to be independent of slag depth and therefore i t should be p o s s i b l e to s e l e c t a slag depth which w i l l u t i l i z e the immersion data to produce the desired i n t e r e l e c t r o d e gap. This allows f o r an estimate to be made to a t t a i n the desired penetration. T h e higher energy f o r the same immersion fo r DCSP i s considered to be due to the lower r e s i s t a n c e i n the DCSP mode. Figure 125 ind i c a t e s there i s a d e f i n i t e r e l a t i o n s h i p betxreen pene- t r a t i o n and energy input per cm of weld. The penetration shows a maximum for both DC modes, with DCSP r e q u i r i n g more energy than DCRP. With a con- stant slag depth t h i s can be explained by assuming the heat i s generated mostly.in the i n t e r e l e c t r o d e gap and, therefore, the maximum i n the penetration curve i s caused by the same e f f e c t as the maximum i n the - 79 - immersion curve. Once again the higher conductivity of the slag i n the DCSP mode shifts the penetration maximum to higher energy requirements. Figure 126 indicates the penetration for both slag depth and inter- electrode gap values. The points to the right of the ve r t i c a l line are due to the slag depth and interelectrode gap values being too high and presumably the current path i s to the sides and not to the molten steel pool. 5.5.4 Velocity Measurements and Their R e l i a b i l i t y Three methods were used and the results are compiled in Table 8. There is a great deal of discrepancy between the methods, particularly when one divides the total length of the weld by the time taken to produce the weld. The tungsten marker bed gives good results i f the tungsten can be cut through and is not too thick. The measurement of the time d i f f e r - ential between the maxima on the thermal profiles for thermocouples, the same distance from the fusion boundary, gave consistently good results. The problem with calculating the velocity using the total length i s that the sump does not necessarily f i l l up; slag i s often l e f t behind with some metal and this amount of metal varies with each experiment. Because of this problem the marker bed method, or more often the distance between peak maxima, has been used for the remainder of the interpretations. 5.5.5 Explanation of the Thermal Profiles Three thermal profiles are shown at positions .318 or .636 intervals from the original position of the root gap. These three thermal profiles correspond with three sets of slag temperature measurements in Figure 44, 46, 68. Experiment number 10 shows a f a i r l y shallow r i s e but large high temperature region adjacent to the molten slag. Experiment number 8 shows a slow r i s e to a small high temperature region and a subsequent drop in - 80 - temperature. Experiment number 14 shows a much slower r i s e i n temperature and small region that i s f a i r l y hot. The respective weld v e l o c i t i e s were .0330, .0339, .0557 cm sec ^ and, therefore, the compressed nature of the p r o f i l e s cannot be caused by large changes i n v e l o c i t y . They a l l have approximately the same slag depth (6.60-7.00 cm) and, therefore, the slag depth has not caused the change i n p r o f i l e s . This suggests that the only cause f o r the compression i s the observed s l a g temperatures which are prob- ably caused by the immersion v a r i a t i o n with p o l a r i t y ; DCRP 1.49 cm, DCSP 3.20 and AC 4.89. The time above 900°C f o r DCRP, DCSP and AC f o r the thermocouple .318 cm from the o r i g i n a l root gap was 100 sec, 0 sec, 25 sec r e s p e c t i v e l y . This i s consistent with the penetration observed when the electrode immersion i s taken into account. I t w i l l be shown l a t e r that t h i s i s due i n part to the current d i s t r i b u t i o n . 5.5.6 Penetration and Thermal P r o f i l e s When a cross sectioned view i s taken (see Figure 127) of the thermal p r o f i l e s , i t becomes s i g n i f i c a n t that the steeper the gradient the l a r g e r the penetration. When t h i s i s integrated with the time above 900°C, the order of penetration becomes r a t i o n a l . 5.5.7 Hardness Traverses Care has to be taken when analysing the hardness data as some welds were made with AISI 1020 wire into A 36 c y l i n d e r s , some with AISI 1042 i n t o A 36 c y l i n d e r s and some with AISI 1042 i n t o AISI 4340 and A 36 c y l i n d e r s . Welds made with AISI 1042 wire i n t o A 36 c y l i n d e r s ( i . e . experiment 10 and experiment 8) show a dramatic decrease i n hardness at the fusion boundary due to lower carbon content and increased g r a i n s i z e . Those with AISI 1020 into A 36 show a small decrease i n hardness due to increased g r a i n s i z e . - 81 - The welds with AISI 1042 into AISI 4340 show an increase in hardness due to the alloying constituents of the AISI 4340 producing a bainitic structure on cooling. A l l welds show a slight increase in hardness at the HAZ due to grain refinement and spheroidization of the carbide banding. Figure 53 indicates typical hardness traverses; the microstructurally determined fusion boundary (FB) and the heat affected zone (HAZ) are also shown. There i s good agreement between the microstructure and hardness changes. 5.5.8 Oxygen Analysis Results The oxygen analysis indicates that DCRP and AC have advantages over DCSP. The weld region for DCRP and AC has an oxygen content very close to the base metal analysis but the DCSP i s up to 100 ppm (or four times the base metal analysis). This would lead to higher susceptibility of DCSP welds to b r i t t l e fracture i n the weld zone (see Figure 54). 5.5.9 Stainless Steel Consumable Guide and Mixing i n the Liquid Steel Pool Figure 55 shows the distribution of the 304 S.S. nozzle as i t has so l i d i f i e d after melting the nozzle. When the consumable guide melts during welding, i t does not do so steadily but melts at regular intervals. This leads to the banding seen in most welds and, in the case where the nozzle i s stainless steel, a simple n i t a l etch reveals the pattern of the s o l i d i f i c a t i o n front. In this case i t i s seen that the 304 S.S. does not mix with the AISI 1042 to any appreciable extent. A trace across one of the bands using the electron microprobe revealed very low concentration of nickel i n the areas between the 304 S.S. bands. This i s true for both the transverse and longitudinal sections. - 82 - 5.6 Consumable Guide ESW with A 36 Plates 5.6.1 An Overview of the Experimental Operating Conditions This s e r i e s of experiments (see Table 9) was undertaken to provide information pertinent to production welding as the procedures used were as close to those used i n industry as i t i s p o s s i b l e to achieve i n the laboratory. Again the thermal p r o f i l e s were measured with time but emphasis was given to DCRP and AC because of t h e i r low oxygen.content and wide In- d u s t r i a l use. A ser i e s of three experiments, one f o r each mode, was under- taken to provide data f o r determination of the e l e c t r i c a l d i s t r i b u t i o n . F i g u r e 128 shows the electrode immersion as a f u n c t i o n of the apparent r e s i s t a n c e . The electrode immersion increases with decreasing apparent r e s i s t a n c e because there i s l e s s heating of the molten s l a g even though the amperage i s increasing. The apparent systematic d i f f e r e n c e between the AC and DCRP modes may be due to an error i n measuring the RMS voltage and current, or a phase s h i f t when welding i n the AC mode. Figure 129 i n d i - cates that the amperage decreases r a p i d l y with i n c r e a s i n g r e s i s t a n c e and therefore the heating of the molten slag should be higher and the immersion shallower. 5.6.2 Explanation of the Temperature Time Cycles The thermal p r o f i l e s are presented i n Figures 61 to 73. The same fi g u r e s also i n d i c a t e the time p o s i t i o n of the s l a g / a i r i n t e r f a c e , the electrode immersion and slag/molten s t e e l i n t e r f a c e . The p r o f i l e s are for thermocouples at 6.35 mm (1/4 inch) spacing with number one 6.35 mm from the o r i g i n a l parent metal root gap edge. The p r o f i l e f o r the number one thermal couple 6.35 mm from t h i s surface simply means that a thermal p r o f i l e - 83 - for a thermocouple at the surface would be s h i f t e d to the l e f t on f i g u r e s 61 to 73 and would i n d i c a t e a much higher temperature. A l l the thermal p r o f i l e s show that s i g n i f i c a n t heating of the parent metal takes place long before the s l a g / a i r i n t e r f a c e i s reached. When the sla g i s medium depth and the electrode immersion f a i r l y shallow as i n experiments P04, P05, P08, P22S, P23S (Figures.62-64, 72, 73} the thermal p r o f i l e s are much shallower. The heat generated i s d i s t r i b u t e d over a much longer parent metal/slag surface f o r a greater period of time. In a l l cases a large amount of heat i s l i b e r a t e d from the molten s t e e l pool as can be seen from the p r o f i l e s ; however t h i s would appear to be l e s s than the sl a g heat. The heat from the molten s t e e l pool i s both superheat and l a t e n t heat of fusion. The comparison between AC and DCRP p r o f i l e s w i l l be discussed l a t e r i n s e c t i o n 5.7. .5.6>3-Penetration and the Heat A f f e c t e d Zone as a Function of the Thermal P r o f i l e s The measured penetration and heat af f e c t e d zone f o r these experimental welds are presented i n Table 10. Experiments P13 through P21 are DCRP, P16 and P22S are DCSP and P04 through P23S are AC. DCRP, as a g e n e r a l i z a t i o n , has l e s s penetration than AC. This i s probably due to the deeper electrode immersion and shallower s l a g depth with the exception of P20S and P21S which - - 84 - were operated with a deeper s l a g depth for the e l e c t r i c a l d i s t r i b u t i o n measurements. In AC the penetration i s deeper even with the shallow thermal p r o f i l e s and deeper s l a g . I f the slag as i n P10 (Figure 65) i s shallower and the p r o f i l e steeper, the penetration i s even deeper. P08 (Figure 64) seems to be anomalous but the innermost thermocouples were' melted and thus the penetration i s believed to be cor r e c t . 5.6.4 Hardness Traverses Figure 74 presents the root gap (RG), fusion boundary (FB) and heat affected zone (HAZ) as determined m i c r o s t r u c t u r a l l y . I t also presents the Vickers Pyramid Number (VPN) hardness traverses. There i s a very good c o r r e l a t i o n between the two methods. The hardness change at the fu s i o n boundary i s due to the very large g r a i n s i z e and the increase at the end of the heat affected zone i s due to the small grain s i z e and spheroidization An attempt has been made to apply the a u t e n i t i z i n g parameter of Bastien et a l (68) to predict the austenite grain s i z e f o r regions that are i n the austenite phase region f o r a s i g n i f i c a n t length of time ( i . e . greater than 10 seconds at temperatures over 850°C). 5.6.5 Measured Grain Size, and Austenite Grain Size Table 11 compares the measured grain s i z e and the c a l c u l a t e d grain s i z e . The cal c u l a t e d g r a i n s i z e i s obtained by a semi-empirical r e l a t i o n - ship which defines an austentizing parameter P: R • t AH l 0 g ) -1 P where T E the equivalent maximum temperature over a period t, °K AH the a c t i v a t i o n energy for austenite g r a i n growth, 110,000 c a l mol -1 - 85 - t - time i n the a u s t e n i t i z i n g phase region t .= time u n i t , 1 sec. T 2 3 N D T E = T M - R A I where = the maximum temperature attained R = the u n i v e r s a l gas constant, 1.98 c a l mol °K A nomograph, Figure 14, i s used to determine the g r a i n s i z e of the austenite grains p r i o r to decomposition. As seen i n Table 11, the decomposition grain s i z e i s (within e r r o r of the measuring technique) the same s i z e as the austenite grain s i z e c a l c u l a t e d by t h i s method. In the two phase f i e l d , alpha plus gamma, there i s no formula that can be a p p l i e d to c a l c u l a t e the f e r r i t e and p e a r l i t e g r a i n s i z e s . However, the la r g e g r a i n s i z e observed and calculated suggests that the heat affected zone near the f u s i o n boundary would cause a decrease i n the impact strength of t h i s region. Figures 78 and 79 show the g r a i n s i z e f o r DCRP and DCSP r e s p e c t i v e l y at various positions.through the weld zone i n t o the unaffected parent metal. 5.6.6 Incl u s i o n Analysis Table 12 presents the Quantimet i n c l u s i o n a n a l y s i s of s i x weld sections and Figure 75 the oxygen an a l y s i s from the corresponding sections. The same t o t a l area was observed i n each case and thus, even though the oxygen an a l y s i s was as expected, the t o t a l area f r a c t i o n of i n c l u s i o n s was found to be higher f o r DCRP than f o r DCSP at AC. This reverse order of area f r a c t i o n with respect to DC modes can not be explained by the oxygen an a l y s i s as i n c l u s i o n d i s t r i b u t i o n s . The i n c l u s i o n s were examined at 3000 X mag n i f i c a t i o n on the scanning e l e c t r o n microscope and analysed f o r q u a l i t a t i v e chemistry. Figure 76 i s a - 8 6 - photo of the DCRP mode and Figure 77 of the DCSP mode. The DCRP i n c l u s i o n s were found to be very high i n MnS with some S i . The DCSP i n c l u s i o n s were found to be very high i n S i and low i n sulphur. This can be explained by the much deeper penetration i n the DCRP mode than i n the DCSP mode and thus much more sulphur has to be removed due to greater d i l u t i o n . 5.6.7 Charpy Impact Values The Charpy impact values using the ASTM V-notch technique are not meant to be accurate, only r e l a t i v e ; they are presented i n Table 12. The notch was oriented i n the plane of the pl a t e p a r a l l e l to the welding d i r e c - t i o n . Table 13 i s consistent with the f a c t that DCRP has greater penetration due to higher temperatures and thus the grain s i z e e f f e c t causes a very large drop i n the r e l a t i v e impact strength. This i s also observed i n DCSP and AC. The very low value of 18 f t . l b s . obtained f o r the AC weld i s only due to the notch being almost at the fu s i o n boundary but s t i l l i n the heat affe c t e d zone. Figures 80 through 86 are fractographs of the Charpy f r a c t u r e surfaces. 5.7 E l e c t r i c a l D i s t r i b u t i o n i n Consumable Guide ESW 5.7.1 Possible I n t e r p r e t a t i o n of Current D i s t r i b u t i o n Figures 99 through 101 give the e l e c t r i c a l d i s t r i b u t i o n f o r DCRP, DCSP and AC f o r very s i m i l a r operating conditions. Between the DC modes there i s very l i t t l e d i f f e r e n c e i n the current d i s t r i b u t i o n . This i s confirmation of the thermal d i s t r i b u t i o n shown i n Figures 50 and 51. Also t h i s i s v e r i f i e d by the e s s e n t i a l l y same shapes of the thermal p r o f i l e s f o r the DC modes. AC on the other hand appears to have an almost even current d i s - t r i b u t i o n and t h i s i s confirmed by the thermal d i s t r i b u t i o n i n the slag - 87 - (Figure 52) but only s l i g h t l y i n the thermal p r o f i l e s . I t i s very s i g n i f i c a n t that the e l e c t r i c a l d i s t r i b u t i o n i n the slag i s assumed i n the l i t e r a t u r e to be almost always from the electrode t i p to the molten l i q u i d s t e e l and very l i t t l e to the side w a l l s . The present research would i n d i c a t e that, w i t h i n a 25% margin of e r r o r i n the p o s i t i o n of the slag with respect to the shunt p o s i t i o n s , the current to the molten pool does not exceed 50% of the t o t a l current. I t i s very p o s s i b l e that the d i f f e r e n c e between the present research f i n d i n g s and published research f i n d - ings i s due to current being c a r r i e d by the consumable guide which has very shallow immersion and thus would tend to pass current to the side walls and not the molten s t e e l pool. The current d i s t r i b u t i o n to molten pool would be much higher i f the slag were shallower or the immersion much deeper. However, i n the present experiments the slag depth was only s l i g h t l y greater than i n i n d u s t r i a l p r a c t i c e . According to D i l a w a r i et a l (76) the slag flow i s dominated by transport caused by the applied current over convectiye transport. 5.7.2 C o r r e l a t i o n with Slag Temperature D i s t r i b u t i o n The temperature i n the slag i s achieved at a quasi-steady state simply by the a b i l i t y of the slag surroundings to remove heat from the s l a g and the source(s) of heat generation. The current d i s t r i b u t i o n of a l l modes of operation shows the maximum current i s j u s t below the electrode t i p (Figures 99 to 101) and the temperature d i s t r i b u t i o n s also show a maximum below the electrode t i p for conditions of deep slag and shallow immersion (Figures 50 and 51). This would appear to be reasonable because the current i s high, the slag w i l l heat up due to resistance heating, and once i t does i t s conductivity increases u n t i l steady state i s achieved between the current, temperature and conductivity. However, - 88 - the induced flow that i s obvious from the scooped out nature of the penetration shows that there i s considerable flow due to electromagnetic forces and t h i s flow i s toward the side of weld away from the electrode. Most of the publications that c i t e slag temperatures also show a maximum temperature i n the slag above the slag/metal i n t e r f a c e . The temperatures i n the l i t e r a t u r e are not for the common e l e c t r o s l a g weld f l u x e s and therefore the a c t u a l temperatures cannot be compared with those obtained i n the present research. Temperatures have been c i t e d as high as 2000°C (52) using a submerged arc welding f l u x f or e l e c t r o s l a g welding. D i l a w a r i and Szekely (76) have modelled the e l e c t r o s l a g process f o r an unusual p h y s i c a l configuration, but have shown that the electromagnetic forces i n small electrodes ( i . e . wire) are predominant and the mass flow i s as shown i n Figure 130. Unfortunately the current d i s t r i b u t i o n assumed i s not reported but the flow i s not unreasonable when re f e r r e d to the current d i s t r i b u t i o n reported i n the present research. 5.8 ESW U t i l i z i n g Bar Electrodes 5.8.1 Overview of the Experimental Operating Conditions The bar electrode welding conditions (see Table 14) were found to be much d i f f e r e n t from those of consumable guide welding. Voltage was con- s i s t e n t l y lower, amperage higher and the welding v e l o c i t y extremely high. The s p e c i f i c power was lower than f o r wire welding except for PP4 where the low voltage apparently caused the excessive penetration and low v e l o c i t y . A l l but three of the welds were made using DCRP. The s l a g depth v a r i e d s i g n i f i c a n t l y because the degree of melting of the s t a r t i n g compact i n the sump was not uniform. The a d d i t i o n of slag was kept constant throughout the - 89 _ experiment. Generally the electrode immersion was constant and never more than 5 mm average. The electrode always melted o f f with a small cone i n the center and four sharp corners which i n e f f e c t made the electrode surface f l a t (Figure 131). Table 14 also i n d i c a t e s that, using bar electrodes, the deposition r a t e i n g sec ^ i s much higher. While the e l e c t r i c a l d i s t r i b u t i o n was not measured i n bar electrode ESW, i t i s apparent from Figure 106 that the current must be near the sur- face, as would be expected from the very shallow immersion. The f i g u r e i s taken from the end of the experiment and therefore shows the e f f e c t of the runout blocks, but the transverse sections show that the weld i s sound and i t i s a t y p i c a l weld achieved when e l e c t r o s l a g welding with bars. The penetration does s t a r t at upper surface of the slag area and not lower down, as i n wire welding. D i l a w a r i et a l (76) have shown that, w i t h i n the l i m i t s of the r e s t r i c t i o n s of t h e i r model, the flow i s reversed with large e l e c - trodes - (compared with wire electrodes) and t h i s could explain the penetratior. s t a r t i n g at the top. In t h i s case the buoyancy forces are assumed to be pre-dominant over the electromagnetic forces. Buoyancy forces are simply n a t u r a l convection. D i l a w a r i et a l (76) have assumed f o r t h e i r model c a l - c u l a t i o n s a l i n e a r v e l o c i t y i n the s l a g pool of 0.40 m.sec ^ for wire and 0.02 m.sec f o r bar electrodes. While v e l o c i t i e s were never measured i n the present research, there i s no doubt from v i s u a l observation of the slag surface that the wire system has a much higher v e l o c i t y than the bar system. It i s f e l t that the penetration morphology i s due to the a c t u a l i n t e r e l e c t r o d e gap occurring between the electrode and the parent metal and not between the electrode and the molten metal pool. - 9 0 - 5.8.2 Explanation of the Temperature Time Cycles The thermal p r o f i l e s are steeper (see Figures 87-94) than f o r con- sumable guide welding, as might be expected f o r much higher welding velo- c i t i e s . However, i n many cases the maximum temperature reached by the thermocouple c l o s e s t to the weld i s lower. This g e n e r a l i z a t i o n applies to A 36 p l a t e s and bars welded i n the DCRP mode. Experiments E l through E4 were made u t i l i z i n g cast T - l equivalent p l a t e and with i n t e n t i o n a l l y shallow slag depths. These experiments r e - sulted i n much steeper thermal gradients. One explanation f o r steeper gradients i n bar electrode welding could be the reverse d i r e c t i o n of the slag flow and lower o v e r a l l flow r a t e of the slag i t s e l f as predicted by Dilawari et a l (76). The heat should then be more concentrated i n the smaller slag volume (due to the higher f i l l r a t i o ) and probably transferred i n the upper regions where the current should pass, t h i s being the smallest i n t e r e l e c t r o d e gap. Figure 106 shows that the penetration does s t a r t at the top of the s l a g . The transverse section of the same figure also shows the uneven penetration due to misalignment of the electrode; the p o s i t i o n of the thermocouples was at the side c l o s e s t to the electrode, thus measuring the maximum temperatures obtained i n the weld. The values of penetration presented i n Table 14 show that i n the DC mode of operation the values of penetration are a l l high with respect to consumable guide e l e c t r o s l a g welding. In the AC mode however, the pene- t r a t i o n i s comparable to consumable guide e l e c t r o s l a g welding. If the same argument i s applicable to penetration as to the thermal p r o f i l e s , then, because the current and temperature are very concentrated, one would expect - 91 - a thermocouple near the fusion boundary to give a very high reading (i.e. the liquidus temperature of the base metal). However, If experiment PP4 i s used as an example, then number 1 thermocouple which i s 0.635 mm from the parent metal face should have recorded a m i l l i v o l t reading equal to the liquidus as the penetration was 0.640 mm. This i s true for many of the bar electrode experiments and was investigated after the f i r s t readings on PP1. The thermocouples were affixed as usual to a plate of the same thickness, placed in a furnace and heated to approximately 1000°C and cooled, readings being taken with exactly the same recording device used i n the welding experiments. No logical explanation can be suggested for this discrepancy; the readings were reasonable in the furnace. Furthermore, the readings for E3 and E4 showed that the liquidus was reached; after that point the readings were erratic and unusable. The problem with experiments PPl through PP7 was the i n a b i l i t y to keep the electrode aligned exactly in the center. This resulted i n the pene- tration being more to one side, the side the electrode approached more closely. This resulted in abnormally deep penetration on one side and shallow penetration on the other. If the thermocouples were situated on the shallow side, there would be intermittent slag inclusions or even a slight slag skin and, even though the profile was steep, the absolute values would be low. Therefore the profiles are not as informative as the total penetration. For experiments PP8 through E4 the guiding apparatus was improved and the profiles were more consistent with much higher temperatures recorded. - 92 - • 5.8.3 Hardness and Charpy Values Figures 96-98 show the hardness values obtained f o r A 36 and T - l equivalent. The values were as expected with the T - l being very much harder except adjacent to the f u s i o n boundary i n the heat a f f e c t e d zone. Figure 114 demonstrates that the e f f e c t of the surface c o o l i n g by the water cooled shoes i s non-existent and the quenching by the parent metal i s at l e a s t as e f f e c t i v e . The Charpy values, Table 15, i n d i c a t e again that the weldment has s u f f i c i e n t notch toughness. However, the area adjacent to the f u s i o n boun- dary i n the heat a f f e c t e d zone was unacceptable f o r most a p p l i c a t i o n s i f a post heat treatment procedure i s not employed. This i s completely consis- tent with the decreased Charpy values obtained f o r A 36 at room temperature, when welded using consumable guide e l e c t r o s l a g . 5.9 Large Electrode Technology and I t s Implication 5.9.1 P o t e n t i a l Advantages and Disadvantages The u t i l i z a t i o n of large ( i . e . , plate) electrodes leads to s e v e r a l advantages: 1) simple mechanical feeding i n comparison with multiwire feeding. 2) r e l a t i v e l y cheap f i l l e r metal. 3) good chemical c o n t r o l i n comparison with multiwire (1). 4) a short term power in t e r u p t i o n does not r e s u l t i n a large defect, such as l a c k of penetration or slag i n c l u s i o n s . 5) the process i s very stable during operation. However, there are s i g n i f i c a n t disadvantages: 1) Long set-up times are required. - 9 3 - 2) A l a r g e r gap between the parent metal plates i s required and thus more energy i s required. 5.9.2 P o t e n t i a l Uses The f i e l d of greatest p o t e n t i a l use for large electrodes i s i n heavy- section welding. The production of large ingots (100 ton) can be achieved ' i n many ways. The most important choice i n f a b r i c a t i o n i s the d e c i s i o n to weld to the f i n i s h e d s i z e before or a f t e r forging and heat treatment. The choice i s dictated by m e t a l l u r g i c a l reasons, since the HAZ developed i n the weld i s extremely large. For s t a t i c a p p l i c a t i o n s , t h i s may w e l l be of no disadvantage and welding of two or more components may be acceptable. In the case where a heat treatment sequence i s man- datory for the f i n a l p roperties, e l e c t r o s l a g welding must be accomplished before such a heat treatment. In t h i s l a t t e r case, there i s no advantage i n having a narrow gap, or small HAZ, and therefore e l e c t r o s l a g welding with plates i s d e s i r a b l e because of process r e l i a b i l i t y . ESW u t i l i z i n g p l a t e electrodes i s p o t e n t i a l l y very usefuly when many heavy s e c t i o n items of the same geometry are to welded and subsequently heat treated. CHAPTER 6 DISCUSSION OF MODELS 6.1 Introduction There are e s s e n t i a l l y three equations to be evaluated f o r the ESW process i n the present work. The f i r s t i s the a n a l y t i c a l s o l u t i o n (64,65) for the temperature f i e l d i n the plates being joined. The second i s the empirical model of P e r t s o v s k i i (67) f o r c a l c u l a t i n g the immersion of the electrode during welding. The t h i r d i s the a p p l i c a t i o n of the austenising parameter of Bastien et a l (68) to cal c u l a t e g r a i n s i z e i n the heat aff e c t e d zone. 6.1.1 The A n a l y t i c a l Solution Equation The a n a l y t i c a l s o l u t i o n equation was solved f o r those experiments u t i l i z i n g a consumable guide and a wire electrode. A comparison of the temperatures, c a l c u l a t e d and measured, i s made. The a n a l y t i c a l s o l u t i o n equation was solved f o r temperature-time d i s t r i b u t i o n s published i n the l i t e r a t u r e (16,17,18,19). A comparison of the temperatures, calculated and measured, i s made. TABLE 17 Input Data Used for Solving the A n a l y t i c a l Solution f o r Consumable Guide E l e c t r o s l a g Welding (CGESW) with A36 Plates Parameter Math Symbol Computer Symbol Units CGS P05 P06 Experiment Number P08 P09 P10 P l l P12 Preheat Temp To PREH °C 25. 25. 25. 25. 25. 25. 25. Voltage V VOLT volt 42. 40.0 35. 40. •35. 40. 44. Amperage I AMP amp 490. 520. 500. 530. 450. 560. 510. Weld Velocity Vw VEL ' cm.sec-1 .0140 .0147 .0131 .0217 .0228 .0194 .0202 Conductivity k AK cal.cm-lsec-l°C-l 0.1 0.1 0.1 0.1 0.1 0.1 0.1 Specific heat Cp CEPE cal.gm-l°C-l .163 .163 .163 .163 .163 ,163 .163 Density P RHO gm.cm-3 7.5 7.5. 7.5 7.5 7.5 7.5 7.5 Time step dt A TIME sec 20. 20. 20. 20. 20. 20. 20. Thickness 6 THICK cm . 3,18 3.18 3,18 3,18 3.18 3.18 3,18 Efficiency % EFF % 58. 58. 58. 58. 58, 58. 58, Slag Depth - SLDEP cm 3.18 2.86 3.49 1.91 1.91 1.27 3.49 Immersion - cm 0.93 1.10 1.27 1.15 1.10 1.00 .93 Pool Depth - POOLDP cm 1.60 1.40 1.75 .95 .95 .70 1.70 Percentage Top Heat Source - SORSI % 25. 30. 30. 30. 30, 20, 20. Percentage Middle Source - S0RS2 % 50. 40. 40. 50. 40. 60. 60. : Percentage Bottom Source - S0RS3 % 25. 30. 30. 20. 30, 20. 20. Root Gap Y rg RTGAP cm 3.18 3.18 3.18 3.18 3.18 3.18 3.18 Surface Heat Transfer coeff. a SHTC .cal,cm-2sec-l°C-l .0005 .0005 ,0005 .0005 .0005 .0005 .0005 TABLE 17 (continued) Parameter Math. Computer Units Symbol Symbol CGS P13 Preheat Temp. To PREH °C 25. Voltage v • VOLT volt 32. Amperage I AMP amp 685. Weld Velocity Vw VEL cm.sec-1 .0159 Conductivity k AK cal.cm-lsec-l°C-l 0.1 Specific Heat Cp CEPE cal.gm-l°C-l .163 Density P RHO gm.cm-3 7.5 Time step dt A TIME sec 20. Thickness 6 THICK cm 3.18 Efficiency % EFF '.%. 58. Slag Depth SLDEP cm 1.59 Immersion - AIMMER cm 1.42 Pool Depth - POOLDP cm ,80 Percentage Top Heat Source - SORSI % 5. Percentage middle Source - S0RS2 % 80. Percentage Bottom Source - S0RS3 % 15. Root Gap Y rg RTGAP cm 3.18 Surface Heat Transfer coeff. a SHTC cal,cm-2sec-l°C ,0005 Experiment P14 25. 36, 760. .0232 0.1 .163 7.5 20. 3.18 58. 1.91 1,42 .95 25. 50, 25. 3.18 .0005 Tiber P18 25. 38. 750. .0191 0.1 .163 7.5 20. 3.18 58. 1.50 1.27 .75 10. 50. 40. 3.18 .0005 P20 25. 39. 570. .0174 0.1 .163 7.5 20. 3.18 58. 3.65 .81 1.80 15 65, 20. 3.18 .0005 - 97 - 6.1.2 The Electrode Immersion Equation : The electrode immersion equation i s solved and a comparison of the c a l c u l a t e d and measured immersions i s made. 6.1.3 Grain Size Determination The method of Bestien et a l (68) i s applied to a number of experiments and a comparison of the predicted and observed grain s i z e determinations i s made. The r e s u l t s of the c a l c u l a t i o n s (Appendix B) are presented i n Table 20. 6.2 A p p l i c a t i o n of the A n a l y t i c a l Model 6.2.1 Model Parameters The value of the various terms i n equation [ 2 . l l were assigned the values shown i n Table 17 or values calculated from data i n Table 17. The values from Table 9 are incorporated In Table 17 and the model input data. The thermophysical properties are obtained from the l i t e r a t u r e (16, 17, 18 and 19).These values vary s i g n i f i c a n t l y with temperature and thus should be varied during computation; but t h i s i s impossible when applying an a n a l y t i c a l s o l u t i o n . The e f f i c i e n c y of 58 percent i s one that i s used by other i n v e s t i g a t o r s (15,16) and i s close to the value i n t h i s research. The low e f f i c i e n c y i s due to the r e l a t i v e l y t h i n p l a t e s (3.18 cm) and the consequent high heat l o s s to the cooling water. - 98 - 6.2.2 Relative P o s i t i o n of the Heat Sources The r e l a t i v e p o s i t i o n of the three heat sources, , q^, and q , i s shown i n Figure 13. q ^ i s situated at the upper surface of the s l a g pool, q^ i s intermediate between the bottom of the electrode and the s l a g / l i q u i d s t e e l i n t e r f a c e and q^ i s s i t u a t e d at the p o s i t i o n equal to one h a l f the depth of the l i q u i d s t e e l pool. This r e l a t i v e p o s i t i o n i s based upon the d i s t r i b u t i o n assumed by Pugin (16). The percentage of the t o t a l heat a v a i l a b l e d i s t r i b u t e d to each p o s i t i o n i s determined by the r e l a t i v e strength of the sources and thus t h e i r r e l a t i v e p o s i t i o n s . • - 9 9 - 6.3 Comparison of the Model Generated and Measured Thermal P r o f i l e s f o r CGESW 6.3.1 Computer C a l c u l a t i o n s , D i g i t a l Output and Graphical Output A simple computer program was used to generate the temperature/time data required to p l o t the isothermal contours and temperature/time graphs displayed i n fig u r e s 132-142. The computer input data i s tabulated i n Table 17. The temperature/time data obtained experimentally f o r t h i s data are represented by a se r i e s of points f o r one thermocouple p o s i t i o n on the same f i g u r e s . The computer program and an example of the output are i n Appendix C. 6.3.2 A n a l y s i s of Graphical Results The predicted temperatures are always greater than the experimental temperatures. This d i f f e r e n c e i s very l a r g e , with the experimental values sometimes only one hal f of the predicted value. The predicted temperatures are f e l t to be too high and the experimental temperature to be r e l a t i v e l y accurate. The curves have the same shape, and heating and cooling rates are approximately the same. The major d i f f e r e n c e i s caused by the maxima i n the predicted temperatures. These values could be e a s i l y reduced by ad d i t i o n of a c o r r e c t i o n f a c t o r to the a n a l y t i c a l s o l u t i o n but t h i s would simply be an exercise i n curve f i t t i n g and would have no basis i n theory. I t i s evident here and i n the next section, that when the predicted temperatures are not above the melting point, the experimental and predicted values do not d i f f e r g r e a t l y . - 100 - The experimental r e s u l t s are believed to be more accurate because there are experimental r e s u l t s f or thermocouples which ceased functioning j u s t p r i o r to a run-out occurring and these thermocouples gave m i l l i v o l t output readings consistent with the melting point with no d i s c o n t i n u i t i e s i n the heating curve. In a d d i t i o n , the predicted temperatures would lead to the run-out c o n d i t i o n i f they were attained. The a c t u a l lower temperatures r e a l i z e d may be a consequence of the consumable guide to some degree. The only data a v a i l a b l e f o r pure wire welding i s shown i n the next section where the c o r r e l a t i o n i s better. The use of a consumable nozzle leads to the electrode a l t e r n a t i n g between a s i n g l e wire electrode and a combination wire-nozzle electrode. The a l t e r n a t i n g between a high and low current density process may cause the temperature p r o f i l e s to be l e s s sharp and thus give lower maxima. 6.4 Comparison of the Model Generated and Measured Thermal P r o f i l e s f o r Published Results 6.4.1 Computer C a l c u l a t i o n s , D i g i t a l and Graphical Output Another, e s s e n t i a l l y the same computer program was used to generate the predicted r e s u l t s f or the temperature/time r e l a t i o n s h i p s . The p l o t t e d and d i g i t a l output i s compared to the values i n the l i t e r a t u r e . The comparison data i s from Pugin (16), Sharapov (17,18) and Trepov (19). The data used i n the computer information input f i l e i s tabulated i n Table 19, with those values which have been estimated or assigned noted by an a s t e r i s k . (these values were not given i n the l i t e r a t u r e c i t e d ) . TABLE 18 Input Data Used for the Analytical Solution for Published Temperature-Time Relationships Literature Thermocouple Reference Placement from Plate edge cm. Experiment Designation Voltage Volts Amperage amps. Thickness cm. Welding Velocity cm.sec-1 Thermal Efficiency % Slag Depth cm.. Electrode Immersion cm.. Pool Depth cm. Pugin (16) 1.8,3.0 4.3,5.0 K a ) 34. 480. 5.0 .0157 58 3.35 2.25* 1.40 Pugin (16) 1.5,2.7 3.0,5.5 Kb) 34. 480. 5.0 .0163 55 2.50 1.50* 1.50 Pugin (16) 2.2,3.4 4.5,5.0 2 35.5 650. 5.0 .0240 55 2.30 1.30* 2.00 Pugin (16) 1.0,2.0 3.0,4.0 3 53. 390. 6.0 .0153 60 3.60 2.50* 1.40 Sharpov(17) 1.0,2.5 3.5,4.0 1 38. 2600. 65.0 .0097 88 4.00 1.40 2.00 Sharpov(17) 0.5,1.5 2.5,5.0 Kb) 39. 2500. 65.0 .0083 83 4.00 1.40 2.00 Sharapov(17)1.0,2.0 3.0,4.0 3 39. 2000. 30.0 .0090 88 4.50 3.30 3.00 Sharapov(17)1.0,2.8 4.0,5.0 4 54. 1500. 16.0 .0330 87 7.50 1.37 7.50 Sharapas(17)l.0,2.0 3.0,4.0 5 43, 1600, 17,0 .0290 83 4,50 2,50 3.40 Sharapov(18)l.0,2.0 3.0,4.0 - 40. 3500. 65.0 .0180 85 3.00 2,50 2.50 Trepov (19) 1.0,1.5 2.0,5.0 1 40. 360, 1.8 .0517 52 2,50 1.50 1.50 Trepov (19) 1.0,1.5 2.0,5.0 2 38. 400. 7.0 .0150 61 2.00 . 1.75 1.00 TABLE 18 (continued) H-1 3 •B. 1) A l l other parameters are the same as those tabulated i n Table 17. 2) The heat source a l l o c a t i o n s are q]_ - 25% (top) q 2 - 50% (middle) q 3 - 25% (bottom) i n a l l cases, these are assumed by the authors. - 103 - 6.4.2 Analysis of the Predicted Temperature-Time Relationships The predicted values and experimental values compare favourably. Once again, the experimental values are lower than the predicted value. The predicted temperature/time curves, see f i g u r e s 143 to 153, have maxima i n the region 1050-1150°C and as with the e a r l i e r lower temp- erature comparisons the c o r r e l a t i o n i s better at these lower temperatures. The reason the predicted temperatures are much lower f o r these experiments -2 are the s p e c i f i c power input i s 30 to 40 kcal.cm f o r the l i t e r a t u r e -2 c i t e d and 50 to 110 kcal.cm for the present research. This d i f f e r e n c e i n power required i s u s u a l l y a t t r i b u t e d to the pl a t e thickness being greater f o r the experiments i n the l i t e r a t u r e and thus the r e l a t i v e l o s s of heat to the co o l i n g water being much smaller. However, the thermal e f f i c i e n c i e s used are not 100% higher as would be required (they are 35% higher). The a n a l y t i c a l s o l u t i o n does p r e d i c t the temperatures r e a l i z e d experimentally f o r these thicker p l a t e experiments. 6.5 A p p l i c a t i o n of the A n a l y t i c a l Solution to a Large E l e c t r o s l a g Weld An attempt was made to apply the a n a l y t i c a l s o l u t i o n p r e d i c t i o n s to a large e l e c t r o s l a g weld. The weld i n question was made with the following conditions: Weld configuration - parent metal 200 cm. cubic blocks - root gap = 10 cm. - slag depth ^ 7.5 cm. - 104 - - pool depth ^ 10 cm. - p l a t e electrodes. Welding conditions - power 800 kilowatts. - voltage 40 v o l t s . - amperage 20,000 amps. - welding v e l o c i t y .00833 cm/sec. - welding time ^ 7 hours. The a n a l y t i c a l s o l u t i o n program i s applied to t h i s weld the same way i t i s applied to the CGESW. This means the electrode immersion must be assigned a value. The value assigned was 2 cm. The r e s u l t a n t contour map from the d i g i t a l output was used to estimate the depth of penetration, t h i s contour map i s pre- sented i n f i g u r e 154. The penetration i s estimated to be equal to the p o s i t i o n i n the parent metal that reaches 1450°C. The depth of the heat a f f e c t e d zone i s assumed to be equal to the p o s i t i o n i n the parent metal that reaches 850°C. The predicted penetration i s 2.6 cm and the HAZ i s 13 cm. These r e s u l t s seem reasonable for a weld of t h i s s i z e as a deep penetration i s necessary and should be large to ensure a sound weld. The r e s u l t s are compatible ( i . e . , a ^ value of 1.5) with the r e s u l t s desired by the welding procedure design. The model p r e d i c t s r e s u l t s that are consistent with large e l e c t r o - slag welding technology and are reasonable. 105 TABLE 19 Electrode Immersion Values CGESW Using PF 201 Slag arid Plates 1) A l t e r n a t i n g Current (AC) Experiment Immersion Measured Calculated P04 1.10 1.10 P05 1.03 0.93 P06 2.06 1.10 P08 0.95 1.27 P09 1.11 1.15 P10 0.95 1.10 P l l 0.64 1.00 P12 0.95 0.93 P13 0.79 0.89 Di r e c t Current Reverse P o l a r i t y Experiment Immersion Measured Calculated P13 1.42 1.42 P14 1.11 1.42 P15 1.59 1.27 P18 1.27 1.27 P20 0.79 0.81 P21 0.79 3) D i r e c t Current Staright P o l a r i t y Experiment Immersion Measured P16 1.62 P22 0.95 0.79 Calculated 1.62 0.995 - 106 - 6.6 Ap p l i c a t i o n of the Electrode Immersion Equation Equation (2.17) was proposed by P e r t s o v s k i i (67) to be ap p l i c a b l e to ESW and was developed from a two dimensional analogue. The equation has been applied i n the present research to both CGESW of plates and in t o c y l i n d e r s . The r e s u l t s comparing the cal c u l a t e d and measured immersions are given i n Table 19. The c a l c u l a t i o n requires use of s p e c i f i c resistances from the l i t - erature (67) for slags of s i m i l a r composition and i n the same temperature range. However, i t was found that an acceptable technique was to c a l c u l a t e the s p e c i f i c resistance from one immersion value known to be accurate. This s p e c i f i c resistance was then used to c a l c u l a t e the other electrode immersion values. This technique was used to obtain the values i n Table 18 The measured and cal c u l a t e d values c o r r e l a t e reasonably considering the v a r i a t i o n of the s p e c i f i c r e s i s t a n c e with temperature. The c a l c u l a t i o n technique f o r electrode immersion i s given i n Appendix B. TABLE 20 Grain Size Determined for CGESW A 36 Plate Experiments Experiment Number Thermocouple Position Maximum Temperature Attained, T„ °K Time above 850°C, t sec. Equivalent Maximum Temperature, T °K E Austenizing Parameter (to = 1 sec) °K °C ASTM Microgram Size Number P13 1 1628 325 1580 1891 1618 1.0 2 1323 205 1291 1473 1200 P14 1 1308 120 1277 1260 1161 7.75 2 1173 75 1148 1434 987 10.50 P15 1 1243 165 1215 1368 1095 8.0 2 1123 20 1100 1169 896 12.0 P18 1 1353 220 1320 1514 1241 6.5 P21S 1 1353 225 1320 1515 1942 6.5 2 1138 80 1114 1221 948 11.35 P16 1 1283 160 1268 1434 1161 7.75 P22S 1 1183 165 1158 1296 1023 10.0 P04 1 1393 230 1358 1566 1293 5.5 2 i i 7 3 ; .: 70 1148 1258 985 10.75 P05 1 1333 450 1301 1518 1245 6.5 P08 2 1223 210 1196 1351 1078 9.0 P10 1 1473 : 270 1249 1518 1244 6.75 2 1278 160 1320 1266 993 10.50 P l l 1 1353 240 1320 1518 1244 6.75 2 1173 . 90 1148 1266 993 10.50 P23S 1 1338 235 1305 1497 1224 6.75 - 108 - 6.7 A p p l i c a t i o n of the Bastien et a l (68) Model to CGESW of A36 Plate The grain s i z e for the plate welds produced by consumable e l e c t r o - s l a g welding was o p t i c a l l y determined at 100X magnification by use of a comparative transparent ASTM grain s i z e overlay. The g r a i n s i z e number was also calculated and determined using the method of Bastien et a l ( 6 8 ) . The calculated grain s i z e method i s presented i n Appendix C and the r e s u l t s presented i n Table 20. The comparable number f o r the experimental g r a i n s i z e and the c a l - culated grain s i z e are presented i n Table 11. The c o r r e l a t i o n between the grain s i z e numbers i s f e l t to be good and a c t u a l l y the same when the experimental er r o r p o s s i b l e i s taken into account. However, there i s a deviation that i s constant and may be the r e s u l t of the e l e c t r o s l a g process. The actual (experimental) grain s i z e i s smaller than the pre- dic t e d grain s i z e when the grain s i z e i s large ( i . e . , f o r small grain s i z e numbers) and the actual grain s i z e i s l a r g e r than predicted when the grain s i z e i s smaller. This could be v i s u a l i z e d as a r e s u l t of the applicable nomogram having a shallower slope than i n f i g u r e 14. It should be noted that the grain sizes measured and predicted are approximately the same and the v a r i a t i o n across the distance of 1.27 cm i s very large. This gives r i s e to the impact strength problems. I t also leads to the unfortunate r e s u l t , that continuous c o o l i n g curves (C.C.T.) are not a p p l i c a b l e throughout t h i s region. Two or more C.C.T. curves would be required to account for the extreme change i n g r a i n s i z e . In order to evaluate the changes i n structures on the b a s i s of the ex- perimental or predicted thermal cycles a c o r r e c t i o n would have to be made for the grain s i z e v a r i a t i o n within the heat a f f e c t e d zone. - 109 - CHAPTER 7 CONCLUSIONS The technique of i n v e s t i g a t i n g e l e c t r o s l a g welding by use of a hollow c y l i n d e r has been found to be u s e f u l i n i n v e s t i g a t i n g the process v a r i a b l e s . This technique removes the problems associated with cooling water and a sla g skin. In ef f e c t , t h e c y l i n d e r could be used as a temperature/time simulator f o r the high temperature h i s t o r y of e l e c t r o s l a g parent metal. Hardness traverse data are u s e f u l i n determining the depth of pen- e t r a t i o n and extent of the HAZ. This technique could be used with a section through the cut-off run-out block assembly used i n many welding shops. It has been shown, from the oxide i n c l u s i o n and consequent weld metal impact strength, that the DC reverse p o l a r i t y and AC modes of welding are preferable f o r e l e c t r o s l a g welding. The slag temperature d i s t r i b u t i o n s i n d i c a t e the maximum temperature i s approximately 1700°C and DCRP has the steepest gradient and AC the shallowest. I t can be concluded, from the maxima i n the penetration and immersion versus s p e c i f i c power input data, that the DCSP mode requires more - 110 - energy due to the lower r e s i s t i v i t y of the s l a g when welding i n th i s mode. It has been concluded there i s l i t t l e mixing i n the l i q u i d pool because the n i c k e l from the s t a i n l e s s s t e e l nozzle d i d not d i s t r i b u t e to any degree through the weldment The s l a g depth and immersion (i n t e r e l e c t r o d e gap) would appear to have the greatest e f f e c t on the rate of heating and co o l i n g as p l o t t e d on the temperature/time p r o f i l e s . Thus, these two parameters have the greatest e f f e c t on the degree of penetration and the s i z e of the HAZ. The DCRP mode has the greatest degree of penetration for equivalent welding conditions and thus, the la r g e s t grain s i z e and lowest impact strength. AC i s the best welding mode to achieve penetration with a modest thermal p r o f i l e while producing the lowest degree of grain growth. When the consumable guide technique i s u t i l i z e d , the welding current i s found to be d i s t r i b u t e d over the depth of the s l a g pool and approximately 50% the current passes to the molten l i q u i d pool. This i s s i g n i f i c a n t i n that most of the l i t e r a t u r e assumes more than 80% of the current passes through the l i q u i d metal/slag i n t e r - face. The current d i s t r i b u t i o n i s confirmed by the temperature d i s t r i b u t i o n . The bar electrode method of e l e c t r o s l a g welding when compared to consumable guide electrode welding has the fol l o w i n g features: - low s p e c i f i c power requirements. - shallow electrode immersion. - I l l - - high deposition rates. - shallow slag depth requirements. - greater degree of penetration. - greater l o s s of impact strength. The a n a l y t i c a l model i s not very u s e f u l when applied to the CGESW experiments i n the present research. I t i s , however, u s e f u l for welding schedules incorporating t h i c k e r plates up to 200 cm. The P e r t s o v s k i i electrode immersion equation i s app l i c a b l e to CGESW. The Bastien et a l austentising parameter i s u s e f u l i n determining the g r a i n s i z e of the austenite i n the HAZ adjacent to the fus i o n boundary and thus the e f f e c t on the impact strength i n the HAZ. - 1 1 2 - CHAPTER 8 SUGGESTIONS FOR ADDITIONAL RESEARCH 1) The a n a l y t i c a l model used i n t h i s research has not been found to be of p r a c t i c a l value fox the p r e d i c t i o n of the penetration or extent of the HAZ for consumable guide e l e c t r o s l a g welding of t h i n p l a t e s . The use of the hollow c y l i n d e r technique allows use of c y l i n d r i c a l symmetry and thus heat flows only i n the l o n g i t u d i n a l and r a d i a l d i r e c t i o n s . A f i n i t e d i f f e r e n c e technique could be developed and would probably be u s e f u l i n p r e d i c t i n g many of the steady state welding conditions. 2) The electrode immersion, i n t e r e l e c t r o d e gap and s l a g depth i n t e r r e l a t i o n - ships could be investigated u t i l i z i n g a s i n g l e s l a g , a non-consumable electrode and a c y l i n d r i c a l parent metal. In the steady s t a t e con- d i t i o n the temperature and current d i s t r i b u t i o n s could also be invest- igated. 3) It would also be enlightening to study the e f f e c t of adding a large percentage of the f i l l e r metal as a powder. This would decrease the temperature of the slag pool and allow f o r chemistry changes i n the weldment. - 113 - Investigation of lower melting point slags and slags that r e a c t vigorously with the parent metal surface chemically or by wetting would be u s e f u l . These slags would allow f o r penetration without excessive austenite grain growth. - 114 - REFERENCES Fatton, B.E., " E l e c t r o s l a g Welding", American Welding Society, 1962. Coplestone, F.W., paper to Commission XII of the I n t e r n a t i o n a l I n s t i t u t e of Welding, 1965. Harvyoshi Suzuki et a l , U.S. Patent 3,352,993, Nov. 14, 1967. Matsuoka,- A r a k i , . Suzuki and Murai, B r i t i s h Welding Journal, June, 1967, pp. 287-298. Linde, E l e c t r i c Welding Instructions, p u b l i c a t i o n F-51-220, Feb., 1969. Paton, B.E., Welding Journal, D e c , 1962, pp. 1115-1123. Saffonnikov, A.N., Avtomaticeskaya Svarka, No. 12, 1967, p. 69. Malysheuskaya, E.G., and Andrew, V.P., Avtomaticeskaya Svarka, No. 3, 1968, pp. 5-7. Ruklin, P.N. and Fokin, N.I., Avtomaticeskaya Svarka, No. 4, 1968, pp. 49-51. Medovar et a l , Federal Republic of Germany Patent 1,917,861, A p r i l 8, 1968. Saffonikov et a l , Avtomaticeskaya Svarka, No. 7, 1963, pp. 29-33. Lefevre, M., Arcos, v o l . 45, 1968, pp. 4205-4216. Miska, H.K., Materials Engineering, v o l . 73, 5, pp. 25-28. Hrivnak, I., Metal Construction, 1, Feb., 1969, pp. 74-77, (discussion pp. 128-129). Mel'bard, S.N., Svarocnoe Proizvodstvo, No. 9, 1960, pp. 5-7. Pugin, A.I., and P e r t s o v s k i i , G.C., Avtomaticeskaya Svarka, No. 6, 1963, pp. 14-23. - 115 - 17. Sharapov, Yu. V., Avtomaticeskaya Svarka, No. 6, 1965, pp. 32-37. 18. Sharapov, Yu. V., Svarocnoe Proizvodstvo, No. 6, 1968, pp. 13-15. 19. Trepov, P.V., Avtomaticeskaya Svarka, No. 9, 1969, pp. 15-17. 20. Shvartser, A. Ya., and Z o l o t a r c e u s k i i , D.B., Avtomaticeskaya Svarka, No. 4, 1968, pp. 16-19. 21. Eregin, L.P. , Svarocnoe Proizvodstvo, No. 7, 1970, pp. 25-27. 22. Makara, A.M. et a l , Avtomaticeskaya Svarka, No. 6, 1963, pp. 24-29. 23. Rote, R.S., Welding Journal, May, 1964, pp. 421-426. 24. Zeke, J . , B r i t i s h Welding Journal, May, 1969, pp. 258-268. 25. Woodley, C.C., and Burdekin, F.M., B r i t i s h Welding Journal, June, 1966, pp. 387-397. 26. Naumchenkov, N.E., Svarocnoe Proizvodstvo, No. 5, 1967, pp. 12-14. 27. Woodley, C.C., Burdekin, F.M. and Wells, A.A., B r i t i s h Welding Journal, March, 1966, pp. 165-173. 28. Malinovska, E. and Hrivnak, I., B r i t i s h Welding Journal, Oct., 1967, pp. 527-532. 29. B e l e n ' k i i , A.M., Toshchev, A.M., and Parkheta, V.K., Avtomaticheskaya Svarka, No. 12, 1967, pp. 56-59. 30. Pokataev, S.E., Gudov, I . I . , Chernykh, V.V., M a l a i , A.E., and Eregin, L.P., Svarocnoe Proizvodstvo, No. 5, 1968, pp. 5-7. 31. Bentley, K.P., B r i t i s h Welding Journal, August, 1968, pp. 408-410. 32. Zeke, J . , and Malinovska, E., B r i t i s h Welding Journal, December, 1968, pp. 621-626. 33. Braun, M.P., et a l . , Avtomaticeskaya Svarka, No. 10, 1969, pp. 7-10. 34. Makara, A.M., Yegorova, S.V., and Novikov, I.V., Avtomaticeskaya Svarka, No. 12, 1969, pp. 1-5. - 116 - Makara, A.M., Egorova, S.V., and Novikov,T.I., Avtomaticeskaya Svarka, No. 3,. 1971, pp. 52-55. Novikov, V.N., Tutov, I.E., and Kondraskev, A.I., Metallovedenie i Obrbotka Metallov, No. 8, August 1958, pp. 38-43. Eichhorn, F., and Shabeeb, A.R., Schweissen und Schneiden, v o l . 25, No. 5, 1973, pp. 174-177. Ikeno, T. et a l . and Takahashi, Y., et a l , Tetsu to Hagane, v o l . 59, No. 4, 1973, Lectures 146-149, pp. S148-S151. Voloshkevich, G.Z., Avtomaticheskaya Svarka, V o l . 6, No. 6, 1953, pp. 3-10. G o t a l ' s k i i , Yu. N., Avtomaticheskaya Svarka, V o l . 7, No. 5, 1954, pp. 38-43. G o t a l ' s k i i , Yu. N., Avtomaticheskaya Svarka, V o l . 8, No.- 5, 1955, pp. 47-49. G o t a l ' s k i i , Yu. N., Svarocnoe Proizvodstvo, No. 3, 1957, pp. 1-3. Ostrovskaya, S.A., Avtomaticheskaya Svarka, V o l . 10, No. 4, 1957, pp. 33-47. Zaitsev, Yu. M., and Tyagun-Belous, G.S., Avtomaticheskaya Svarka, Vo l . 11, No. 11, 1958, pp. 57-60. Muller, R., Schweissen and Schneiden, V o l . 10, No. 9, pp. 359-367. Poznyak, L.A., Zaitsev, Yu. M., and Tickhonovskii, A.L., Avtomatiches- kaya Svarka, V o l . 11, No. 10, 1958, pp. 67-74. Kozulin, M.G., Syatishev, A.P., and Emel'Yanov, Y.V., Avtomaticheskaya Svarka, No. 5, 1969, pp. 47-49. Dubovetskii, V., Ya., et a l , and Bronshtein, L.M., and Sigarev, V.S., Avtomaticheskaya Svarka, No. 8, 1969, pp. 44-47. - 117 - 49. Ivochkin, I . I . , and Sosedov, A . F . , Svarocnoe Proizvodstvo, No. 11, 1969, pp. 18-19. 50. Yushchenko, K.A. , et a l , Avtomaticheskaya Svarka, No. 5, 1970, pp. 72-73. 51. Makara, A.M., Egorova, S.V., Novikov, I.V., and Bronshtein, L.M.,. Avtomaticheskaya Svarka, No. 10, 1970, pp. 43-46. 52. Ando, K., and Wada, H., The Welding Journal, (Japanese), V o l . 39, No. 7, 1970, pp. 62-76. 53. Ando, K., Nakata, S., and Wada, H., The Welding Journal (Japanese), V o l . 40, No. 10, 1971, pp. 62-71. 54. Ando, K., Nakata, S., and Wada, H., The Welding Journal (Japanese), V o l . 40, No. 11, 1971, pp. 28-34. 55. Feldman, M. et a l , Soudage et Techniques connexes P a r i s , V o l . 25, No. 9/10, Sept.-Oct., 1971, pp. 359-377. 56. Zeke, J . , Slabon, I. and Mosny, J . , Welding Research Abroad, V o l . XXIII, No. 2, Feb., 1972, pp. 18-31. 57. Lanyie, L., and Zeke, J . , Welding Research Abroad, V o l . XXIII, No. 2, Feb., 1972, pp. 32-45. 58. Shaheeb, A.R., Schweissen und Scheiden, V o l . 25, No. 6, 1973, pp. 223-224. 59. Sharapov, Yu. V., Svarocnoe Proizvdstvo, No. 7, 1967, pp. 21-24. 60. Shcherbina, M., Ya. et a l , Avtomaticheskaya Svarka, No. 9, 1970, pp. 45-48. 61. Rosenthal, D., Transactions ASMF,, V o l . 68, 1946, pp. 849-866. 62. Rosenthal, D. and Cameron, R.H., Transactions ASME, V o l . 69, 1947, pp. 961-968. - 118 - Rykalin, N.N., " C a l c u l a t i o n of Heat Flow i n Welding", t r a n s l a t e d by Z. Paley and CM. Adams, U.S. contract number UC-19-060-001-3S17, 1951. Rosenthal, D., 2-eme Congres National des Sciences, 1935, pp. 1277-1292. Goldack, J.A., Burbridge, G., and Bibby, M.J., Canadian M e t a l l u r g i c a l Quarterly, V o l . 9, No. 3, "1971, pp. 459-466. Goldack, J.A., Burbridge, G. and Bibby, M.J., Canadian M e t a l l u r g i c a l Quarterly, V o l . 9, No. 3, 1971, pp. 467-473. P e r t s o v s k i i , G.A., Svarka, Vol. 1, 1958, pp. 187-193. Bastien, P.G., D o l l e t , J . and Maynier, Ph., Metal Construction, V o l . 2, No. 1, Jan., 1970, pp. 9-14. Rykalin, N.N., Soudage et Techniques Connexes, V o l . 15, 1961, p. 5. Haveri, J . , Moffat, W.G. and Adams, CM., Welding Journal, Jan., 1962, p. 125. Maynier, Ph., Martin, P.F. , and Bastien, P., Soudage et Techniques Connexes, May-June, 1966, pp. 197-218. Etienne, M., Ph.D. Thesis, U n i v e r s i t y of B r i t i s h Columbia, 1970. Burel, B.C., M.A.Sc. Thesis, U n i v e r s i t y of B r i t i s h Columbia, 1969. Jo s h i , S.V., Ph.D. Thesis, U n i v e r s i t y of B r i t i s h Columbia, 1971. Vinokurov, V.A., Avtomaticeskaya Svarka, No. 4, 1966, pp. 18-21. Di l a w a r i , A.H. and Szekely, J . , M e t a l l u r g i c a l Transactions B, Vol . 8B, June, 1977, pp. 227-236. Beynon, G. T., Ph.D. Thesis, U n i v e r s i t y of B r i t i s h Columbia, 1971 - 119 - APPENDIX A CALCULATION OF IMMERSION A . 1 The Equation The equation to be used i s p _ P I n /2L, where: R i s the apparent r e s i s t a n c e i n ohms, p i s the s p e c i f i c r e s i s t a n c e i n ohms cm. \ L i s the e l e c t r o d e immersion i n cm. and r i s the ra d i u s i n cm. A.2 Determination of the Parameters The apparent r e s i s t a n c e i s found by d i v i d i n g the o v e r a l l v o l t a g e by the o v e r a l l amperage. The s p e c i f i c r e s i s t a n c e can e i t h e r be c a l - c u l a t e d or can be used to c a l c u l a t e the immersion by use of a computer program. The r a d i u s i s a constant 0.12 cm. (diameter i s 3/32 i n c h , or 0.24 cm.). Thus i f the immersion i s known, then the s p e c i f i c r e s i s t a n c e can be c a l c u l a t e d • The apparent- r e s i s t a n c e and measured immersions are presented i n Table 9. - 1 2 0 - A.3 C a l c u l a t i o n of the S p e c i f i c Resistance I f the measured immersion f o r the f i r s t experiment i n each mode used ( i . e . , P04 for AC, P13 f o r DCRP and P16 for DCSP) the s p e c i f i c r e s i s t a n c e i s : 1) f o r AC, . 2 7 r L R 2TT(1.1Q) (.078) ~ « , « « : • t . - 1 P = l 7T = --, T7T~Tn\ = 0.185 ohm cm. ln.2L. ln,2(1.10)., (r'} L .12 J 2) f o r DSRP, 2.TT(1.42)(.047) n , -1 p = 0 / 1 = 0.133 ohm cm. l n . 2 ( l . 42) -. 1 .12 J 3) f o r DCSP, p = 2ir(1.62 ) C 1049l = . 1 5 1 ohm cm."1 P l n . 2 d . 6 2 ) , 1 .12 1 These s p e c i f i c resistances are then u t i l i z e d to c a l c u l a t e f u r t h e r immersions. A.4 C a l c u l a t i o n of Immersions The immersions of each mode were calculated using the above s p e c i f i c r e s i s t a n c e s . The c a l c u l a t i o n involves a s o l u t i o n using a IBM program and the following d e r i v a t i o n : - 121 - L = P -. ,2L. (A 9f( X ) 1 1 OT o p A - In 2L. 8L I^L 2 1 r C = X Q - f(X )/f (X ) l u o o (A Equation [A.2] can be solved by an i t e r a t i v e process u n t i l the s o l u t i o n converges to a value. The immersions ca l c u l a t e d by t h i s method are presented i n Table 18. - 122 - APPENDIX B GRAIN SIZE CALCULATION B. l Introduction The c a l c u l a t i o n of the austentizing parameter, P, requires two experimentally determined values. These values are the maximum temp- erature attained and the time above the austentizing temperature, 850°C. Figures 61-73, the temperature/time r e l a t i o n s h i p s f or CGESW of A36 plate s , are used to d i r e c t l y obtain the maximum temperature f o r the p o s i t i o n and the time above the austentizing temperature. B.2 The Equivalent Maximum Temperature An equivalent maximum temperature must be c a l c u l a t e d because the temperature/time r e l a t i o n s h i p i s not rectangular. The c a l c u l a t i o n allows for the heating and cooling cycle portions of the temperature/time r e - l a t i o n s h i p . Mathematically the equivalent maximum temperature, T i s c a l c u l a t e d E as follows: 2 RT T - T = t B . l ] AH - 123 - where T , = the maximum temperature experimentally M achieved, °K. -1 -1 R = the un i v e r s a l gas constant, 1.98 cal.mol. °K AH = the a c t i v a t i o n energy f or aust e n t i t e g r a i n growth. and re w r i t i n g [ C . l ] , T E = T M -[(RT M 2/(AH)] [B.2] and for P04 (see Table 20), T, = (970+273) - [(1.98)(1243) 2/110,000] °K E = 1215 °K B.3 The Austentizing Parameter The au s t e n t i z i n g parameter, P, i s calculated as follows f o r P04: P = r i _ _ l n . t ^ - l [B.3] LT AH  Vt  n E o ,1 _ 1.98 l n ,165,.-1 L1215 110,000 1 1 , J = [.000823 - .000018 (5.1059)]" 1 [(.000823) - (.0000919)] 1 (.000731) 1 1368°K or 1095°C - 124 - where t = 165 sec. i s the time above 850°C from the temperature/time r e l a t i o n s h i p and to = 1 sec. i s used so fi g u r e 14 can be used d i r e c t l y ( f i g u r e 14 i s p l o t t e d f o r (t = 1 sec. only). B.4 Grain Size Determination The grain s i z e i s determined from the nomogram, f i g u r e 14, where the a u s t e n i t i z i n g temperature i s obtained by c a l c u l a t i o n as i n C.3 and the AFNOR grain s i z e number i s read. The AFNOR grain s i z e i s the same as the ASTM Micrograin Size Number from ASTM E-112. The Grain Size Number equal to the number of nominal grains per square inch at 100X magnification The r e s u l t s of the above c a l c u l a t i o n s and grain s i z e determinations are presented i n Table 20. - 125 - APPENDIX C COMPUTER CALCULATIONS AND OUTPUT C. 1 Introduction The temperature at any p o s i t i o n i n the pl a t e and any time are cal c u l a t e d using equation (2.16) and the r e l a t i v e p o s i t i o n s are calc u l a t e d using equation (2.17). The computer programs w r i t t e n and used to solve these two equations are THEORY and DROG r e s p e c t i v e l y and are presented below. C. 2 THEORY Program This program was used to solve equation (2.16) with the input data from Table 17. The program generates the data required to characterize the welding procedure. This data i s stored i n a per- manent f i l e t i t l e d "run number - S" (e.g., 05S ). The program also generates the time temperature r e l a t i o n s h i p data required to plot thermal p r o f i l e s and isothermal p r o f i l e s . Some of t h i s data i s stored i n a permanent f i l e t i t l e d "run number - SI" (e.g., 05S1). Two pri n t e d p l o t t e r outputs are also generated to evaluate the input data and c a l c u l a t i o n s before the f i n a l p l o t t i n g and contour mapping i s undertaken. , An example THEORY program i s : - 126 - QS (3) , Y ( 10) #TI.1E (2 00) ,POSN (3) REXTE3?1 (3) , RADIUS (3) SQTERM (3) , BESTE?.(3) , BFUN (3) , CONTEB (3) , TE&P(1Q,200) 5 (3) , POSIT (200) , T{3), TT (3) , TEMPI (200 V 4) PRSH, VOLT, AMP, VEL, AK, CEPE, RHO, DTI^S, THICK, E? SLDE?, AI33ER, POOLDP, SORS1, S03S2, SORS3* SHTC, START DIMENSION DIMENSION DI*"NSTON RFAO (S, 1) READ (5,2) READ (5, 3) 1 RTGAP, Y(1)=3 DO 35 J=2,8 Y (J)=Y(J-1) +1. CONTINUE POWEE=0„2**VOLT*AMP*EFP SPWE=POWEE/THICK TFK= T./(2. *3. 1415S264*AK) PWBS1=SOBS1*SPWH/100„ PWRS2=SORS2*SPRR/100„ PBHS3=SORS3*SPWR/100. QS ( 1) =PWRS1*TPK QS(2) =PWRS2*TPK QS<3)=PWES3*TPK BEE=2» *SHTC/ (THICK^EHO^CEPE) DIFFUS=AK/(RHO*CEPE) BEYA=BEE/DIFFUS • SE?1=(SLDEP*AIMMER)/2. SEP2=SLDEP+{POOLDP/2.) X1=START*VEL Tia E ( 1) = STA RT S (1) =~SEP1 S(2)=0. S (3) =SEP2~SEP1 TT(1)=-SEP1/VEL TT(2) =0- TT (3)= (SEP2-SEP1) /7EL E LG A P= SL D E P— AIM 3 E R Q44=POWER/(EFF*7EL*10 0 0.) H? 173 WRITE WRITE SPITE v? SIT E WHITE WRITE WRITE WRITE WRITE WRITE WRITE R-,RITE T I T W?.I73 WRITE WRITE WRITE * r i -T E 77172 WRITE 6,5) 6,6) 6,7) 6,8) 5,9) 6,10) 6,11) 6,12 6, 13 6,14 6, 15 6,16 6, 17 6 ,18 fi , 19 6,2 3 6, 21 6 ,22 6, 23 6 ,2 <4 6, 25 6 ,2fi 6,27 RHO AK CEPE DIFFUS VOLT AMP SPWR THICK Q44 EPF, POS2R SHTC (QS (I) ,1=1 ,3) V EL SLDE? POOLDP AI1"SR EI, GAP S(1) , S(2) , S(3) - 127 - WRITE ( 6 , 2 9 ) T T ( 1 ) , T T ( 2 ) , TT (3) H SITE ( 7 , 3 0 ) BBITE (7^11) DO 60 K- 1 ,200 POSN { 1) =X1-SEP1 P0SN(2)=X1 POSN (3) -POSN (1) +SEP2 c DO 50 J=1,8 C DO 40 1=1,3 EXTESH ( I ) = { - 0 . 5 * V E L * ? O S N ( I ) ) / D I F F U S RADIUS (I) =SQHT( POSN (I) *PQSN (I) + Y (J) *Y (J) } SQTEan.(I) = { ( {VEL*7EL) / {4. *DIFFUS*DIFFUS} ) *-EBTA) BESTEH (I) =RADIUS(I) *SQRT (SQTERH (I) > BFUN (I) =BESSK0 (BESTEB (I) ) CONTER (I) =QS (I) *EXP (EXTERH (I) ) *BFDH (I) 40 CONTINUE C TESP (J,K)=PREH*CONTER{1)+CONTER{2) +CONTEH{3) 50 CONTINUE C TIHE (K + 1) =TIME (K) *DTIME X1 = X1- VE.L*DTrS3 T(2)=TI:1E (K) POSIT{K)=POSN (2) WRITE (7 ,32) T ( 2 ) , POSIT (K) ^ (TEMP (J , K ) ,J= 1 ,8) 60 CONTINUE C AA=. 5 CALL DRAWC (TEMP , POSIT , J , Y) CALL ALA3L(HU3RUN,VOLT,ABP, SLDEP,AIHttEB,POOLDP,VEL,AA,T3ICK,Q44) CALL OT7TPLT (TEflP, POSIT ,VEL) CALL ALA BL (NU.'lEUN, VOLT, AMP, SLDEP f AIMMEH , POOLD?, VEL , A A, Tf!ICK:,Q44) CALL PLOTND STOP • . 1 FORM AT (14) 2 FORMAT (8F10. 5) 3 F O R « A T ( 8 F 1 0 . 5) 4 FORMAT(4F10.5) .5 FOH.1AT {' * * * * * * * * * * * * * * * , / / , 1 X , I 4 , / / , i x , , / / / / ) 6 F 0 R M A T ( 1 X , T H E FOLLOWING ABE THE•THERMOPHYSICAL PROPERTIES * * 1 * * V / / / ) 7 FORMAT ( I X , • DENSITY = ' , ' 3 X , F 1 0 . 2 , 'G!1-C?1--3 t , / / / / ) 8 F033AT(1X,*THERMAL CONDUCTIVITY = » , F 1 0 . 5 , • C A L . S E C - - 1 C M . - 1 D E G . C - V 1 . / / / / ) 9 F O R M A T { 1 X , » S P E C I F I C HE AT= ' , 6 X , F 1 0 . 5 , * C A L - G M . - 1 D E G - C . - 1 » / / / / ) 10 FORMAT(1X,'THERMAL DIFFUSIVITY = * , ,1X, F 10-5 , 5 X , « C 3-2 SEC--1 ' , / / / / ) 11 FORM AT (1 X, 1 **** THE FOLLOWING ARE THE -T EL D IN G CONDITIONS ****',/// 1 /) 1-? ?C?-'.\7 { 1X, ' VOLTAGE - 1 , 1 3.(,? 1 0 - 1 , 3 X VOLTS ' ,////) 1 3 FO 3 M AT (1 '( , * A "1 ? ER A G 3 = , , 1 ? X / F 1 0 . 0 , 5 X , * A M P S * , / / / / ) 14 FOHMAT(1X, ' HEAT IN PfJT PER UNIT THICK NESS» , 2 X , F 10-0 , • CA L. SEC- - 1 * , / ) 15 FOPMAT(1X,'PLATE THICKNESS = ' , 5 X , F 1 0 - 2 , 5 X * CM- * , / / / / ) 16 FORM A T ( 1 X , ' S P E C I F I C POSER INPUT = » , F l 0-0 , » KCAL. CM - 1 » , ////) 17 ^i" R 1 AT (1 X, ' K EAT INPUT AT ' , 2X , F 4 . 2 , 2 X , • =» , 2X , F6 . 0 , 2 X , • C AL-SEC. - 1 CM. I " 1 ' , / / / / ) 13 F O R M A T ( I X , » S U R F A C E HEAT TRANSFER COSEFICIENT = » ,E 1 0 . 3 , 5 X r » C A L . C M - 1 - 2 . S E C . - 1 D E G - C - 1 » , / / / / ) - 128 - C c c c c 19 20 21 22 23 24 25 26 27 1 23 29 30 31 32 FORMAT FORMAT FORMAT ?C. 3 M \ T FORMAT FORMAT FORMAT FORMAT FORMAT FORMAT FORMAT FORMAT FORMAT 1X,'TBE THREE HEAT SOURCES IN CAL-CM. -15 EC. -1 ARE;*,////) 1X,10X,»Q(1)«, 10X,'Q(2)' , 10X, 9Q(3) » ,//,3F10 .0) 1X, 'WELDING VELOCITY = *,5X„F8„5 0.5X,* CM.~SEC»- 1* »////} 1X,'SLAG DEPTH = ' , 10 X , F'4 . 2 , 5X a ' CH» ,/) 1X,'P00L DEPTH =» , 10X,F4.2,5X,'CM B,/) 1 X, 'IMMERSION =» ,10X,F4. 2 f f5X, ,CB* ,/) 1X,' ELECTRODE GAP =*>8X,F4.2,5X,»CM',////) 1X,'TS£ RELATIVE POSITION OF THE HEAT SOURCES : %////) - A 1 X , Q ( 2 ) « , 2 X , F 6 - 3 t 2 X # ' C H . " . /, 1 n , ' Q ( 1 ) ' ,2X,F6.3,2X,*CM« X,«Q (3) »,2X,F6- 3,2X,'CM„ B IX,'THE RELATIVE TIMES OF THE HEAT SOURCES :»,////} 1X,' Q (1) • ,2X,F&. 1,2X,» SEC B,/,1X,°Q (2) «,.2X,F6. 1,21, »S2C», 1X,«Q(3) %2X,F6. 1,2X, "SEC* ,////) 1H1) 1 X, ' TIME' , 3X, ' POSITION' , 5XV* THERMOCOUPLE POSITION 8 ,/„ 1 X* « SE A 1C0NDS«,3X,»CM',3X,» ONE CENTIMETER INTERVALS FORMAT {IX, E10. 2*4X,F8-3, 4X, 1 0F8,0) BHD SUBROUTINE DRABC (TESP, POSN,J, Y) DIMENSION TEMP (4,200) <, POSN (200), T ( 4 ) , Y(10) DO 20 K=1,200 DO 10 J=1,8 TEMPI (K,J)=TEMP (J,K) 10 CONTINUE 20 CONTINUE XHIN=POSN (1) DO 30 K=1,200 POSN (K) =POSN (K)-XMIN 30 CONTINUE ,//> TEMPI (200,4) x r i ? i = ? o r > N ( 1 ) XMAX=POSN (200) DX=(XMAX-XMIN)/10. YMIN=0.0 YMAX=5-0 ASELT1=7 2 3- AMELT2=1450. 40 DY=5./7. DO 40 K= 1,200 POSN (K) = (POSN (K) -XMIN) /DX CONTINUE DO 45 J= 1, 4 Y f 7) - v (J) /D Y 4 5 CCKTINUE CALL AXIS(0.,0 1 XMIN,-DX) CALL AXIS (0. ,0 1 CALL O'TOUR (?OS M,200,Y,4,TEMPI,200, A MELT 1 ,4. , A MELT 1) CALL CNTOOR (POSN,200,Y,4,TEMPI,200,AMELT2,4.,AMELT2) DO 70 K = 1,200 ,'POSITION ALONG WELD CENTEHLINE', - 30,10_,0„, ,»DISTANCE FROM WELD CENTEHLINE*,29,7-,90.,YMIN, 129 - POSN (K) = POSN (K) *DX 70 CONTINUE DO 80 J= 1,4 Y (.7) =Y (J) *D Y 80 CONTINUE RETURN END SUBROUTINE A LABL ( NUMRU N, V ,AM P,SDff E, PD, VI? , AA, DELTA, Q44) RUNNUM=FLOAT (NUMRUN) HT=.07 BE=AA>2.25 CC=B3+0.5 CALL SYMBOL (AA, 6.5, HT, • EXP o 5,HT,BUNNUL1,0. , - 1) SYMBOL (AA,6. 25, HT,* VOLTAGE* a 0..„7) SYMBOL (CC, 6» 25, HT, 9 VOLTS', 0 - ,5) NUMBER (BB,6*25,HT,Y,0. , - 1) ,6. , 6 < SYHBOL(AA SYHBOL (CC NUMBER (BB, 6 SYHBOL (AA,5 SYMBOL {CC„5 NUMBER (BB, 5 SYMBOL (AA,5 CALL NUMBER (BB,6 CALL CALL CALL CALL CALL CALL CALL CALL CALL CALL CALL CALL CALL CALL CALL CALL CALL CALL CALL CALL CALL CALL !IUM3ER" ,0. 10} 0,HT, ' AMPERAGE' , 0.p8} ,0,HT,*AMPS',0.,4) ,0,HT, AMP,0. , - 1) 75,HT,'SLAG DEPTH*,0.^10} 75,HT,'CM.',0.,3) 75, HT, SD, 0. , 2) 5,HT,'IMMERSION',0.,9) 3 ) SYHBOL (CC, 5.5, HT, •CH.1 ,0. NUMBER (BB,5.5,RT,E,0-,2) SYMBOL (AA,5.25,HT,'POOL DEPTH* „0. , 10 ) SYM BOL (CC, 5. 25, HT, »CM. • , 0- , 3) NUMBER (BB,5.25,HT,PD,0. ,2) SYHBOL (AA,5-0, HT, ' WELD VELOCITY' „ 0-, 13} SYMBOL (CC,5. 0,HT, 'CM. SEC—1 ' ,0. ,8) NUMBER (BB,5.0,HT,VW,0-,3} SYMBOL (AA, 4. 75, HT,'THICKNESS',0., 9) SYMBOL (CC,4.75,HT,'CM.' ,0. ,3) NUM 3EE (33,4. 75, FIT, DELTA, 0. , 2) SYMBOL (AA,4.5,HT, 'SPECIFIC FOHEE INPUT*,0-,20) SYMBOL (CC,4.5,HT, 'KCAL.CM.-1*,0.„ 1 0) 10 CALL CALL NUMBER (BB,4. 5„HT,Q44,0. , - 1) RETURN END SUBROUTINE OUTPLT (TEM P, POSN, VI?) DIMENSION TEMP (10, 200) , POSN (200) „ DO 10 K= 1, 20 0 TIME (K) = POSN ( K ) / V W CONTINUE Y(10) r TI3E (200) „ H(200) XMIN=TIM E{1) XMA X = TtME (200) DO 20 K=1,200 TIME(K) =TIME(K)-XMIN 20 CONTINUE XMIH=TIM E(1) XMAX=TIME(200) DX= (XMAX-XMIN)/10, - 130 DO 10 K=1 , 2 1 0 Tia.E(K) = T I 1 E (K)/DX 30 CONTINUE C y«iN= o . o YKAX=1400. DY=200. CALL PL0T{14. „ C L , - 3) CALL AXIS (0.,0-,»TINE IN SECONDS *, - 15,10_ , 0.,Ximi rDX) CALL AXIS (0o „ 0. , ? TESP. IN DEG. C. * , 15, 7. , 90» , YHIN,D Y) C DO 50 J=1,8 C DO 40' K= 1, 200 W (K) =TSHP (J,K)/DY 40 CONTINUE C CALL LINE (TIHE, 9, 200, + 1) 50 CONTINOE C RETURN END - 131 - C.3 THEORY Program Output The complete output data from a computer run of THEORY r e s u l t s i n the creation of two f i l e s and two p r i n t e r p l o t t e r outputs. The l i s t i n g of two example f i l e s 05S and 05S1 and the corresponding p r i n t e r p l o t t i n g s follow below (this output i s f o r experiment P05). t l _C5S 2 c 10 11 _12_ 13 1 5 16. I T J 8 _ « « T K S F C L L C W I NG A R E T H E T H E ? !*C P H Y S I C A L P R C P E S T I r S • « • « • C E N S I T Y 7.5CG>.C*.-3 Z I 2 2 2 3 2 6 2 7 2 0 T H E P ! » A L C C N C L C T I V I T Y » 0 . I C C O C C J L . S S C . - 1 C M . - I C E G . C - 1 S P E C I F I C H E A T = C. 14300C »L.GP';-ICEG.C.-I 2 9 2 9 _ 3 0 _ 3 1 3 2 3 3 _r>f_»UL C I E F L S I V I T V •-• • o . c a i a o c * . 2 s £ c . - i 3 * 3 5 3 6 3 7 3 S 3 9 4 0 41 '2 '"43" * * * * T H E FCLlCWINfi AHE T H E ELCING CCNOITIONS • •*• VOITACE = .42.0 ,'. VOLTS 45 »fF;P<GE 4 9 0 . AMP S '- '. : / . . . 4 7 • . v ' - - ; 4 8 49 50 51 H E A T I f i F U T P E P U N I T T H I C K N E S S 5 7 5 . C A L . S E C - 1 r H O T I I M C * , V £ S S J J ' E C I F K »r>t,fl» INPLT 3 5 3 . XC A l . . C r< - 1 LI 63 65 F - F A T IMPUT AT C . 3 7 = I U 2 8 . C » L . S - X , . - I C M . - I 6 7 S L P F J C ' f ' F E A T " T R A N S F E R " C C E E M C ' l E r i T " = 0 . 5 C U P - C 3 7 c: 71 C A I . I . X . - 2 . S E C . - 1 n c G . C . - l 2 ........ ' H E I h P E E J 5 S £ C J S _ J i L l i i . C M , - 1 S S . C . - 1 A » l : 7 h 7 7 nr. _ai e?" fl:i HE S(> Al ' ae sc " i 2 2 0 . 4 5 7 . . E L D I N G V ' U r C I T Y = 2 2 9 . 0.C1400 ii.Hr, "tf i»TM =" t-coi r E P T h a 3 . i a 1 .ta I P H E R S I C F = •• SLECfRlifcf'c/.f «' Q.<-,i 2 . 2 5 C M C M CP' C ' t . S F C . - l 1<, <;5 ' * ' ' I t - E _ R E I A T I V E P O S I T I C N Of THE F - E A T S O U R C E S S ':6 i si 1 'J2 1 0 1 I C ' i 1 0 5 1 0 6 1 C 7 I C i i 1 OS 1 I 0 1 1 1 c m - 2 . C E 5 c> : 0 ( 2 ) _ 0 . 0 C f . Hit " l . S Z S C > . T H E ' R E L A T I V E T I P E S C F T H E H E A T S O U R C ' S ISI*"--1_L'1 HEFHCCCUFLE PCS !T ICK .'. GNE CENTIMETER INTERVALS" 18 I S 20 21 ' 22 C.16E*C4 <>• 16E * C 4 C.lFf •O'i :6k 0.1EE*04 1 7 . 3 6 0 I 7 . CSC ' 1 6 . 8 0 0 ' It.520 • 23 :i<o.\et*a it.no 2* 0.1SE«C4 15.96C 2 5 2 6 2 7 " 2 3 2 9 • :; -J 0. KEtC4 30 31 32 33~ 3* 35 _0.1<;E«C4 0". 1SE«Q4* .1SE*C4 i ; . e a c 15.40C 15.120 0.2CftC4 ; 14.56C C.2CEK4 14.280 36 37 3B 3<T 40 1 • C*C4 F.t04 "tC4 0.2 1E*0<: 0.21E«04 0.2OF »C4 ,i-!&'0.?0E«04 S; _ 1 E « 0 4 0 . 2 I E » C 4 .' 0 . 2 1 E K 4 . 0 . 2 1 E » C 4 " 0 . 2 ? E * C 4 ' ' . i 5 y o . 2 2 F ( C4 ' _4 4 2 43 u.ccc; 13_.720 13.4*0 13.160 12.860 0.22EK4 0.22ftC4 12.320 12.040 11.76C 11.460 11.2 " 44 _J 0. • "' "" 4 5 • . 0. «V0.22E»C4 23E»C4 23E»C4 :': 0.23f «C4 : •- 0.23F< C4 0 0 ?2C 1C 1C 1C.360 IC.'CEO S.BOO Co 4 H 0 . 2 3 E * 0 4 9 . 2 4 0 2 2 3 . 2 1 9 . 2 1 2 . 2 C 7 . 4 9 0 . 2 4 E K 4 8 . 9 6 C 2 3 6 . 2 : 1 . 2 2 6 . 2 1 9 . 5 0 0 . 2 4 F . . C 4 6 . 6 6 0 2 5 0 . 2 4 5 . 2 3 9 . 2 1 1 . 5 1 0 . 2 4 F « 0 4 8 . 4 0 0 2 6 6 . 2 6 C . 2 ' - 3 . 2 4 5 . 5 2 0 . 2 4 F K 4 8 . 1 2 0 2 8 3 . 2 7 6 . 2 6 8 . 2 5 9 . y— » 0 . 2 4 E K 4 7 . 6 4 0 3 0 0 . 2 9 3 . 2 8 4 . 2 7 4 . ' 5 4 0 . 2 5 E K 4 7 . 5 6 0 3 2 0 . 2 1 2 . 3 0 2 . 2 9 C . 5 5 0 . 2 5 E . C 4 1 . 2 8 0 3 4 1 . 3 3 1 . 3 2 0 . 3 C S . 5 6 0 . 2 5 E . C 4 7 . 0 0 0 3 6 3 . 35 2 . 3 4 0 . 3 2 6 . 5 7 0 . 2 " ; f * 0 4 6 : 7 2 0 " 3 8 7 . 37 5 . 2 t 1 . 3 4 5 . 5 8 0 . 2 5 E * C 4 t . 4 4 0 4 1 3 . 4 0 0 . 3 G 4 . 3 6 6 . 5 9 0 . 2 t E « 0 4 6 . 1 6 0 4 4 2 . 4 2 6 . 4 0 6 . 3 6 8 . 6 0 0 . 2 6 E . C 4 5 . 8 8 0 4 7 2 . 4 5 5 . 4 3 4 . 4 1 1 . • — t t 0 . 2 6 E . C 4 5 . 6 0 0 • 5 C 5 . 4 E 5 . 4 6 2 . 4 3 6 . 6 2 0 . 2 6 E . 0 4 . 5 . 3 2 0 5 4 1 . 5 1 6 . ' 4 9 1 . 4 6 2 . 6 3 , 0 . 2 6 E K 4 . 5 . 0 4 0 5 8 0 . 5 5 3 . 5 2 2 . 4 9 C . " " " " ' 6 4 0 . 2 7 E . C 4 4 . 7 6 0 6 2 2 . 5 9 1 . ' 5 5 5 . 5 1 9 . 6 5 0 . 2 7 E . 0 4 4 . 4 3 0 ' 6 6 8 . 6 3 1 . 5 9 1 . 5 5 0 . 6 6 0 . 2 7 E . C 4 4 . 2 0 0 7 1 8 . 6 7 4 . 6 2 8 . 5 6 2 . 6 7 0 . 2 7 E . C 4 3 . 9 2 0 7 7 2 . 7 2 C . 6 6 7 . 6 1 5 . 6 8 0 . 2 7 E . 0 4 3 . 6 4 0 8 2 9 . 7 6 6 . 7 C 8 . 6 5 C . 6 9 0 . 2 E E « C 4 3 . 3 6 0 0 9 1 . 6 1 9 . ' 7 5 0 . ' 6 6 5 . 7 0 0 . 2 8 E . 0 4 . 3 . C 8 0 • 9 5 7 . 6 7 2 . . 7 9 3 . 7 2 2 . 71 0 . 2 e E « C 4 2 . 8 0 0 1 0 2 5 . 9 2 6 . e s s . 7 5 ? . 7 2 0 . 2 8 E . C 4 2 . 5 2 C : 1 0 9 5 . 9 6 2 . 6 6 3 . 7 9 6 . 7 3 0 . 2 8 E < 0 4 2 . 2 4 0 1 1 6 . 4 . 1 0 3 7 . 9 2 1 . 8 3 3 . 7 4 0 . 2 9 F . C 4 1 . 9 6 0 1 2 3 2 . 1 0 9 I . 9 7 2 . 8 7 C . 7 5 0 . 2 9 E * C 4 1 . 6 8 0 . 1 2 9 7 . 1 1 4 5 . 1 C 1 6 . 9 0 6 . ' 7 6 0 . 2 9 E « 0 4 . 1 . 4 0 0 1 3 6 0 . 1 1 9 6 . 1 C 5 8 . 9 4 2 . 7 7 0 . 2 9 E . C 4 1 . 1 2 0 1 4 2 0 . 1 2 4 5 . I C 9 8 . . 9 7 5 , 7 8 0 . 2 9 E 1 C 4 C . E 4 0 W 1 4 7 9 . 1 2 9 1 . 1 1 3 6 , 1 0 C 7 , 79 0 . 3 C E » 0 4 C . 5 6 0 ;. Y: 1 5 3 3 . ,-! 1 3 3 4 . 1 1 7 1 . 1 0 3 7 , •'• • • . : , • • 8 0 0 . 3 C E K 4 £ . 2 8 0 1 5 8 0 . 1 3 7 1 . 1 2 ' J 3 . 1 0 6 4 . - 81 0 . 3 0 E . C 4 ' . - e . e o o 1 6 1 8 . 1 4 C 3 . 1 2 2 0 . 1 C 6 6 , - 6 2 0 . 3 0 E + C 4 - 0 . 2 8 0 ' 1 6 4 4 . 14 2 8 . 1 2 5 4 . 11 1 C , 8 3 0 . 3 C E . C 4 - C . 5 6 0 1 6 5 9 . 1 4 4 6 . 1 2 7 2 . 1 1 2 0 . 8 4 0 . 3 1 E » C 4 - C . 8 4 0 1 6 6 5 . 1 4 ! E . 1 2 6 7 . 1 1 4 3 , 6 5 0 . 3 1 E » C 4 - 1 . 1 2 0 1 6 6 3 . 1 4 6 4 . 1 2 9 7 . 1 1 5 4 , . £ 6 0 . 3 1 E » C 4 - 1 . 4 0 0 : U 5 5 . 1 4 6 5 . 1 3 0 2 . 1 1 6 3 . . : ; ' . 0.3iE « C 4 - 1 . 6 8 0 . .•. • • U 4 2 ; ' 1 4 6 C . " 1 3 C 4 . " U t b . "' " . ' ' ' y:» " ...,T „. 6 8 0 . 2 1 F » C 4 - 1 . 9 6 0 ". •!•'.; 1 6 2 2 , , 1 4 5 1 . . 1 3 0 1 . 1 1 7 0 , : ; , . : i i : ' : > ' . { • ' :• •• '• •• 8 9 0 . 3 2 E < C 4 - 2 . 2 4 0 ' ' 1 5 9 6 . 1 4 3 7 . 1 2 9 6 . U 7 0 . ' '••''" • . ' : : ; ; : 9 0 0 . 3 2 E . C 4 - 2 . 5 2 0 • : ' . - 1 5 6 5 . 1 4 1 9 . . 1 2 8 7 . : U 6 7 , - . ' • . ; •.•' ~ . i ; > - . ! ' • - . ' • , 1 " 9 1 0 . 3 2 F t C 4 , - 2 . 6 0 0 . ^ 1 5 3 1 . 1 3 9 9 . • 1 2 7 5 . 1 1 6 1 . • . " i ; : ^ ; ' . ' - ; C r S 1 , • ' • •.'"'' ' -; >•"'' 'V.. :• 9 2 0 . 3 2 E . C 4 -3 . C 8 0 •".''.'< 1 4 9 4 . 1 3 7 6 . 1 2 6 1 . 1 1 5 4 . . - ; - ' • . - i vH ''• ' . : • : . • . : : v . 9 3 0 . 3 . 2 E K 4 -3 . 3 6 0 :•• 1 4 5 6 . 1 3 5 1 . 1 2 4 6 . 1 1 4 5 . '"" " """ ."-T."••?~T~TTr~~'~:7~r~~~~r ~'""" 9 4 0 . 3 3 E U 4 -3 . 6 4 0 ' ' • 1 4 1 9 . 1 3 2 5 . 1 2 2 9 . 1 1 3 5 . ' • ' ' . . 9 5 0 . 3 3 F » 0 4 - 3 . 9 2 0 1 3 8 2 . 1 2 9 5 . 1 2 1 1 . 1123. " • 9 6 . ' 0 . 3 3 E » C 4 - 4 . 2 0 0 , 1 3 4 7 . 1 2 7 3 . 1 1 9 2 . 1 U C . . • . 't :• •• 9 7 . 0 . 3 3 E * Q 4 . , - 4 . 4 0 0 . • . - 1 3 1 4 . 1 2 4 7 , 1 1 7 3 . • 1 0 9 7 . . ' , ' . ! " ' . • • • ' . • • • • : • ' • ' : * ' • ' ' • . . . ' ; . ' : . , ' ' . • . ' • • : • • • • . • • ; " . ' , : ' 9 6 . 0 . 3 3 E » 0 4 - 4 . 7 6 0 ' ' I 1 1 2 8 2 . M 2 2 2 . 1 1 5 4 . . 1 0 8 3 . ' ' - , ' • ' ' . : - ; • ( ' ; • > " • • ' ' . • . ; , . iz.' : .' 9 9 0 . 3 4 E » C 4 - E . C 4 0 . 1 2 5 1 . ' 1 1 9 7 . " " " 74 1 3 5 . • 1 0 6 9 , • . .- -•• * , V i ; : ^ , ' . . ' . ^ ^ ; - : ' - 1 0 0 0 . 3 4 E . 0 4 • - J . 3 2 0 - ? : 1 2 2 3 . 1 1 7 3 . ' 1 1 1 6 . 1 0 5 4 . > • : ; ; ; " , : '"' 1 0 1 0 . 3 4 E t C 4 V 5 . 6 0 0 1 1 9 5 . 1 1 5 0 . 1 C 9 7 . 1 0 4 C . ; . '• :':t: '.•''•-' 1 0 2 0 . 3 4 F ^ C 4 -5 . 8 8 0 • 1 1 6 9 . 1 1 2 8 . 1 0 7 9 . 1 0 2 5 . . " , ; ' " 1 0 3 0 . 3 4 E . C 4 - 6 . 1 6 0 1 1 4 4 . 1 1 ( 6 . 1 0 6 1 . l o i c . • . ' . y ' ••••••• I C 4 0 . 3 5 E K 4 - 6 . 4 4 0 1 1 2 0 . 1 0 8 5 . i 0 4 3 . . .. . . . U 5 ' " C . 3 5 f > C 4 ~ ~ ~ - 6 . " 7 2 0 " t ' 0 9 7 " . 1 0 6 5 . " 1 0 2 * ; " ~ 1 0 6 0.35F .04 - 7 . 0 0 0 1 0 7 5 . 1 0 4 6 . 1 C C 9 . 9 6 7 . ':' y '• <Jl ' ••; . •. U 7 0 . 3 5 I . C 4 - 7 . 2 8 0 ' 1 0 5 5 . 1 0 2 7 . 9 9 3 . 9 5 3 . ':!• - 135 - O -Jj m U Ci; iT. -r r \ o « r>- O C !? 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LU U J -4* o* >r cj* m o |m m i n ^ ^ , 0 0 0 0 0 0 m o >r I M o n o tr I M I T co o rvi rv m m pn |(NI P4 pg W AJ f\J I t i l < UJ HI Ul u tu U.' o o 0 0 o o o 0 0 0 o c j O o 0 | 0 o o O O O j O o o i O CJ 0 0 o o l o o 0 , 0 o o o o o o o o i O o o o o o 1 «C.r- en D-.] r m L o r— cc V o . > m m m m -jj - 136 - • r o -o <M r r - r— <i< ~u C7> vl" O ' f - r - fvj ^* NT! -O tr\ !u» m tM oi -r o , f i o* m —* co «> - r >r m f i •C -J- »T >T vT V i n ui <r sr •* - r ^ m w i t n IM y u . u-1 nj r>4 - J — - o j O O U*.(J» 0* co. >r >r >r *r .j- , -o ft <-» r» J _ . - fl W n f l ft ft ft V in W CO tn r\l| *r >r ,f> f t m J -X}—» a i n w j i*0->J' n u T> i n ft;«—1 co -^J f> —• (*" >r I'W o * * i > -J" -J" -J flj«"> f l .-a f»J;<M —• —• <-» — I U O >-> O W* u 1 i n f t r*\ f» m ,1 m M f i j f ft f t ft m f i i f t f i f i <M c\* \ INJ tr -*J m i o f~ J O r» N (vr"- sO -O <r> ft ro ft rn f» f i ; f i f i ft f i f i !—• uj -j- —> mj <NI 1 O B Ul M J j ^ - r * <M i\j ,-o ;;\i r\j —. , , — i ' 0 rj o T1 -J* i ni w m m n IM I\I | v *r —• C* - ft n i ni (v : - ui ft -— ca j") t-J O O O KT !ft f i **i ft ,—< rsj ft C >0'm U" O —* O O O C 1 0> .# -r!<j- f i f l ! C -O <"i o J i 33 H W tt) f i ft f i f l f l ftj i O O | 0 c i »r <NIIO a i -a v\e<4 s o o o o O I <r f\i o co o • O f l ) CO —I • •o vo o r - f o o o ( M O CO CD fO[ U l LU Lu i O O O O O O l ) O O ( ( I j I I t LU UJ a l Uj l o o o , o o o p O O ' C J o o O * ^ M o C O f l %0 (M «T p o o p o o ^ —• » u\ rg CT» r - r» & *i> ^} u>| ft f | ft (* O O O t_) O O •o «r M U D o . O CI * J) - j J f l ft f l f l -T 1 fk f l f> f t f f I I I I I O O O O O O i •o ft CD tn f i u i m m ^ ^ ^ f i f*i f t ft ft f*> p o *-» o o < rsi O co sQ » -r f— O ip L <f ir- ^ L I L I I t I I in u i ui JI L , u o <- o ' o o c o vf\ <j r - »f f l ft f l f l (NI f l f • ft ft f l ft IA IM O N lA i*t fM (Nt t\i — i —» ft ft f l f l f , f l O O fM O O C O O O ft n ft f i ft OJ ' <J *J> (Ni » O O r I I ; ) I o o o p o o CC C O —• <\i f i ^ p u o a o —« —• <Ni r\J t\i <\i| •t IT O N CC CT |o O O O O o U rw , \ N rg o p o o o o iM a co o .-M U» .CO t_» f ^) u* ii m m in v l u o o p o i ( j | G O O O O O 0 | < " C ^ r v o C <v;k r * o m * co j tnjw- L/. o w o o - — •— - • ^ ^ ^ ^ l I I i o o o o o o - —• rvj :\jtf\j r\j r\j r\. *\t j 'M f\J N |(M (\j fvj N (\) N - 137. - 1<J* 'J* CO J J r - T— i ~ i ~ I c M f M P g p g C M r g l r g r M p g f M r g -r r\i:i_) co -»J -T n ' r\i fM rg rg .-M pgj pg f\i N O -T W i t ' j u* u* y 3 cO| j CM CM CM 1M ."Mi fM r\i <M; (M (N) pgr £ j :o r - '(*• r*- r - i r - ^ -ii -i) ^ -o •* CM (M r\l f\l CM-CM CM <M (M rg (M o in ui „ i j i IV (M <N1 (N( fM I c c* 0*- tr* co cO' | CM CM CM; CM eg fM p O O j O o o (—• -j- r - o CM m I A c i — ! j> r - u su 3 3 '-O P - i— r- rg rg rg. r-j pg rg O O O O o o | M O CO - f CU —t <J LP (Ml l O O O j O O O CO CT r*:—> (M ( iM (M CM CM Pg ( >]-0 «u -Oj-O -U M J o C J o o o o r m «o!p- cu fji •* f> f> f i ft f i J (M CM. CM. IN) fM f i —. u*'r— u- r * f— -Oi-O gj -o! (M (M CM-CM (M-rg; O O O p *-* O fM O CO !-0 «J- fM O u U O u O - * • * : * * • ,UJ LU M , L U Ui U l (M l ^ (M CM ft f i o —• o 0 0 0 f— *0 ^- (M —; pg fM fM fM fM CM fM (Ml m f i tMJ o ^ r - m f i (MJ o o» r*. in 1 m <r -r ^- f i ft n f t i- >: 1 CM CM IM rg <M (Ml CM (M (M fM (M (Mj I fM fM fM fM fM (Ml fM O COr— Ut f o 4 m m ui 1 (M IM (Mi fM CM ( O O O l O o o O CD gj.gr- fM C 1 fM -T r— . 0 en gj —• o cot -o u* m tn sr: -r 4-] (M (M CM'fM n o o oj«-» o O 3D »0 -J-!fM O CO aj — *j-|r- a fM a o o | o o o O —• fM.fi -J* m h O ^- aj'tr- c — •j- v *f'*T g* \<r -r -T in m ru «. r\j rg.fM fM (M: 0» l U O i A f l ft f i ft f i - m mi »M CM (Mi (M- fM fM O CT I— -« ^ f l f i (M CM! fM fM (M CM CM CM: CM CM CM 0 0 " r\» <M —• - (M CM CM' C CM (Ml fM —* t7* CO r - CM CM CM;CM fM CM 1 C co;-U ut ' 1 CM CM! CM rg 1 1 CM CM CM CM I -4 QS i CO r*> U lilfM CM CM;t>* CM » <M (M ' i O U > f * | * CM CM CMl ^ HI (\l CJ O 1 o o »-> O t> tr CM CM CM CM —> -O m - r ' CM — O O u o o o | CM (M CM,CM Ii JiCM fM (MfM <"* f CM O.co;***- i n f l | ft ft ft! CM IM (M-CM CM CM O O O O O O <i >f fM O CD -O i n cu — i ^ - u tp CM —< o ! co o m f ft CMifM CM CM CM r\t f\(i(M CM CM o o CJ o a o i CM m « > I Q ft <u f i fM —• i C71 CJ 43 (M CM CMt—• —« —1 (M CM CMICM CM pg o o o o o o , rg ~T[r- KJ O o o l o O O o o o j o o O C l i t I I > o <-» o p o o m CM —1 O co -« -< o CM CM CM fM fM CM O O O O O O ] O ID 4} ^ (M O * yj -J- 1— o |<r »r u i m «> uv m u> m I • 1 l l o o o p o o r - MJ i n - f t rg - o o 0 0 o o | CM CM fM|CM CM rgI O O O O O O l to O - t M O ccl CM m cu — ^- un m t n . m m m | I l 1 I- 1 p o o p o o | J f i m >o r - •* m L A m -j-. in j« u . J (M fM *M fM fM CM CM C7* O —1 fM f- m in gj -o * ~ " CM CM fM ^ in >0 N « ? o • * J >0 >U ^ -U -tl N f ]fM CMCMrMCMCMCMC 1 rg f t -J" ir.l-o r— r - r - r— jr- I CMTM «M CM CM tn i/> o r— p- r - ro cc[3* CM Pg CM fM fM CM Pg CM CM PM CM] 28H 289 29C 291 292 293 294 295 296 "297 296 299 300 301 302 "303 " 304 305 0 . 7 1 E K 4 0.72EiC4 0. J 2 F t C 4 0.72E«C4 0. 72E «04 0. 72F.»f.< 0.73E.C4 C.73F*C4 C.i3E«C4 0.73r.C4 0.73E«C4 0. 7 4 E K 4 -E7.96C -58.240 -51.52C -58. 60C ' -59.080 - f 9 . 360 -59.64C A -59.920 - f C . 2 0 0 -6C.486""' -6C.760 -61.C40 EKO OF FILE 0.74EtC4 C.74E«C4 _ 0 . 7 4 E « C 4 . 0.74E.C4 0.i=E*C4 0.75£<C4 -fcl.320 - t l . 6 0 0 _-61.680 -6 2". 160' . -<2.4<,0 -£ 2 . 7 2 0 200. 119. 199. 1 98. 196. . 118. 197. l c . 7. 197. 1 9 7. 196. 196. 196. l i t . 195. ' 1 14. 195. . 194. 194. 193. 1 94. 193. 193. 1 92. 192. IS 2. 192. 1 SI. 19 1. 19 1. 19C. 19C. 190. ISC. 189. 1 69. I 8 i . 1E9. 166. 1 66. 168. 16 7. 187. 1 66. 187. 1 t-l. 186. 1 £5. 186. l£i. 1(5. 1 64. 165. 164. 164. 163. 1 83. 16 3. 1 E3. 162. 16 2. 16 2. 162. 181. l e t . 16 1. 180. 1 HO. 1 s c . 1EC. 179. 1 79. IB PFINAL1»»PRPL0T 5=C5 4*C5S 7=C":S! EXECL'TiCN BEGINS , . ^ O T E - THE FGLICKING STPING CGUIC NOT BE PRINTED CUE TO OVER O.0W C IKG OF THE PLOT PAGE. - 139 _ O » * * fNJ * * ft #|# tf * * » # * # * tt * -ft ft * » * * • * » * > ] •B- O ' O o 9 «» »• e- ^ o m o " . I - 'J3 O O I - 140 - * * * * • * * » ft * * 1 _. 14.1 _ C.4 PROG Program ' The DROG program was run i n i n t e r a c t i v e *IF made of f o r t r a n and thus the output from t h i s program was not i n hard copy. The program i s presented below. . C 'THIS PEOGrfiS IS USED TO CALCULATE THE IE a EESION C TEE ELECTEODE DURING ELECTSOSLAG SELDIHG,IE THE c FOLLOFJMG paaanETEas ABE KNOWN: C VOLTAGE. c A'UPEBAGE C OVE'SALL RESISTANCE "(V/I) • . C SLAG RESISTIVITY (BHO) '• " C BADIDS 0? THE ELECTRODE (RAD) C c c BEO=0-151 RAD=- 238 75 READ[5,800,END=999) E,DELE 800 FOBM AT I2T10-5) ' 76 H=0 1 E=H+1 DP~EH0/(6.283*XL**2) * {1,,-ALOG (2-*XL/EAD) ) F=BHO/(6„283*XL)*AL0G{2-*XL/RAD)-R X1=XL-P/D? DX=X1-XL I P JABS [DX/X 1)-LT-0.01) GO TO 10 I?(N-GT.50) GO TO 11 XL=X1 GO TO 1 10 WHITE (5,2.) XL,N,H 2 FORil AT(1 X, ' IHME2SI0N IN Cfl- EQUALS* ,/FIO. 5 r V r • if' ,/X3p. 1f,X, ' r,ESXSTA2JCE=» , 3X,F10_5) •GO TO 19 11 WRIT? (5, 3) 3 POn.MAT(1X#'TOO MANY ITERATIONS') 19 R-2+UELR I ? {2-GT--03) GO TO 999 GO TO 76 999 STOP END - 142 - This data ( i . e . , immersion) was used to c a l c u l a t e the r e l a t i v e p o s i t i o n s of the three heat sources and therefore became input data for THEORY. I t could e a s i l y be incorporated i n THEORY' and be run simultaneously as a si n g l e program. C.3 Computer Output Considerations The computer generated data could have been used to generate cooling curves and then these curves could be plo t t e d over a stored CCT curve and then the combined plo t s could have been used to pr e d i c t the resultant expected microstructures. A major problem arose because there are few, i f any, CCT curves a v a i l a b l e f o r the range of grain s i z e s observed. The cooling times for various temperature ranges for experiment P05 and two po s i t i o n s from the weld counterline are shown: Temperature Time Time Range (2 cm) (4 cm) 900 - 500 1000 sec. 1000 sec. 700 - 500 700 sec. 680 sec. 700 - 600 360 sec. 280 sec. 800 - 500 820 sec. 860 sec. The cooling rates are not s i g n i f i c a n t l y d i f f e r e n t f o r the 800-500° C coo l i n g time and are applicable to a curve that i s s h i f t e d f a r to the r i g h t . Thus i t would be expected that the transformation would not be a f f e c t e d as greatly f o r these slow c o o l i n g r a t e s . Even though CCT curves are not a v a i l a b l e f o r the large grain s i z e s obtained i n ESW, - 143 - the grain s i z e e f f e c t on har d e n a b i l i t y i s obviously minimized. I t should be noted that i n the HAZ adjacent to the fusion boundary a b a i n i t i c s tructure was observed (also some retained austenite and. some martensite) and thus the curve was s h i f t e d very f a r to the r i g h t f o r large grain s i z e s . E L E C T R O S L A G WELDING : Wire Porent Metol H A Z Consumable Guide Liquid Slag Liquid Metol -Penetrotion Weldment SECT ION IN P L A N E OF P L A T E Wire Consumable Guide Porent Metal Copper Shoe Liquid Slag Liquid Metal Solid Slag 5 SECTION PERPENDICULAR TO PLANE OF PLATE Figure 1 Schematic representations of Consumable Guide Electroslag Welding (CGESW) - 145 - Amperage Figure 2 Schematic f l a t welding c h a r a c t e r i s t i c 146 - *4 to I 1 'I Ul Time, mtn Thermal cycles in welding with" covered electrodes, by submerged arc, and electroslag welding: I—Heating; II—maximum temperature; III—cooling Figure 3 Schematic comparison of thermal c y c l e s ; a, covered electrodes b, submerged arc; c, ESW (Ref . 6 ) Electroslag 3 Electrode Elect ros lag 2 Electrode Elect ros lag I Electrode Submerged Arc (I Arc) Flux Cored Manual " 4 " E 7 0 2 4 Relative Deposition Rates For Welding (Ref. 13) 50 100 150 Deposition Rate (Ib./hr.) Figure 4 R e l a t i v e deposition rates f o r various welding processes (Ref. - 147 - Thermal cycle: (a) arc and electroslag welding o f steel, 100 mm thick, using one electrode; ( b ) electroslag welding o f steel, 100 mm thick, using two electrodes (solid lines), and electroslag welding of same steel using a plate electrode (dotted lines) Figure 5 Thermal cycles f o r welding conditions i n Table 2 (Ref.1.) - 148 - U o cT I . 3 o <u CL E 1100 l O O O h 800 600 400 \ \ V \ Distance From Root Gop |o 10 mm. 2X 22 mm. 3 v 20mm. 4 * 32 mm. 5 0 20mm. 32 mm A — A — A — A ^ . ^ • - • - D v / 2 0 0 1 ® 1 » 1 1 _ i 1 -20 0 20 40 60 80 100 120 140 Distance of Thermocouple From Slag Surface (mm.) Figure 6 Thermal cycles from Mel-bard (Ref. 15) 2000 0 10 T © Case I x Case 2 2 0 30 4 0 50 Distance (mm.) From Slag Surface Figure 7 Slag temperatures from Mel'bard (Ref. 15) - 149 - Weld Fusion line ' 1 Parent plate level at -HOT / / / \ 80 70 X + 0 -As welded -Stress relieved -Normalised 60 & Or LU / I I \ UJ H50 / ; / < / \ + +' / / V J \ \ Inside weld 4a / / /.X \ \ i 30 / N / 20 V- Hio Outside weld 125 - 0 6 - 0 4 -0-2 0 +02 +0-4 +0-6 +0-8 DISTANCE FROM FUSION LINE, in. Charpy V-notch impact test results for different regions of weld and H A Z , at -•• I0°C. Figure 8 Charpy impact traverse values f o r ESW from Woodley et a l (Ref.17) - 150 - D C S P lA= 2 4 5 A m p s l B= 2 3 Amps l c = 7 2 Amps ' • r o t o r 3 4 0 A m P s D C R P l A = 4 5 A m p s l B = 8 5 A m p s l c = 2IO Amps I, To to I Vo l tage = 4 - 8 Vo l ts S lag Depth = 5 0 m m . Root Gap = 2 5 m m . 3 4 0 Amps V f = 2 - 9 cm. sec7 V, = 3-3 cm.secT Figure 9 DCRP and DCSP immersion and e l e c t r i c a l d i s t r i b u t i o n (Ref. 53) - 151 - E S c o c Q_ 25 20 15 10 5 0 I 1 T— O - 43 Volts o — 36 Volts A — 3 0 Volts 200 300 400 500 600 700 800 Amperage —*- Figure 10 Penetration versuc amperage f c r various voltages (Ref. 54) E E c O OJ c O) 0_ 25 20 2 15 10 0' A—30 mm. Slag Depth • — 50 mm. Slag Depth 200 300 400 500 600 700 800 Amperage —»• Figure 11 Penetration versus amperage f o r two sl a g depths (Ref. 54) - 152 - I < o —• => > C M O ^ 12 10 8 6 4 2 0- 1 1 1 1 O -Slag H 380 Si02 26-4 MnO 101 CaO 4-8 FeO A-SlagTH" 640 Si0 2 360 MgO / i / G -Si o .A / A A / A 1000 1400 1800 Temperature (°C) 2200 c o c Q) a. 25 20 15 10 5 0 O - S l a g H A—Slag n r 200 400 600 Amperage -*~. 800 Figure 12 Penetration versus amperage f o r slags of d i f f e r e n t c o n d u c t i v i t y (Ref. 54) Actuai Conditions -Electrode .Molten Slog Immersion -Molten Interelectrode Gap (EG) 1 Metd P o o l x D e P t h f- Penetration •Heat Affected Zone FB HAZ q Slog Surface Position of Heat Sources for Model [A/ Center of 3/ interelectrode Gap Center o1 Pool Depth 3 I x u x=0 Figure 13 Schematic representation of the a n a l y t i c a l model. - 154 - 500 200 100 h 6 h 7 %S0\ I ? 20] I I I 10] 7 11 12 13 V- f 1100 V r 1200 1300 < 1400 I 7500 P'K(t0*1sec) A /Vc V 3 tod sec O 12 to 80 sec 0 7 min X10 to 20 min © 1 hr o 3hr + 9hr &H= 110000 cat/motes A L_i !_!_ 1S00 1700 1800 800 900 1000 1100 1200 1300 P'C(t0*1sec) I n f l uence of t ime a n d tempera tu re o n g a m m a gra in s i ze . 1400 1500 Figure 14 Nomograph f o r determining the austenite g r a i n s i z e - 155 - Figure 15 Thermocouple placement for i n i t i a l block welds - 156 - V e r t i c a l S e c t i o n OJ X J c o CD c a> & - 0 1 0 0 i n c h 0 E l e c t r o d e - 1 / 8 " M i l d S t e e l C o n s u m a b l e G u i d e cu T3 c o o 0_ I inch I c m . , • P o s i t i o n e r S i l i c a Insulation H o r i z o n t a l S e c t i o n — P o s i t i o n e r S i l i c a Insulator —Consumable G u i d e ' E l e c t r o d e Figure 16 S i l i c a guides f o r Consumable Guide E l e c t r o s l a g welding i c y l i n d e r s . - 157 - Hor i zon ta l V iew Figure 17 CGESW guiding head for welding and c y l i n d e r s - 158 - Horizontal Section Plane 7-62 cm. From Top Of Cylinder Surface Thermocouples) Verticol Section 11- CITI| , I inch Plane 22-9 cm. From Sump Bottom Figure 18 I n i t i a l thermocouple set-up f o r CGESW of c y l i n d - 159 - N° 13-16 N° 17-20 N° 21-24 N° 9-12 Horizontal Section N ° 5 - 8 I 20, 16, 19, I5| l 8 . '4| 17. 13, 12 .4 II ,3 10 ,2 -»^ri--2-54 cm From Cylinder Top 8 Vertical Section Vertical ) Position Of Thermocouples 2 cm i i I inch i_ I Figure 19 F i n a l thermocouple set-up f o r CGESW of c y l i n d e r s - 160 - Double 3 r Eye Loop Holder Parent Metal To Guide Plates in Fixed Head cm. 105 inch , Two Holed Alumina Tube (0-32 cm. Diameter) n Boron Nitride Sheathing (0-65 cm. O.D.) Nozzle -Electrode •A -Graphite Powder Figure 20 Slag temperature measuring thermocouple -16.1 I I cm. | 10-5 inch [ To Fixed Head Guide Molten Slag i Figure 21 Immersion measuring loops 2 Figure 22 The welding j i g for butt welding - 163 - gure 23 The eooling shoes Figure - 165 - Side View Run Out Sump Run In Sump N Top View Section (Thermocouple Hole Side) I inch i 1 Figure 25 Thermocouple p o s i t i o n i n g i n p l a t e butt welding. Front Section Plate Plate Plate To Shunt 4 -To Shunt 3 -To Shunt 2 -To Shunt Top Sect ion Parent Plate Shunt Lead 7 Asbestos inch Figure 26 Shunt set-up f o r measuring the current d i s t r i b u t i o n - 167 - Figure 27 Hobart 750 d i r e c t current power c h a r a c t e r i s t i c s 100 CO 5 8 0 d < CD 6 0 o I J l 4 0 o 2 0 0 1—: : 1 1 r— Output Volt-Ampere Characteristics Model T-500 Tap 5 0 N. v \ \ N \ V \ \ \ \ ^ \ \ \ I \ \ \ H \ \ \ 100 2 0 0 300 4 0 0 5 0 0 6 0 0 7 0 0 8 0 0 Output Amperage (A.C. Amps) 00 Figure 28 Hobart T-500 alternating current power characteristics Welder ( "Vo l tmete r 1 Recorder j 0 _ r- R4 R s Figure 29 V a r i a b l e r e s i s t a n c e current measuring apparatus INCLUSION DISTRIBUTION i f " THICKNESS 100 80 -0 60 Z) O 4 0 20 0 40 cr- Q - 20 0 LthL 0 J — 1 — 1 • 4 0 Irr-r-i - > — 1 — 1 — 1 t 1 6 8 /xm - 10. '2 14 Figure 30 T y p i c a l Canron weld i n c l u s i o n d i s t r i b u t i o n - 170 - W E L D O X Y G E N C O N T E N T DISTANCE F R O M W E L D C E N T E R L I N E ( in) igure 31 Canron oxygen d i s t r i b u t i o n and penetration - 171 - 400 Weld Oxygen Content "TT" 7/8" Thickness 200 c cn >» x O 0_' 0 400 G I G G + I 1/2" Thickness 200 0 G _L 0-5 10 1-5 Distance From Weld Centerline (in.) Figure 32 Canron oxygen d i s t r i b u t i o n and penetration 200 or LU 9 < rr >- c/> 150 or LU > — 1 1 1 ' 7"1 1—~~ i f " Thickness * M o o $ < X 200 O | o o ° 0 ° ° ! o O O O 0 n l O ° ° °" ) 1 150 ' 1 : 1 ; L O O O O | i O N N < < O X X O Thickness * o o 0 0 ° o ° o o cp o o o o o o o o o o 200 150 7 II NI NI < < X X Thickness o o o o o o o o o 0.0 o i o o o 9 2.0 4.0 •DISTANCE FROM C L . (cm) 6.0 Figure 33 Canron weld, microstructurally determined (solid line) and hardness traverse determined (broken line) HAZ 2-50 2 0 0 Electrode Velocity (cm/sec) 4 0 5 0 6 0 7 0 8 0 E X J CO _0J CD E o Q 0-50 h 0015 0017 0019 0021 0023 0 0 2 5 0027 0 0 3 0 0032 Welding Velocity (cm/sec) t-O I Figure 34 Electrode v e l o c i t y and weld v e l o c i t y f o r d i f f e r e n t weld s i z e s  Figure 36 Thermal p r o f i l e for DCSP - 100 - 110 - 120 " 130 - 140 - 150 . 160 . 170 I e l e c t r o d f e ! Figure 37 Thermal p r o f i l e f o r AC I mould - 177 - 15001 1200 900 o 600 300 / / / ' ' / / / Electrode Positive Electrode Negative -; Alternating Current O 500 700 900 — R E L A T I V E . T I M E (sec) 1100 Figure 38 Thermal p r o f i l e comparisons f o r l a r g e electrode experiments s p a c e r 6 3 t h e r r a o c o u P l e , l / 4 - c h from o r i g i n a l weld 100 90 80 *S 70 > 6 0 r | 50f I 40f 30 20 10 0 0 C 50 40 30 20- 10 0. X 3 4 5 Microns —*• 5 6 7 8 Microns— 6 10 Ii 12 13 14 15 Figure 39 Inclusion d i s t r i b u t i o n f o r DCRP 100 90 80 70 CD | 601 E 50f c3 401 30 20 10 0. 0 50 40 - £ 30 20t T J ioh 0 J L ,.,.1, „ 1 3 4 Microns- 3 4 6 7 Microns 8 10 12 13 14 15 Figure 40 Inclusion distribution for DCSP lOOr 9 0 - 8 0 - cr- 7 0 - > 6 0 - 5 0 - E E =j 4 0 - o 3 0 - 2 0 - 1 0 - oL 0 50 40 30 20 10 0 0 -1. .. 1 1 1 i il LL i 1 m , I 2 3 4 5 6 7 Microns—> J 1 _ L _ 12 !3 14 3 4 6 7 8 9 10 II Microns —> Figure 41 Inclusion distribution for AC V) v . CD e E o v . >> Q_ to V . CD CJ 250 200' 150 100- 250- 200- 150- 100- 250 200' 150 T O l © © © e © © 100 DCSP • 9 © © © < X © e DCRP © o o or o © s e e AC N < X Q Q O 0 0 2 0 4-0 6 0 Distance From Weld Centerline (cm.) i i i. 00 1 u_ I I© "4 I 8 0 i CO Figure 42 Hardness traverses for large electrode welds T T 400 300 200 100 120 3. looi- o 50 0 120- 100 50 0 CL CL 0 F B . I F B . ~i r D C S P DCRP 20 4 0 6 0 Distance From Weld Centerline (cm.) F B . A C 8 0 i i — 1 DO Figure 43 Oxygen content traverses for large electrode welds - 183 - DCRP Experiment Number 10 DCRP Experiment Number II o 1600 ° c o >I700°C O 1̂600 °C O »I500°C o o 1300 °C O 1600 °C " CO >I700°C^/ -5 2 x Ac tua l S ize Figure 44 Slag temperatures for DCRP (experiment 10 and 11). Figure 45 Experiment 10, vertical section macrograph (corresponding above temperatures. - ± o t - DCSP Experiment Number 3 DCSP Experiment Number 8 1 j ' IS «: 1 o r 1650 °C 1650 °C O G 1750 °C 1700 °C O O >I750°C 1750 °C O O 1750 °C 1650 °C M GO 2 x Actua l S ize Figure 4 7 - 185 - AC Experiment Number 14 Figure. 49 © 1550 °C O 1550 °C I O 1650 °C O 1650 °C | j o I700°C o CD N CO O U < 2x Actual Size Figure 48 Slag temperatures f o r AC (experiment 14) I I P i i • • •  fimg| Experiment 14, v e r t i c a l s e ction macrograph (corresponding to c b o v c t e m p e r a : u r e s ) 14 F 13' 12- 11 • CM 11 * 10-O X 9 • O o 8 • c 7 - O i_ 6 • d> Q. e 5 • 4 - 3 - 2 - i u v a —s 1 I  r T I r i 1 r N° 1 2 3 5 7 cm •318 •636 •953 1-588 2-223 -I 1 —I L. -I 1 K. I I -1 1 I  i R 0 1 2 3 4 5 6 7 8 9 10 II 12 13 14 15 16 17 [8 19 2 0 T i m e in S e c o n d s (x 1 0 2 ) 00 a-- Figure 50 Thermal cycle for DCRP, experiment: 10. Figure 51 Thermal cycle for DCSP, experiment 8.same thermocouple positions as i n F i g . 50. Figure 52 . Thermal cycle for AC, experiment 14>same thermocouple p o s i t i o n s as i n F i g , 50. - 189 - 4 0 0 3 0 0 2 0 0 03 .O e 3 ro E o w >% 0. O 100 3 0 0 200 100 c o 3 O L L CD Root Gap C O DCRP N°I0 1042 Wire A36 Cylinder 304 Consumable Guide N < x -4- 0 4 0 0 300 200 c CJ 5 m DCSP N°3 1020 Wire A36 Cylinder Mild Steel Guide N < X CO [U. AC N°I4 1042 Wire J ^ 4340 x Cylinder 3 0 4 S S Guide 0 2 4 Distance From Weld (cm.) DCRP N°l l 1042 Wire 4340 Cylinder 304 Consumable Guide to DCSP N°8 1042 Wire A36 Cylinder I Mild Steel N Guide < x A ure 53 Example Vickers hardness traverses f o r CGESW c y l i n d e r s - 190 - c QJ X O 100 75 50 25 0 100 50 °~ 25 100 75 50 251- 0 J L DCRP Experiment P18 DCSP Experiment PI6 AC Experiment P5 I I 2 3 Distance From Weld Centerline (cm) 'Figure 54 Representative oxygen values f o r CGESW c y l i n d e r s - 191 - Figure 55 Macrograph of etched weld produced using a 304 S.S. consumable guide - ] 0 2 - Welding Velocity (cm./sec) „ „ 0 O005 0010 0015 0020 0025 U U o i 1 1 r 100 200 300 400 500 Slag Feeder RPM Figure 56 C a l i b r a t i o n curve for sla g feeder and method f o r determining s l a g a d d i t i o n r a t e o o o D v. <D CL E <u I - 14 13 12 II 10 9 8 7 6 5 4 3 --2. 1 r r T i r 1 T 1 r — i 1 1 r 1 r 0 8 9 10 II Time in Seconds (x 102) -i 1 1 «_ 12 13 14 15 _j !6 — i • • 17 18 19 20 Figure 57 Thermal p r o f i l e s f o r a l l thermocouples h inch from the o r i g i n a l surface f o r experiment P15. T i m e in S e c o n d s (x 1 0 2 ) Figure 58 Thermal p r o f i l e s f o r a l l thermocouples % inch from the o r i g i n a l surface f o r experiment P15. Figure 59 Thermal profiles for a l l thermocouples 3/4 inch from the original surface for experiment P15. 0 6 7 8 9 10 II 12 13 14 15 16 17 f8 19 20 Time in Seconds ( x lO 2 ) Thermal p r o f i l e s f o r a l l thermocouples 1 inch from the o r i g i n a l surface f or experiment P15.  141 I  1 1 1——i 1 1 r 1 1 1 1 1 1 1 1 r 13- I2r Time in Seconds (x !0 z ) Figure 62 Thermal p r o f i l e s f o r experiment P04. - 661 - Figure 64 Thermal profiles for experiment P08. Figure 65 Thermal profiles for experiment P10. Figure 66 Thermal profiles for experiment P l l , T i m e in S e c o n d s (x 1 0 2 ) Figure 67 Thermal profiles for experiment P13. Figure 68 Thermal profiles for experiment P14. - soz - Time in Seconds (x 102) Figure 70 Thermal profiles for experiment P18, - LUZ - I4i • i 1 i 1 1 1 1 1—-i 1 1 1——i }——i 1 1 r~—r 13' 12- . I w ' — I Q '• I I I I I I 1 1 1 1 1 l I ' • t i • • I 0 1 2 3 4 5 6 7 8 9 10 II 12 13 14 15 16 17 [8 19 20 Time in Seconds (xlO 2) Figure 72 Thermal profiles for experiment P22S, - uUZ - - 2 1 0 - Distance From Weld . Centerline (cm) Figure 74 T y p i c a l hardness traverses f o r various degrees of f u s i o n and modes of operation - 211 - cz OJ >» x O 125 100 75 50 25 i i 0 125 100 75 CL Q- 50 25 0 125 100 75 50 25 0 6= o CQ U. T T D C R P E x p e r i m e n t N ° I 0 X-—x E x p e r i m e n t N ° l I o O -O m J t D C S P E x p e r i m e n t N°3 A C E x p e r i m e n t N ° I 4 0 I 2 3 4 5 Distance From Weld Centerline (cm.) Figure 75 Oxygen values corresponding to the same traverses as Figure 74, - 212 - Figure 77 Inclusions i n DCSP ( =3000 X magnification) a g Figure 78 Micrographs of g r a i n s i z e v a r i a t i o n with distance from f u s i o n boundary In DCRP (P21); a, weld metal; b, adjacent to F.B.; c, 2mm; d, 4mmJ e, 6mm; f, 10mm; g, 15mm; h, base metal. 211X magnification H F i g u r e 79 Micrographs of g r a i n s i z e v a r i a t i o n w i t h d i s t a n c e from f u s i o n boundary i n DCSP (P22); a, weld m e t a l ; b, adjacent to F.B.; c, 0.5mm; d, 1.0mm; e, 3.0mm; f , 6.0mm; g, 10mm; h, 15mm; i , parent m e t a l . 21IX magnification b, 0. Figure 80 55mm into HAZ^20QX Fractograph of Charpy f r a c t u r e surface (P21) - 216 - Figure 81 Fractograph of Charpy fracture surface 3.0mm into HAZ (P21) Gt* 20 OX) Figure 82 Fractograph of Charpy fracture surface in weld zone showing du c t i l i t y and inclusions (P22) (*V1000X) - 2 1 7 - Figure 83 Fractograph of Charpy f a i l u r e surface (P22) (^lOOOX) - 218 - Figure 84 Fractograph of Charpy fra c t u r e surface i n spheroidized zone (P22) (^200X). a. weld o u t e r l i n e Figure 85 Fractograph of Charpy f r a c t u r e surface (P23) Q&20QX) - 219 - c. 1.5mm i n t o HAZ from F.B. Figure 85 Fractograph of Charpy f r a c t u r e surface (P23) (£*200X) - 220 - Figure 86 Fractograph of Charpy surface 1.5mm i n t o HAZ from F B C ^ I O O X ) O J T" Temperature in °C (x 102) ->j Oi CD i r~ CD - T " CO —1— o ro OJ H M ><! ro >i o PJ c 3 o 3 o 1—1 V n TJ ro o w f-i c 1—1 ro o CO H O f t U l M W O •f- O s i LO 3 O Ln O U i O 3 H ro t-i 3 o n o c XI - 1ZZ - c •-( n> OO •OO O O _ 3 co 5 " £ CD o o K 5 </> x — O ' ro o i 63 CD ro o ro —r- OJ Temperature in °C (x I02) - f i 0 1 C T ) - > I O D C D O = : ro —r~ OJ W M 2 C 3 CD H 3* CD i-( 3 o n o c to H M O rt . . . . H- O H O M CP O 4>- O s i U> O O Oi O Ui W X •d M a pd n M Pd g >-d Ed P3 H *d . 0 0 H to i-i 3 o o o o id 3 M • (0 CO CP - Z Z Z - TO C H ro oo H ro 3 o H> H-M <° H> O H -P- Temperature in °C (x 102) oo —i— OJ Oi CT) O ~_ — i r- ro —r— H ro H 3 o o o c fd X M i—i H M I—' I—1 o On \0 M ON O W O O Oi O Ot fj H a" ro i-i g o o o ^ c n x) 3 i—1 • ro . > O - £ZZ - OQ C H (D O Temperature in °C (x I02) -i> OI CD -si 00 CD O ro OJ OJ CD sj 55 col H rr ro i-i 3 o o o c xi M X N H H O j> o s i u O U i O L n H rr ro i-( 3 o o o O T3 3 M • ro T l s ] > L O C P 8' - vzz - 14 r 13' te- l l - CJ 10- g X 9 • O 0 8 - c 7 - o I . 6 - a. £ 5 - 4 • 3- 2 - 0 T T T" EXPERIMENT AC Thermocouple Number . ' ' 1 2 3 '• 4 V t I l I P P 8 A36 Thermocouple P o s i t i o n (cm.) 0.635 1.270 1.905 2.540 3 4 5 6 7 8 9 |0 II 12 13 Time in Seconds (x 102) U I 14 15 16 17 18 19 20 Figure 91 Thermal profile for PP8. 14 r 13- te- l l - 10- o X 9 • O 0 8 • c a? 7 • O i _ 6 • CD CL E 5 cu H- 4 3 2 0 T r — T ~ — T r EXPERIMENT DCRP Thermocouple Number 1 2 3 4 E2 T - l Thermocouple P o s i t i o n (cm.) 0.635 1.270 1.905 . 2.540 • • ' i 1 1 1 1 1 1 1 J 3 4 5 6 7 8 9 10 II 12 13 14 15 Time in Seconds (x 102) 16 17 18 19 20 Figure 92 Thermal profile for E2. 1,1. CTQ c (0 V O C O H rr ro H a P > t-* o F> M (0 t-h H O * 3' M C D O J * 5" CO C D o o Q. j C O I . . 1 X i . 1 1 o IM o ro OJ cn cn CO to o OJ cn 63 ro o Temperature in °C (x 102) OJ ~r~ — r — ~ i — CD i — CO 10 o "T— ro OJ J > - U ( O H ro 5 3 a ° cr o ro o H C 13 M X w a ^ o I—I pa S i d M H W CO L n v o t o H - J>- O s j u o O C n o C n 3 H nr ro i-t a o. o o c - AZZ - o x o o CO w . o cu CL e h- I 4 r 13- 1 2 II 1 0 9 8 7 6 5 4 0 ! T 1 £ i EXPERIMENT AC Thermocoupli Number 1 2 3 A E4 T - l Thermocouple P o s i t i o n (cm.) 0.635 1.270 1.905 .2.540 3 4 6 0 7 8 9 1 0 II . 1 2 13 1 4 15 Time in Seconds (xlO 2) 1 6 1 7 1 8 1 9 . 2 0 Figure 94 Thermal profile for E4. O' ' 1 ' I I, 8 U_ I II I I I Pi ' ' » ' • » 0 I 2 3 4 5 6 7 8 9 10 I! 12 13 14 15 16 17 f8 19 20 Time in Seconds (xlO 2 ) Figure 95 Comparison of thermal p r o f i l e s f o r some s p e c i f i c power inputs - 230 - 0> E 3 in <u c -o o I •o £ o k_ >> • Q . o 2 3 4 5 D i s t a n c e F r o m Weld C e n t e r l i n e (cm.) Figure 96 Hardness values f o r ESW u t i l i z i n g bar e l e c t r o d e s . (PP7 A36, PP1 A36, E4 T-l) 0 N3 L J • S P E C I F I C P O W E R ' k c a l . c m - | Figure 97 Penetration Versus s p e c i f i c power input, bar electrodes - 232 - Figure 98 Hardness values along the c e n t e r l i n e and near the surface of h o r i z o n t a l sections of T - l equivalent cast p l a t e s . - 233 - CZZ22 I cm I—I DCRP Experiment P2IS Electrical Distribution A MPS % . 216 38 17 3 30 115 20 24 4 • 47 8 575 Figure 99 E l e c t r i c a l d i s t r i b u t i o n i n DCRP. - 234 - DCSP Experiment P22S Electrical Distribution Amps % 177 30 177 30 1 47 25 34 10 29 5 590> Icm. Figure 100 E l e c t r i c a l d i s t r i b u t i o n i n DCSP. - 235 - AC Experiment P23 S Electrical Distribution Amps 127 •102 102 102 - 7 7 510 % "55" 20 20 20 15 I cm i 1 Figure 101 E l e c t r i c a l d i s t r i b u t i o n i n AC - 2 3 6 - Figure 103 Shunt set-up f o r P22S, DCSP Figure 104 Shunt set-up f o r P23S, AC  - 238 - Fig u r e 106 P e n e t r a t i o n shape w i t h bar e l e c t r o d e s i n the DCRP mode. - 239 - Figure 107 Macroetched longitudinal section of a weld produced during unstable welding conditions. - 240 - Figure 108 Macroetched l o n g i t u d i n a l section of a weld produced during s t a b l e welding condit i o n s . - 241 - Figure 109 T y p i c a l e l e c t r o s l a g weld surface. Figure 110 E l e c t r o s l a g surface f o r DCRP and AC. - 243 - Figure 111 Weld run-out. F i g u r e 112 H o r i z o n t a l a n d v e r t i c a l s e c t i o n s o f t w o t y p i c a l w e l d s , u s e d f o r o x y g e n a n a l y s i s , h a r d n e s s t r a v e r s e s , g r a i n s i z e m e a s u r e m e n t s a n d C h a r p y i m p a c t v a l u e s . - 245 - - 2 4 6 - 420 i 41-5 41 0 «5 40-5 4 0 0 39-5 39-0 3 8 5 1 —r Canron Welds Arcos PF201 Slag DCRP 400 500 600 Amperage - 700 800 Figure 114 Voltage-amperage c h a r a c t e r i s t i c s f o r Canron welds. 800 700 v 600 CP o I. <u £ < 500 400b Canron Welds Areas PF20I Slag DCRP 05 06 -07 08 Resistance (OHM) - •09 •10 Figure 115 Amperage versus apparent resistance f o r Canron weld - 2 4 7 - 800j 750] 7 0 0 6 5 0 | E 6001 a. < 550f- 5 0 0 4 5 0 h . 4 0 0 T 1 Conron Welds Arcos PF20I Slag DCRP 3 81 cm. S lag Depth • J - A - Wire Velocity (cm.sec" ) Figure 116 Amperage as a function of wire v e l o c i t y . •05 ' 1 1 1 1 : i _1 4 5 6 7 8 9 Wire Velocity (cm.sec" 1)—*• Figure 117 Apparent resistance as a function of wire v e l o c i t y for Canron welds. — LlU _ 4 o CD to n r — ~ 1 — Canron Welds Arcos PF20I Slag 3-81 cm. Slog Depth ! 31- 0> O rr CD 0 0 I 2 3 4 5 Energy Per Grri: of Metal Deposited (kcal.gm - 1) Figure 118 Energy per gram of metal deposit for Canron welds ^ ; — j — j ^ i:y,-̂ ,.v,- f 6 0 5 0 CD S 4 0 o > 3 0 2 0 A A A |& Q A I © Ref. 16 D Ref. 18 O Ref. 52 * Ref. 2 © Ref. 35 & Ref. 23 O Ref. 4 present research 0 2 0 0 4 0 0 6 0 0 8 0 0 Amperage 1000 1200 1400 Figure 119 L i t e r a t u r e c i t e d voltage-amperage c h a r a c t e r i s t i c s - 249 - Figure 120 L i t e r a t u r e c i t e d amperage-xrelding rate r e l a t i o n s h i p s . - 250 - Welding .Velocity (cm.sec." 1)—*• Figure 121 L i t e r a t u r e c i t e d apparent r e s i s t a n c e dependence on welding v e l o c i t y . - 251 - 8 A Ref. 16 x Ref. 52 DCRP O Ref. 52 DCSP ® Ref. 2 M Ref. 35 A Ref. 4 — Present Research o cu to E o> o o 4 | r r CD 0 0 I 2 3 4 Energy Per G m . of Metal Deposited (Kcal.gm."') Figure 122 L i t e r a t u r e c i t e d energy requirements. •II •10 m s •09 o cu •08 o c o w •07 to o or •06 0 5 •04 —1 1 — • D C R P O D C S P Immersion (cm.)—*• Figure 123 Electrode immersion versus apparent s l a g r e s i s t a n c e . E & 3 c .2 v> co 2 E E - I 0 E -ti 4 c o 2 3 tu E E 2!- 0 _L + j 1- l 1 -j ~ r DCRP Cylinders _1_ H 1 h H h DCSP Cylinders _ L JL 9 0 100 110 120 130 140 150 160 170 180 190 2 0 0 210 Kcal. cm." Figure 124 Electrode immersion versus energy input - 253 - Figure 125 Penetration versus energy input 0-5 i i r 1 * i r DGRP Cyl inders 0 ' ' 1 1 — — i i J I 0 I 2 3 4 5 6 7 Interelectrode Gap (cm.) Figure 126 Penetration versus s l a g depth. - 255 - i Figure 127 Temperature p r o f i l e across weld sections - 256 - c o CO u 2 E E 0 i 1 1 r T T • DCRP X DCSP O AC \ \ \ \ \ c?o0. J L • 0 4 - 0 5 06 07 08 09 10 II 12 13 R = V/i Figure 128 Immersion as a function of apparent resistance 8 0 0 7 0 0 0) CL E < 6 0 0 5 0 0 4 0 0 — r 1 C G E S W I 1/4" A36 Plates 1042 Wire X DCRP O AC O O <5> G O O •04 •06 08 Resistance (ohms) 1 0 Figure 129 Amperage as a function of apparent res i s t a n c e - 257 - IOcm/sec 0 4 8 12 16 20 24 28 .32 3 5 3 ° Radial distonce (cm) The computed v e l o c i t y f i e l d i n an i n d u s t r i a l s c a l e E S R system Figure 130 Th e o r e t i c a l s l a g flow pattern from Dilawari (75) 258 _ Figure 131 T y p i c a l bar electrode t i p s EXP NUMBER VOLTfiGE RMPERRGE SLRG DEPTH IMMERSION POOL DEPTH VELD VELOCITY . . THICKNESS S P E C I F I C POWER INPUT 5 42 490 3 . 1 8 0 . 9 3 1 . 6 0 3 . 1 8 3 5 3 VOLTS P.MPS CM. CM. CM. 0 . 0 1 4 C M . S E C - 1 CM. K C H L . C M . 00 —i 1 1 1 1 1 1 i i i r 0 4 0 8 12 16 2 0 24 28 32 36 4 0 4 4 T I M E IN S E C O N D S (X 10 ) Figure 132 Model predicted thermal p r o f i l e fo I Figure 133 Model predicted thermal p r o f i l e f o r P06, EXP NUMBER VOLTAGE flMPERRGE SLAG DEPTH IMMERSION POOL DEPTH VELD VELOCITY THICKNESS S P E C I F I C POWER INPUT 8 35 500 3 . 4 9 1 .27 1 .75 3 . 1 8 321 VOLTS HMPS CM. CM. CM. 0 . 0 1 3 C M . S E C - 1 CM. K C A L . C M . 0 0 0 4 0 8 12 16 2 0 24 28 T I M E I 32 IN 36 S E C O N D S 4 0 {X 10 4 4 Figure 134 Model predicted thermal p r o f i l e f o r P08. EXP NUMBER VOLTHGE AMPERAGE SLAG DEPTH IMMERSION POOL DEPTH . VELD VELOCITY TH1CKNE55 S P E C I F I C POVER INPUT 9 40 530 . 1.91 1.51 0 . 9 5 0 .021 3 . 1 8 234 VOLTS AMPS CM. CM. CM. C M . S E C - 1 CM. K C A L . C M . - 1 0 0 I 0 4 0 8 12 16 I 2 0 24 — 1 I 1 1— 2 8 32 36 4 0 T I M E IN S E C O N D S (X I0') 4 4 48 ro ro 52 ~I 56 — r 6 0 Figure 135 Model predicted thermal p r o f i l e f o r P09. EXP NUMBER VQLTRGE RMPERRGE SLRG DEPTH IMMERSION POOL DEPTH WELD VELOCITY . THICKNESS S P E C I F I C POWER INPUT 10 35 450 1.91 1 .10 0 . 9 5 3 . 1 6 166 VOLTS RMPS CM. CM. CM. 0 . 0 2 2 C M . S E C - 1 CM. K C R L . C M . 0 0 0 4 08 - i 1 1 1 : — i 1 1 r 12 16 2 0 24 2 8 32 36 4 0 T I M E IN S E C O N D S (X I0 1) 4 4 Figure 136 Model predicted thermal p r o f i l e for P10. EXP NUMBER 11 VOLTAGE 40 VOLTS RMPERBGE 560 RMPS SLRG DEPTH 1 .27 CM. IMMERSION 1.00 CM. POOL DEPTH 0 . 7 5 CM.. VELD V E L O C I T Y . 0 . 0 1 9 CM. S E C - THICKNESS 3 . 1 8 CM. S P E C I F I C POWER INPUT 277 K C R L . C M 0 8 *T~ 12 T 16 T 2 0 24 T 2 8 T 1 r 32 36 40 T I M E IN S E C O N D S (X 10 ) Figure 137 Model predicted thermal prof i l e for P l l . ru 1 * EXP NUMBER 12 VOLTRGE 44 VOLTS RMPERRGE . 510 AMPS SLRG DEPTH 3 . 4 9 CM. IMMERSION 0 . 9 3 . CM. POOL DEPTH 1.70 CM. VELD VELOCITY 0 . 0 2 0 C M . S E C - 1 THICKNESS 3 . 1 8 CM. S P E C I F I C POWER INPUT 267 K C f l L . C M . Figure 138 Model predicted thermal p r o f i l e f o r P12. Figure 139 Model predicted thermal prof i l e for P13. Figure 140 Model predicted thermal p r o f i l e for P14. Figure 141 Model predicted thermal prof i l e for PIS. Figure 142 Model predicted thermal prof i l e for P20, Figure 143 Predicted thermal profile for experiment 1(a)? Pugin (16), Figure 144 Predicted thermal profile for experiment 1(b), Pugin (16), Figure 145 Predicted thermal profile for experiment 2, Pugin (16)V Figure 146 Predicted thermal profile for experiment 3, Pugin (16)V  Figure 148 Predicted thermal profile for experiment 1, Sharapov ,(17), Figure 149 Predicted thermal profile for experiment fc, Sharapov ,^7)« Figure 150 Predicted thermal profile for experiment 3, Sharapov (17). Figure 151 Predicted thermal profile for experiment 1, Sharapov (18), Figure 152 Predicted thermal profile for experiment 1, Trepov <19;). 141 i i j-*—i i r—i 1 1 1 s — i i H I • » .— • R Time in Seconds (x 10*) 153 Predicted thermal profile for experiment. 2,'Trappy (19). 

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