ML20080F293
| ML20080F293 | |
| Person / Time | |
|---|---|
| Site: | Oconee  |
| Issue date: | 02/04/1984 |
| From: | Bassett B, Boyack B, Burkett M LOS ALAMOS NATIONAL LABORATORY |
| To: | NRC OFFICE OF NUCLEAR REGULATORY RESEARCH (RES) |
| References | |
| LA-UR-83-3182, NUDOCS 8402100418 | |
| Download: ML20080F293 (237) | |
Text
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JU LA-UR 3182 Y jd v
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.- l TITLE. TRAC ANALYSES OF SEVERE OVERC00 LING TRANSIENTS FOR THE OCONEE-1 PWR AUTHOR (St B. Bassett, B. Boyack, M. Burkett, J. Italand, J. Koenig, J. Lime, and R. Nelton Compiled by J. Ireland SUBMITTED TO- Nuclear Regulatory Commission Division of Accident Evaluation Nuclear Regulatory Research Washington, DC i
I Sy acceptance of tais article. tne ovoksner recognizes that tne U $ Governenent retains a noneocluseve. royalty-free hcenes to puosien or reproduce eneovoksned form of tn s contriouten, or to enous otners to de so ter U S Governenent purposes fne Los Alamos Peettonal Lacoratory requests tnet tne pupsnee 16entdy tnes arbcle as work performed under tne auspe.of tne U s Deoertment of Energy i
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CONTENTS ABSTRACT................................................................ 1 I. INTRODUCTION AND
SUMMARY
.......................................... 2 II. TRAC INPUT MODEL DESCRIPTION AND STEADY-STATE RESULTS............. 4 A. Primary Side......................................., ......... 6
- 1. Vesse1.................................................... 6
- 2. Hot Legs.................................................. 8
- 3. Steam Generators.......................................... 9
- 4. Cold Legs................................................. 11
- 5. Emergency Core-Cooling Systet (ECCS)...................... 11 B. Secondary Side................................................ 11
- 1. Feedwater Train........................................... 11
- 2. Steam-Generator Control Va1ves............................ 12
- 3. Emergency Feedwater....................................... 14
- 4. Steam Lines............................................... 14 C. Control Systen................................................ 15
- 1. Trips..................................................... 15
- 2. Integrated Control Systea................................. 20 D. Steady-State Calculation...................................... 30 III. TRAC TRANSIENT CALCULATIONS....................................... 35 A. Oconee-3 Turbine Trip......................................... 35
- 1. Introduction.............................................. 35
- 2. Model Description and Assumptions......................... 36
- 3. leansient Calculation..................................... 36
- 4. Summary................................................... 44 B. Main Steam-Line Break......................................... 44
- 1. Int r od uc tion and S ummary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 44
- 2. Model Description......................................... 47 -
- 3. Results................................................... 48
- a. Base Case (Case 1).................................... 48
- b. Parametric case - Case 2.............................. 70
- c. l'aramet ric cas e - Cas e 3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 79
, d. Par amet ric cas e - Cas e 4. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 90
- 4. C on c l us i ons . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10 3 C. PORY L0CA.................................................... 105
- 1. Int roduc tion and Summary. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10 5
- 2. Model Des cription and Assumptions . . . . . . . . . . . . . . . . . . . . . . . . 105
- 3. Transient Calculation.................................... 107
- 4. Summary.................................................. 112 D. Turbine Bypas s-Valve Failures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 119
- 1. One Bank of Two TBVs..................................... 119
- s. Int roduction and Summary. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 119
- b. Mod el Des crip t ion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 119 j
- c. Re s u l t s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 120
- i. B as e Cas e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 120
- 11. Paramet ri c Cas e 1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12 8 111. Par amet ric Cas e 2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 137 l
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- d. Concl us i ons . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13 8
- 2. Two Banks of Two TBVs.................................... 139
- s. Introduction and Summary....................... 2...m. 139 .
- b. Results.............................................. 140
- i. B as e C as e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14 0
- 11. Paramet ric Cas e 1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 148 111. Parametric Cas e 2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 156
- c. Conclusions.......................................... 157 E. No t-Les Bre ak LOCAs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 159
- 1. 2 in. Break.............................................. 159 3
- a. Int roduction and Summary. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 159
- b. Model Description and Assumptions. . . . . . . . . . . . . . . . . . . . 160
- c. Transient Calculations............................... 160
- d. Analysis of the Loop-Flow Oscillations............... 171
- 2. 4 in. Break.............................................. 174 l a. Int roduct ion and Summa ry. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 174
- b. Model Description and Assumptions..................... I'5 l
- c. Transient Calculation................................ 175
- d. S umma ry . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18 2 F. Rancho-Seco Type Transient (SC Dryout Followed by EFW Overfeed)................................................ 188
- 1. I nt r oduc t ion and S umma ry . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 186
. 2. Model Des cription and Assumptions . . . . . . . . . . . . . . . . . . . . . . . . 189
- 3. Results.................................................. 190
- 4. Conclusions.............................................. 197 IV. CONCLUSIONS AND RECOMMENDATIONS.................................. 197 REFERENCES............................................................. 201 1
APPENDIX A OCONEE ICS CONTROLLER FOR L00P-A........................... 202 APPENDIX B EXTRAPOLATIONS ............................................ 213 APPENDIX C UNCERTAINTIES IN OCONEE PTS CALCULATIONS................... 234 I
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-iii-LIST OF FIGURES .
- 1. Primary-side model for Oconee-1................................... 7
- 2. Vent-valve acde1.................................................. 8
- 3. Second ary-side mod el f or Oconee-1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10
- 4. Detail of f eedwat er-heat e r mod el (cross-section) . . . . . . . . . . . . . . . . . . 13
- 5. Steam generator control-valve schematie........................... 13 ..
- 6. Reactor trip system............................................... 17
- 7. HP1, RCP, and NFW realignment trip systes......................... 18
- 8. Hotwell, condensate-booster, and main feedwater pumps trip system....................................................... 19
- 9. SG isolation trip system.......................................... 20
- 10. Emergency feedwater trip systen................................... 21
- 11. ICS eganization................................................... 22
- 12. Feedwater-control section of the 1CS.............................. 23
- 13. BTU-limiter control b1ocks........................................ 24
- 14. Neutro n power cross-limit er cont rol blocks . . . . . . . . . . . . . . . . . . . . . . . . 25
- 15. Feedwater-flow control b1ocks..................................... 27
- 16. Level-limiter control b1ocks...................................... 29
- 17. Feedwater-valve adjustment control b1ocks . . . . . . . . . . . . . . . . . . . . . . . . . 29
- 18. MFW pump-speed control b1ocks..................................... 31
- 19. Main-steam safety-valve modeling for Osonee-3 transient........ .. 37
- 20. Measured main feedvater flow rates................................ 38
- 21. TRAC-calculated decay power and Oconee-3 measured th e rm al powe r . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 38
- 22. Calculated and seasured primary pressures......................... 39
- 23. Calculated and measured pressurizer water levels.................. 39
- 24. Calculated and measured hot- and cold-leg temperatures for loop A..................................................... .. 42
- 25. Calculated and measured hot- and cold-leg temperatures for loop B........................................................ 42
- 26. Calculated and measured steam generator A secondary pressures..... 43
'27. Calculated and seasured steaa-generator B secondary pressures..... 43
- 28. Calculated and measured water levels in steam generator A......... 45
- 29. Calculated and seasured water levels in steam generator B......... 45
- 30. Pressurizer pressure (0-900 s) - base case........................ 51
- 31. Pres surizer pres sure (0-7200 s) - bas e cas e. . . . . . . . . . . . . . . . . . . . . . . 51 .
- 32. Pressurizer water level (0-900 s) - base case..................... 52
- 33. Pressurizer water level (0-7200 s) - base case.................... 52
- 34. Downconer liquid temperatures (0-900 s) at vessel axial level 6 (all azimuthal sectors) - base case............................. 53
- 35. Downconer liquid temperatures (0-7200 s) at vessel axial level 6 (all azimuthal sectors) - base case............................. 53
- 36. Total vent-valve flow into downconer - base case.................. 55
- 37. Hot-leg liquid subcooling (0-900 s) - base case................... 55
- 38. Hot-leg liquid subcooling (0-7200 s) - bas e casa. . . . . . . . . . . . . . . . . . 56
- 39. Ho t-le g mas s flows (0-900 s ) - bas e cas e. . . . . . . . . . . . . . . . . . . . . . . . . . 56
- 40. Ho t-leg mas s flows (0-7200 m) - base cas e. . . . . . . . . . . . . . . . . . . . . . . . . 57
- 41. Candy-cane void fractions - base case............................. 57
- 42. Upper plenum liquid volume f raction - base case. . . . . . . . . . . . . . . . . . . 58
- 43. Loo p-A cold-leg mas s flows - bas e case . . . . . . . . . . . . . . . . . . . . . . . . . . . . 58
- 44. Loop-A cold-leg liquid temperaturae - base case.................. 59
- 45. Loop-B cold-leg mass flows - base cas e. . . . . . . . . . . . . . . . . . . . . . . . . . . . 59 .
- 46. Loop-B cold-leg liquid temperatures - bas e case. . . . . . . . . . . . . . . . . . . 61
- 47. SG A secondary-side water inventory - base case................... 61
- 48. SG A secondary-side pressure - base case.......................... 62
-iv-
- 49. SG A steam-line flow - base casa.................................. 62 '
- 50. CCFL phenomena in affected steam generator (SG A) - base case..... 63
- 51. SG B secondary-side water inventory (0-900 s) - base case......... 63
- 52. SG B secondary-side water inventory (0-7200 s) - base case........ 64
- 53. SG B secondary-side pressure (0-900 m) - base case................ 64
- 54. SG B secondary-side pressure (0-7200 s) - base case... . .. ..... . .. . 66
- 55. MFW pump speed - base case........................................ 66
- 56. MF W liquid temperature - bas e cas a . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 67 ..
- 57. MFW m as s f lows - bas e ca s e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 67
- 58. EFW mas a flows (0-900 s) - bas e cas e. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 63
- 59. EFW mas s flows (0-7200 s) - bas e cas e. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 68
- 60. EFW liquid temperature at pump discharge - base case.. . .. . .... . .. . 69 61, EFW liquid temperatures at injection locations - base case........ 69
- 62. Loo p-A HPI flows - ba s e cas e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 71
- 63. Loo p-B HPI flows - bas e cas e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 71
- 64. Accusulator water levels - base case...................~........... 72
- 65. Accumulator liquid volume discharged - base casa.................. 72
- 66. PO RV mas s f l o w - bas e c a s e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 73 67.- Pressurizer pressure - Case 2..................................... 73
- 68. Pressurizer water level - Case 2.................................. 75
- 69. Downcomer liquid temperatures at vessel axial level 6 (all
, a zimuthal sec tors) - Cas e 2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 75
- 70. Bo t-leg mas s f lows - Cas e 2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 76
- 71. Candy-cane void fractions - Case 2................................ 76
- 72. Loo p-A cold-leg mas s f lows - Cas e 2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 77
- 73. Loo p-B cold-leg mas s flows - Cas e 2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 77
- 74. Loo p-A cold-leg liquid temperatures - Cas e 2. . . . . . . . . . . . . . . . . . . . . . 78
- 75. Loo p-B cold-leg liquid temperatures - Cas e 2. . . . . . . . . . . . . . . . . . . . . . 78
- 76. SG A second ary-side water inventory - Case 2. . . . . . . . . . . . . . . . . . . . . . 80
- 77. SG A secondary-side pressure - Case 2............................. 8C
- 78. SG A steam-line flow - Case 2..................................... 81
- 79. SG B secondary-side water inventory - Case 2. . . . . . . . . . . . . . . . . . . . . . 81
- 80. SG B secondary-side pressure - Case 2............................. 82
- 81. MFW pump speed - Case 2........................................... 82
- 82. HP I f l ows - C as e 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 83
- 83. Pres surizer pres sure - Cas e 3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 83
- 84. Pressurizer water level - Case 3.................................. 85
- 85. Downconer liquid temperatures at vessel axial level 6 (azimuthal -
sectors) - Case 3................................................. 85
- 86. Ho t-le g mas s f lo ws - Cas e 3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 86
- 87. Candy-cane void fractions - Case 3................................ 86
- 88. Loo p-A cold-leg mas s flows - Cas e 3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 87
- 89. Loo p-B cold-leg mas s flows - Cas e 3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 87 CO. Loop-A cold-leg liquid temperatures - Cas e 3. . . . . . . . . . . . . . . . . . . . . . 38
- 91. Loop-B cold-leg liquid temperatures - Cas e 3. . . . . . . . . . . . . . . . . . . . . . 88
- 92. SG A secondary-side water inventory - Case 3...................... 89
- 93. SG A secondary-side pressure - Cas e 3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 89
- 94. SG A s t e am-line flow - Cas e 3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 91
- 95. SG B secondary-side water inventory - Case 3...................... 91
- 96. SG B secondary-side pressure - Case 3............................. 92
- 97. MFW pump s pe ed - Cas e 3 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 92
- 98. System pressure for Case 4 and base case. . . . . . . . . . . . . . . . . . . . . . . . . . 93
- 99. Downcomer liquid temperatures at vessel axial level 6 for Case 4 and base case.............................................. 95 100. Break flow for Case 4............................................. 96 101. SG-A second y pressure for Case 4................................ 96 102. SG-B secondary pressure for Case 4................................ 97
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-v-103. EFW flow (SG A and SG B) for Case 4............................... 97 104. SG A total feedwater flow for Case 4.............................. 98 ,
105. SC B total feedwater flow for Ceee 4.............................. 98 106. SG-A tube-bundle-region mas s inventory for Cas e 4. . . . . . . . . . . . . . . . . 99 107. SG-B tube-bundle-region mas s invent ory for Cas e 4. . . . . . . . . . . . . . . . . 99 108. Ho t-leg A and B mas s flows f or Cas e 4. . . . . . . . . . . . . . . . . . . . . . . . . . . . 100 109. Hot-leg A and B liquid t emperatures f or Cas e 4. . . . . . . . . . . . . . . . . . . 100 110. Loop A cold-leg mass flows f or Cas e 4. . . . . . . . . . . . . . . . . . . . . . . . . . . . 101 111. Loop A cold-leg liquid temperatures for . Case 4. . . . . . . . . . . .. . . . . . . 101 - -
112. Loop B cold-leg mas s flows f or Cas e 4. . . . . . . . . . . . . . . . . . . . . . . . . . . . 102 113. Loop B cold-leg liquid temperatures f or Case 4. . . . . . . . . . . . . . . . . . . 102 114. To t al HPI flow for Cas e 4. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10 4 115. Total mas s flow through vent valves f or Case 4. . . . . . . . . . . . . . . . . . . 104 116. Upper plenum liquid temperature f or Cas e 4. . . . . . . . . . . . . . . . . . . . . . . 105 117. SG A secondary pressure.......................................... 108 118. SG B secondary pressure.......................................... 108 119. Main-f eedwat er flow - loops A and B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 109 120. Main-feedvster liquid temperature - loops A and B................ 109 121. Realignment mass flow - loop A (negative flow is into SG)........ 110 122. Realignment mass flow - loop B (negative flow is into SG)........ 110 123. St eam generator s econdary inventory - loop A. . . . . . . . . . . . . . . . . . . . . 111 124. Steam generator secondary inventory - loop B..................... 111 125. Pressurizer pressure............................................. 111 126. Pressurizer water leve1.......................................... 113 127. PORV sass f1ow................................................... 114 128. PORV vapor fraction.............................................. 114 119. Ho t-leg mas s flows - loo ps A and B. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 115
'30. Hot-leg liquid temperatures - loops A and B. . . . . . . . . . . . . . . . . . . . . . 115 131. Cold-leg mass flows - loops Al and A2............................ 116 132. Cold-leg mas s flows - loops B1 and B2. . . . . . . . . . . . . . . . . . . . . . . . . . . . 116
- 33. Cold-leg liquid temperatures - loops Al and A2. . . . . . . . . . . . . . . . . . . 117 134. Cold-leg liquid temperatures - loops B1 and B2................... 117 135. Candy-cane void f rac tions - loops A and B. . . . . . . . . . . . . . . . . . . . . . . . 118 136. Vessel upper plenum void fractions - all azimuthal cells......... 118 137. Downcomer liquid temperatures at vessel axial level 6 (all a zimuth al se c t o rs ) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 119 138. Pressurizer pressure............................................. 121 139. Steam generator-secondary pressure - loop A...................... 121 140. Steam generator-secondary pressure - loop B...................... 122 141. Main-feedwater flow - loop A..................................... 122 142. Main-f eedwat e r flow - loop B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 124 133. Main-feedwater liquid temperatures - loop A...................... 124 144. Main-feedwat er liquid temperatures - loop B. . . . . . . . . . . . . . . . . . . . . . 125 ItS. Flow through emergency-feedwater header - loop A. . . . . . . . . . . . . . . . . 125 146. Flow through emergency-feedwat er header - loop B. . . . . . . . . . . . . . . . . 126 147. Liquid temperatures in the emergency-feedwater header
-loopA.........................................................126 148. Liquid temperatures in the emergency-feedwater header
- loop B......................................................... 127 149. Steam generator-secondary inventory - loop A..................... 127 150. Steam-generator-secondary inventory - loop B..................... 129 151. Het-leg flow - loop A............................................ 129 152. L;t-leg flow - loop B............................................ 130 153. Cold-leg flow - loop A1.......................................... 130 154. Cold-leg flow - loop A2.......................................... 131 155. Cold-leg flow - loop B1.......................................... 131 ,
156. Cold-leg flow - loo p B2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13 2 i I
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157. Cold-leg liquid t emperatures - loop B1. . . . . . . . . . . . .'. . . . . . . . . . . . . . 132 158. Cold-leg liquid t emperatures - loop B2. . . . . . . . . . . . . . . . . . . . . . . . . . . 133 -
159. Cold-leg liquid t emperatures - loop A1. . . . . . . . . . . . . . . . . . . . . . . . . . . 133 160. Cold-leg liquid temperatures - loop A2. . . . . . . . . . . . . . . . . .. . . . . . . . . 134 161. Candy-cane v apor f raction - loop A. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 134 162. Candy-cane vapor f rac tion - loop B. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13 5 163. Pre s suri zer wat er leve1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 135 164. Downconer liquid temperatures (base case) at vessel axial level
6 (all azimuthal sectors)........................................ 136 165. Downcomer liquid temperatures (p'araaztric case 1) at vessel axial level 6 (all azimuthal sectors)....................e............. 136 166. - Downcomer liquid temperatures (parametric case 2) at vessel axial level 6 (all azimuthal sectors).................................. 139 167. P re s s u ri z e r pre s s u re . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 141 168. Steam generato r-secondary pres sure - loop A . . . . . . . . . . . . . . . . . . . . . . 141 169. Steam generator-second:ry pressure - loop 1...................... 142 170. Main-f eedwat er flow - loo p A. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14 2 171. Main-feedvater flow - loop B..................................... 144 172. Main-feedsater liquid temperatures - loop A....................... 144 173. Main-feedwater liquid temperatures - loop B...................... 145 174. Emergency-feedwater flow - loop A................................. 145 175. Emergency-feedwater flow - loop B................................ 146 176. Emergency-feedwater liquid temperatures - loop A................. 146 177. Emergency-feedwater liquid temperatures - loop B................. 147 178. S te am generat or-secondary invent ory - loop A. . . . . . . . . . . . . . . . . . . . . 147 179. Steam generator-secondary inventory - loop B..................... 149 180. Ho t -l e g flo w - lo o p A . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14 9 181. Ho t-le g f lo w - lo o p B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 0 182. Col d-le g flo w - loo p A1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 150 183. Col d-leg flow - lo o p A2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 151
, 184. Co ld-l e g flow - loo p B 1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 151 185. Cold-leg flow - loop B2.......................................... 152
.186. Cold-leg liquid temperature - loop A1. . . . . . . . . . . . . . . . . . . . . . . . . . . . 152 187. Cold-leg liquid temperature - loop A2. . . . . . . . . . . . . . . . . . . . . . . . . . . . 153 188. Cold-leg liquid temperature - loop B1. . . . . . . . . . . . . . . . . . . . . . . . . . . . 153 189. Cold-leg liquid temperature - loop B2. . . . . . . . . . . . . . . . . . . . . . . . . . . . 154 190. Candy-cane vapor fraction - loop A............................... 154 191. Candy-cane vapor fraction - loop B............................... 155 192. Pres surize r wat er leve1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 5 193. Downcomer liquid temperatures (base case) at vessel axial level 6 (all azimuthal sectors)........................................ 158 194. Downcomer liquid temperatures (parametric case 1) - vessel axial level 6 (all a zimuthal sectors) . . . . . . . . . . . . . . . . . . . . . . . . . . . . 158 195. Downcomer liquid temperatures (parametric case 2) - vessel axial level 6 (all azimuthal sectors)............................ 159 196. Steam generator secondary-side pressure - loop A................. 161 197. Steam generator secondary-side pressure - loop B................. 161 198. Steam generator secondary-side inventory - loop A................ 162
'99. Steam generator secondary-side inventory - loop B................ 162 200. MFW mas s flows - loops A and B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 164 201. MFW liquid temperatures - loops A and B . . . . . . . . . . . . . . . . . . . . . . . . . . 164 202. Eme rgency/ realigned mas s flows - loops A and B. . . . . . . . . . . . . . . . . . . 165 103. Emergency / realigned liquid temperatures - loops A and B.......... 165 204. P r e s s u ri zer p r es s ure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 166 205. Pressurizer water leve1.......................................... 166 206. Break mass f1ov.................................................. 167 207. B re ak void f ra c tion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 167 1
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-vii- I 208. Candy-cane void fractions - loops A and B........................ 168 209. Cold-leg mas s flows - loops Al and A2. . . . . . . . . . . . . . . . . . . . . . . . . . . . 168 210. Cold-leg mass flows - loops El and 52............................ 169 211. Cold-leg liquid temperatures - loops Al and A2. .. .. . . . . . .. ..... .. 169 212. Cold-leg liquid t emperatures - loops B1 and 52. . . . . . . . . . . . . . . . . . . 170 213. Ho t-leg mas s flows - loo ps A and B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 17 0 214. Hot-leg liquid t em peratures - loops A and B. . . . . . . . . . . . . . . . . . . . . . 172 215. Downconer liquid temperatures - vessel axial level 6 (all a zimuthal sec t ors ) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 172 216. Downcomer liquid temperature comparison for 2-in. break cas e (vent valves versus no vent va1ves ) . . . . . . . . . . . . . . . . . . . . . . . . . 173 217. To t al vent v Alve mas s f1ow. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 173 218. Steam generator secondary-side pressure - loop A................ 176 219. Steam generator secondary-side pressure - loop B................. 176 220. Ste am generator secondary-side inventory - loop A.. .. .. . . . .. . .. . . 177 221. Steam generator secondary-side inventory - loop B................ 177 222. MFW flows - lo o ps A and B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 178 223. MFW liquid temperatures - loops A and B. . . . . . . . . . . . . . . . . . . . . . . . . . 178 224. Realig'nsd mass flows - loo ps A and B. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 179 225. Realigned liquid temperatures - loops A and B. . . . . . . . . . . . . . . . . . . . 179 226. Pre s s u ri z e r pre s s u re . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 180 227. Pre s s u ri ze r wat e r leve1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 180 218. Break mass f1ow.................................................. 181 229. B r e ak void f ra c t ion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 181 230. Candy-cane void fractions - loops A and B........................ 183 231. Ho t-le g mas s flows - loo ps A and B. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 183 232. Hot-le g liquid temperatures - loops A and B. . . . . . . . . . . . . . . . . . . . . . 185 233. Cold-leg mass flows - loops Al and A2. . . . . . . . . . . . . . . . . . . . . . . . . . . . 185 234. Cold-leg mas s flows - loops B1 and B2. . . . . . . . . . . . . . . . . . . . . . . . . . . . 186 235. Cold-leg liquid temperatures - loops Al and A2. . . . . . . . . . . . . . . . . . . 186 236. Cold-leg liquid temperatures - loops B1 and 52. . . . . . . . . . . . . . . . . . . 187 237. Tot al po sit ive vent-valve vapor mass f1ow. . . . . . . . . . . . . . . . . . . . . . . . 187 238. Downcomer liquid temperatures - vessel axial level 6 (all azimurhaJ sectors).......................................... 188 239. Prim ar y sy s t em pre s s ure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 191 240. Pressutizer water leve1.......................................... 191 241. PORV mass f1ow.................................................... 192 242. SG-A secondary-side pressure..................................... 192 .
243. SG-B secondary-side pressure..................................... 193 244. E FW m as s f 1 o ws . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19 3 245. EFW liquid temperatures at injection point....................... 194 246. SG-A secondary-side inventory.................................... 194 247. SG-B secondary-side inventory.................................... 195 248. Loop A cold-leg liquid temperatures.............................. 195 249. Loop B cold-leg liqcid temperatures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 196 250. Ho t-leg liquid t em peratures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 196 251. Loo p A HPI m as s f1ows . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 198 252. Loo p B H P I a sa s f 1 o ws . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 198 253. Loo p A cold-J eg mas s f1ows . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 199 254. Loo p B c old-leg mas s f1ows . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 199 255. Ho t-l e g m as s f 1ows . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 200 256. Downcomer liquid temperatures-vessel axial level 6 (all azimuthal sectors).......................................... 200 B-1. Pressurizer pressure............................................. 214 B-2. Downconer liquid temperatures at vessel axial level 6 (all azimuthal sectors).......................................... 215 e
w - e m- - --s. , , s- p-.-- --g---g.- - ,aa--,er-.ggq-:-w
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-viii-B-3. Heat-transfer coefficients at vessel axial level 6 '
(all azimuthal sectors).......................................... 216 B-4. Pressurizer pressure histories for Case 5 (Case 5A-base; case 5B parametric 1; case 5C parametric 2)..................'.... 217 B-5. Pressuriser pressure histories for Case 6 (Case 6A-base; case 6B parametric 1; Case 6C parametric 2) .. . . . . . . . . . . . . . . . . . . . . 217 3-6. Downcomer liquid temperatures at vessel axial level 6 l (all azimuthal sectors) for Case 5A.............................. 218 B-7. Downcomer liquid temperatures at vessel axial level 6 (all a zimuthal sectors ) for Cas e 55. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 218 B-8. Downcomer liquid temperatures at vessel axial level 6 (all a zimuthal sectors) for Case 5C. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 219 B-9. Downeomer liquid temperatures at vessel axial level 6 (all azimuthal sec tors) for Cas e 6A. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 219 B-10. Downcomer liquid temperatures at vessel axial level 6 (all a zimuthal sec tors) for cas e 6B. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 220 B-ll. Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors) for Case 6C.............................. 220 B-12. Heat-transfer coefficients at vessel axial level 6 (all azimuthal sectors) for Case 5A.............................. 221 B-13. Heat-transfer coefficients at vessel axial level 6 (all a zimutbal sectors ) for cas e 5B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 221 B-14. Heat-transfer coefficients at vessel axial level 6 l (all azimuthal sectors) for Case 5C. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 222 l B-15. Heat-transfer coefficients at vessel axial level 6 (all azimuthal sectors) for Case 6A.............................. 222 B-16. Heat-transfer coefficients at vessel axial level 6 (all azimuthal sectors) for Case 6B.............................. 223 B-17. Heat-transfer coefficients at vessel axial level 6 (all azimuthal sectors) for Case 6C.............................. 223 B-18. PORV LOCA extrapolat ed sys tem pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . 226 B-19. PORV LOCA extrapolated downconer liquid temperature.. . . ...... ... 226 B-20. PORV LOCA extrapolated downcomer heat-transfer coefficient....... 227 B-21. 4-in-diam. SBLOCA extrapolated sys tem pres sure . . . . . . . . . . . . . . . . . . . 227 B-22. 4-in-diam. SBLOCA extrapolated downconer liquid t em pe r a tu r e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22 8 B-23. 4-in-diam. SBLOCA extrapolated downconer heat-transfer coefficient...................................................... 228 -
B-24. 4-in-diam. SBLOCA extrapolated system pressure. . . . . . . . . . . . . . . . . . . 231 B-25. 4-in-dias. SBLOCA extrapolated downcomer liquid
- t em pe r a t u r e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 231 B-26. 4-in-dias. SBLOCA extrapolated downconer heat-transfer Coefficient...................................................... 232 B-27. Rancho-Seco type transient extrapolated system pressure.......... 232 ,
B-28. Rancho-Sece type transient extrapolated downconer '
liquid t am pe ra t ure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 233 B-29. Rancho-Seco type transient extraplated downcomer heat- l trans f er coe f ficient s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 233 l
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-ix-LIST OF TABLES I. LOS ALAMOS OCONEE-1 PTS OVERC00 LING TEANSIENT CALCULATIONS...... 3 II. TRAC OCONEE-1 TRANSIENT RESULTS................................. 5 III. SIMPLE TRIPS.................................................... 16 IV. BTU-LIMITER C0hTROL-BLOCK EQUATIONS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 24 V. NEUTRON-F0WER CROSS-LIMITER CONTROL-BLOCK EQUATIONS............. 26 VI. FEEDWATER-FLOW CONTROL-BLOCK EQUATIONS................c......... 27 VII. LEVEL-LIMITER CONTROL-BLOCK EQUATIONS........................... 28 VIII . FEEDWATER-VALVE ADJUSTMENT C0hTROL-BLOCK EQUATIONS . . . . . . . . . . . . . . 30 IX. MFW-PUMP-SPEED CONTROL-BLOCK EQUATIONS.......................... 32 X. PRIMARY-SIDE STEADY-STATE CONDITIONS............................ 33 XI. SECONDARY-SIDE STEADY-STATE CONDITIONS.......................... 34 XII. INITIAL STEADY-STATE CONDITIONS................................. 40 XIII. MAIN-STEAM-SAFETY AND TURBINE-BYPASS VALVE SETPOINTS............ 40 XIV. SEQUENCE OF EVENTS.............................................. 41 XV. COMPARISON OF TRAC AND OCONEE-3 RESULTS......................... 46 XVI. SEQUENCE OF EVENTS.............................................. 49 i
XVII. MSLB (CASE 2) SEQUENCE OF EVENTS................................ 74 XVIII. MSLB (CASE 4) SEQUENCE OF EVENTS............................... 106
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IIX. MSLB (CASE 3) SEQUENCE OF EVENTS...............#................ 106 XX. PORV LOCA EVENT SEQUENCE....................................... 107 XXI. TBV EVENT SEQUENCE (BASE CASE)................................. 123 XXII. TBV EVENT SEQUENCE PARAMETRIC CASE 1........................... 137 XXIII. TBV EVENT SEQUENCE FARAMETRIC CASE 2........................... 138 XXIV. TBV EVENT SEQUENCE BASE CASE................................... 143 XXV. TBV EVENT SEQUENCE PARAMETRIC CASE 1........................... 148 XXVI. TBV EVENT SEQUENCE PARAMETRIC CASE 2........................... 157 XXVII. HOT-LEG BREAK LOCA - 2 IN. BREAK SEQUENCE OF EVEPTS............ 160 XXVIII. HOT-LEG BREAK LOCA - 4 IN. BREAK EVENT SEQUENCE AASE CASE...... 182 XXIX. RANCFO-SECO TYPE TRANSIENT INITIAL CONDITIONS AND POSTULATED EVENT SEQUENCE................................................. 189 XXX. RANCHO-SECO TYPE TRANSIENT SEQUENCE OF EVENTS.................. 190 B-I. EXTRAPOLATED RESULTS FOR TBV TRANSIENTS AT 7200 s.............. 225 C-I . SYSTEMS AFFECTED BY UNCERTAINTIES . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 236 I
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ANS American Nuclear Society BTU British Thermal Unit R&W Babcock & Wilcox ECCS Emergency Core-Cooling System .-
EFW Emergency Faedwater EFWV Emergency Feedwater Valve FSAR Final Safety Analysis Report HPI High-Pressure Injection ICS Integrated Control System IDCA Loss-of-Coolant Accident LPI Low-Pressure Injection LPIS Low-Pressure Injection System MFCV Main-Flow-Control Valve MFW Main Feedwater
( MSLB Main Steam-Line Break MSSV Main Steam Safety Valve NDT Nil-Ductility Temperature NRC Nuclear Regulatory Commission ORNL Oak Ridge National Laboratory PORV Power-Operated Relief Valve PTS Pressurized Thermal Shock PWR Pressurized Water Reactor
, RCP Reactor Coolant Pump 1
l SG Steam Cenerator ,
SUFCV Start-Up Flow-Control Valve l TBV Turbine-Bypass Valve TRAC Transient Reactor Analysis Code TSV Turbine-Stop Valve USI Unresolved Safety Issue 9
I mm
-xi-ACKNOWLEDGDENTS T1.= authors wish to acknowledge the extraordinary, efforts of word processors Jean Martines and Cecilia Gonzales in the organization and processing of this document. Also, the efforts of Sylvia Lee in preparing the graphics for the calculations are greatly appreciated.
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4 TRAC ANALYSES OF SEVERE OVERC00 LING TRANSIENTS Foil THE OCONEE-1 PWR*
by B. Bassett, B. Boyack, M. Burkett J. Ireland, J. Koenig, J. Line, and R. Nelton Compiled by J. Ireland ABSTEACT This report describes the results of several TRAC-PF1 calculations of overcooling transients in a Babcock & Wilcox lowered-loop pressurized-water reactor (Oconee-1). The purpose of this study is to provide detailed thermal-hydraulic input to Oak Ridge National Laboratory for pressurized thermal-shock analyses. The trant hot calculations performed were plant specific in that details of the primary systen, the secondary system, and the plant integrated control system of Oconee-1 were included in the TRAC input model. The results of the-calculations indicate that the turbine-bypass valve failure transient was the most severe in terms of resulting in relatively cold liquid temperatures in the downconer region of the vessel. The power-operated relief-valve LOCA transient was the least severe in terms of downconer liquid temperatures because of vent-valve firid mixing and near saturated conditions in the primary systee. .It is recommended that future calculations consider a wider range of operator actions to cover the spectra of overcooling traasient sequences more completely.
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- Work performed under the auspices of the United States Nuclear Regulatory
- Commission.
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I. INTRODUCTION AND
SUMMARY
Pressurized thermal shock (PTS) in pressurized water reactors (PWRs) has ,
been identified by the Nuclear' Regulatory Commission (NRC) as an unresolved safety issue (USI A-49). Because of this, the NRC has a major prc; ram distributed among several organizations to help resolve the PTS issue. The goal of this project is to determine the potential risk of older reactor vessels to severe overcooling transients that rapidly cool the primary system.
The Los Alamos contribui: ion to this project is to use the multi-dimensional two'-fluid, non-equilibrium numerical simulation code, TRAC-PFil, to provide accurate thermal-hydraulic conditions during postulated PTS accidents in selected PWRs. This report presents the results of several TRAC-PF1 thermal-hydraulic calculations performed for the Oconee-1 PWR. The Oconee-1 PWR is operated by Duke Power Company, and the nuclear steam-supply system was designed by the Babcock & Wilcox (B&W) Company. The main purpose of these calculations was to determine which of the overcooling transients specified by Oak Ridge National Laboratory (ORNL) was the most severe in terms of cold liquid temperatures in the downcomer region of the reactor vessel. These ORNL-specified transients are listed in Table I.
The concern over PTS arises because the material properties of the vessel wall change after ;everal years of irradiation.2 The vessel wall becomes embrittled and its nil-ductility temperature (NDT) increases. If during an accident, overcooling of the primary-system liquid cools the vessel wall below the NDT (the NDT Foc Oconee-1 is ~365 K) and the system subsequently repressurizes, the possibility exists that defects could be initiated or propagated in the vessel wall. Such overcooling of the primary-system liquid may result from the high pressure injection (HPI) system or rapid cooling by the secondary system. -
Because the risk of initiating or propagating flaws in the vessel wall depends on t'ie coupling of the thermal stresses produced by overcooling with the mechanical stresses from repressurization, detailed system models are required.
! Modeling both the primary and secondary systems of the reactor plant is l necessary to properly analyze the PTS phenomena. The steam generator (SG) secondary-side inlet conditions directly affect primary temperature, pressure, and the emergency core-coolant injection. Secondary-side inlet conditions are highly dependent on main-feed pump and SG control-valve operations as well as the termination of the extracted steam supply to the feedwater heaters. Other impo. tant systems modeled in the TRAC inp,ut deck include a model of the B&W Integrated Control System (ICS) used at the Oconee-1 plar:t. The ICS sonitors 5
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TABLE I LOS ALAMOS OCONEE-1 PTS OVERC00 LING TRANSIENT CALCULATIONS Transient Description
- 1. Oconee-3 turbine trip Simulate actual plant transient
that occurred on March 14, 1980.
- 2. Main-steam line break 34" steam-line break; all systems l . operate as designed; steam generators l isolated at 10 min., unaffected steam generator refilled at 15 min.
stuck open) ICS fails to run back main feedwater; primary coolant pump trip.
- 4. Turbine-bypass valve One bank of TBVs f ail to failure (one bank of two valves) reseat af ter opening.
- a. SG 1evel control fails
- b. SG 1evel control does not fail
- c. RCP restart; HP1 throttled I
- 5. Turbine-bypass valve Two banks of TBVs fail failure (two banks of two valves) to reseat after opening.
- a. SG 1evel control fails
- b. SG 1evel control does not fail
- 6. Small-break LOCA Two-inch diameter hole in (2 in. hot-leg break) pressurizer surge li'ne; RCP trip; all systems operate as designed.
- 7. Small-break LOCA Four-inch diameter hole in (4 in. hot-leg break) pressurizer surge line; RCP trip; all systems operate ~
as designed.
- 8. Rancho-Seco type Initial loss of feedwater transient followed by run-away-emergency feedwater to both steam generators-i the primary flows ad temperatures to determine the feedwater demand. It also regulates the main- and startup flow-control valves, the main-feedwater (MFW) pumps, and the turbine-bypass valves (TEVs). Details of the ICS are presented in Appendix t..
Several overcooling transients have been identified by ORNL,3 and ad'ditional transients may be specified after these initial results are l
evaluated. The initial transients include a main steam-line break (MSLB) with a delay in isolating the affected steam generator, a small-break LOCA [ full-open I
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i l failure of the power-operated relief valve (PORV)] with failure of the ICS to '
i throttle main-feedwater flow and trip the reactor coolant, pumps (RCPs), and turbine bypass-valve transients with steam generator overfeed. An actual plant l transient (Oconee-3 turbine trip) was also simulated by TRAC to compare with actual plant data to verify the cose models of the primary side. In addition, ,,
two small hot-leg-break loss-of-coolant accidents (LOCAs) were analyzed to investigate the effects of vent-valve flows on downconer fluid mixing.
l Except for the small hot-leg-break cases, all calculations showed significant primary-system depressurization followed by repressurization if the
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HPI was not throttled. System repressurization did not occur for the small-break LOCA cases because the break sizes were too large. Some overcooling was obtained in all calculations as evidenced by highly subcooled liquid temperatures in the downconer. The most severe transient in terms of 1
overcooling was the TBV transient in which both banks of TBVs were assumed to fail open, and the least severe was the PORV-LOCA transitat. Table Il summarizes the key results calculated for these overcooling transients. Not all the transients were run to 7200 s because of computer-time limitations. Once f
the primary system had stabilized, the calculations were terminated, sometimes as early as ~1500 s. For these cases, the results were extrapolated to 7200 s using engineering judgment. The results of these extrapolations are presented in Appendix B. An assessment of the infitence of uncertainties in the calculations is included in Appendix C.
It is recommended that other calculations be performed to fully address the Oconee-1 PTS issue. Specifically, other operator sctions should be considered to fully cover all possible overcooling scenarios. Also, in the case of the small-break LOCAs, other break sizes and locations should be investigated.
II. TRAC INPUT MODEL DESCRIPTION AND STEADY-STATE RESULTS l At the time of these calculations, the Oconee-1 model developed for the PTS study represented the most comprehensive modeling of any nuclear power plant assembled for use with the TRAC code. The model contains a primary side, a secondary side, and a complex control system consisting of both trips and controllers. This model operates at steady state over a pressure range from 4.01 MPa (~1.5 psia) in the condenser of the secondary side to ~15.2 MPa (~2200 psi) in the primary side and over a temperature range from ~300-590 K (40-600 oP).
~5-TABLE II ,
TRAC OCONEE-1 TRANSIENT RESULTS Minimus Minimus Downconer Cold-leg System Temp. Pressures Temp. Repressuri-Transient (K) (MPa) (K) ization? ._
- 1. MSLB a.
RCPresta{t;HPI 475 6.5 A-Loop 475 Yes throttled B-Loop 400 b.
NoRCPrestarg; 450 8.5 A-Loop 435 Yes HPI throttled B-Loop 445
- c. Same as (1.a.) 405 3.5 A-Loop 402 Yes with EFW B-Loop 422
- 3. TBV failurec (One Bank)
- a. SG 1evel control 365 17.0 A-Loop 448 Yes fails B-Loop 375
- b. SG 1evel control 440 17.0 A-Loop 477 Yes does not fail B-Loop 430
, c. RCP restart; EPI 430 4.0 A-Loop 491 No throttled B-Loop 491
- 4. TBV failurec (Two Banks)
- a. SG 1evel control 350 17.0 A-Loop 441 Yes fails B-Loop 446
- b. SG 1evel control 465 17.0 A-Loop 463 Yes does not fail B-Loop 465
- 5. SBLOCAc (2 in. hot-leg) 425 1.0 A-Loop 370 No B-Loop 410
- 6. SBLOCAc (4 in. hot-leg) 320d 0.5 A-Loop 430 No B-Loop 425
- 7. Rancho-Seco type 452 14.0 A-Loop 450 Yes e B-Loop 450 transient l
- This is the system pressure at the time of minimum downcomer temperature.
b For these MSLB calculations, the EFW system did not actuate because of input errors in the ICS and trip logic.
cThese calculations were extrapolated to 7200 s and the temperatures, pressures shown represent estimated values.
d For this calculation, the minimum temperature corresponded to temperature
" spiking" as a result of accumulator injection.
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l
- A. Primary Side l
The primary side of the Oconee-1 model is similar to ,other TRAC models that have been used (Fig. 1). It consists of the three-dimensional' vessel, two hot legs, two once-through steam generators, four cold legs, and some parts of j the emergency core-cooling system (the low pressure injection system was not . , ,
modeled).
t The volumes, elevations, and pipe lengths from the Oconee-1 plant are i
- closely matched in the model. The wall areas and thicknesses of the primary piping are also modeled so that the thermal response of the piping is predicted.
Heat transfer from the pipe walls to the environment is also allowed.
- 1. Vessel. The three-dimensional vessel is made up of 56 cells. These cells are arranged such that the vessel is divided into eight axial levels, two radial rings, and six azimuthal segments. The lower plenum is made up of one level; the core four levels; the upper plenus, two levels; and the upper head region contains one level. The two radial rings are divided so that the core region le bounded by the inner ring, and the downcomer annulus is modeled in the outer ring.
The vessel metal structures and wall masses are nodeled by assi. Ing a representative thickness and area of a heat slab to each three-dimensional cell.
Using these heat slabs, the stored energy of the vessel structure is approximated. The thermal conductivity for each heat slab is calculated by assigning five nodes across each slab thickness. These nodes are spaced so that they are closer together on the fluid of 3e of the slab. There is no heat transfer through heat slabs between cells, that is, each cell's heat slab is isolated from any other cell's heat slab except through the fluid-dynamic coupling.
The six azimuthal divisions were chosen to allow for the angle and separation of the penetration of the hot and cold legs. Also, the accumulator injection ports were modeled closs to their exact positions (both a:.ially and azimuthally) in the vessel downconer region.
l With only one radial division for the inner portion of the vessel modeled, the circulation of hot water rising up from the core into the upper head, then back down and out the hot legs is lost. Without finer noding of the vessel, the upper head is effectively isolated from the rest of the vessel in many circumstances. To alleviate this problem, a connection is made between the upper head and the hot legs so that about one third of the normal steady-state flow passes through the upper head. This connection is called the upper-head tee. , ,
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Primary-side model for Oconee-l.
n ~ ,
The vessel includes vent valves that are modeled to allow flow from the upper plenum directly to the downconer. This flow path is only available when the upper plenum pressure is higher than the downconer pressure. A composite vent valve made up of one-sixth the total vent-valve area is modeled in each azimuthal cell of the inner radius of axial level 7. Figure 2 shows the vent-valve model used for these calculations. The vent valves are fully open when .
the pressure drop between the upper plenum and downcomer exceeds 0.12 psi.
- 2. Hot Legs. The loop-A and loop-B hot legs are modeled symmetrically except for the surge-line connection to the loop-A candy cane. The candy canes represent the highest elevation in the system. The surge line includes a "small-break", which is activated for the small hot-leg break transients.
The pressurizer is modeled with a very small cell at the bottom and two small cells at the top. This was necessary to allow correct fluid conditions to either enter or leave the pressurizer.
Pressure relief for the primary system is provided by a single PORV at the top of the pressurizer. This valve supplies adequate relief for the cases where secondary cooling is provided and the reactor has tripped (all cases in this study).
- VENT VALVE MODEL ISL7
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ti 8
9.
as 00G E12 AP(INSIDE-0Uf51DE) (PSI)
Fixed closed value allows leakage for AP<0.08 Fixed open value for AP>0.12 Closed value based on design leakage (W. Jensen. NRC)
Open value based on BAW-1628 table.
' Fig. 2.
Vent-valve model.
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- 3. Steas Generators. A camplete model for BE once-through steam ,
generators was developed for this study. The primary side of the steam generator is made up of 12 cells whereas the secondary side uses 26. Heat transfer from the primary occurs through the steam generator tubes to the secondary coolant. Celin 2-11 (Fig. 1) represent a composite of the volume ,
inside and wall area of all tubes, and f e is in these ec11s that heat transfer from the coolant to the walls of the tubes occurs. Cells 1 and 12 model the upper and lower plenuns. The wall area and thickness of the plenuma are modeled so that the heat capacitance and heat transfer of the external wall can be calculated.
The secondary of the steam generator is divided inte four components (Fig. 3), two model the tube-bundle region, one models the downconer, and the final one models the steam-outlet annulus. The tube-bundle region is made up of 10 cells (SG components 32-1 and 12-2, for loop A and 2-1 and 2-2 for loop B in Fig. 3). These cells acdel the total volume and the tube wall area for the region between all the steam generator tubes, but model the heat-transfer characteristics of a unit cell of tubes. The top 4 of these 10 cells comprise the superheated-steam region of the model during normal full power operation.
The connection at the top cell of the tube-bundle region is . for the emergency-feedwater (EFW) flow and the realianed main-feedwater flow. This connection closely models the correct plant geometry so that any flow from the " upper header" is correctly injscred into the tube bundle.
i The steam generator downconer consists of seven cells (SG Components 12-3 and 2-3 of Fig. 3), plus one cell for the main-feedwater injection and one cell for the aspirator port. During normal operation, the downconer condenses enough superheated steam drawn through the aspirator port to heat the main feedwater to saturation temperature prior to entry into the tube-bundle region. The final six steam generator secondary cells model the steam-exit annulus. The outlet from this annulus is near the midpoint of the steam generator, close to the actual location in the plant.
Because of insufficient information regarding pressure tap locations on i the SG secondary side, the SG water levels were modeled using a collapsed liquid level calculation. This method is adequate for transients in which relatively slow changes in the secondary side occur, but may not be accurate for rapid changes such as in a MSLB transient. However, the response of a pressure
- transducer also say not be accurate for a violent secondary-side transient. '
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- 4. Cold Legs. All four cold legs of the plant are modeled. These legs each consist of a loop seal (lowest primary-loop point), a reactor coolant pump (RCP) and the HPI cosnection. The HPI nozzles are positioned closely to their correct height and distance from the vessel entrance. The RCPs are speed controlled during steady-state operation to obtain the required mass flow, but their steady-state speed is maintsined fixed while they are running during a .
transient. During normal operation, the four cold legs of this model have synaetric flows.
- 5. Emergency Core-Cooling System (ECCS). Two major components of the ECCS were included in this model; the HPI system, and the accumulators (core flood tanks). If a transient were to be run that caused the primary pressure to fall below the low pressure injection system (LPIS) setpoint, then the LPIS would also have to be added to the model.
The HPI system was modeled as four boundary conditions that can inject 283 K (50 0F ) water into the primary through the four side nozzles of the cold legs. These nozzles are located so that they enter the main cold Itgs from the side at an elevation somewhat higher than the centerline of the cold / hot legs.
The pressure-dependent flew rate of the two loop-B ports is identical, as is also the case for loop-A, but the loop-A capacity is greater than that of loop-B.
There is an accumulator tank for each loop that allows emergency coolant to flow directly into the downcomer region in axial level 7 of the vessel (Fig. 1). The accumulator flow is controlled by check valves such that when the primary pressure falls below the accumulator tank pressure, the check valves open. The initial accumulator pressure was 4.2 MPa (410 psia) with a coolant temperature of 305.4 K (90 0F ). .
B. Secondary Side All major components of the secondary side are modeled except for the turbine generator equipment and various valves that were not necessary for any transients of immediate interest. With the exception of the vessel, the secondary side required much more modeling detail than was previously necessary when only the primary system was modeled.
- 1. Feedwater Train. In this discussion, the feedwater train describes the secondary-side modeling from the condenser (component 55 in Fig. 3) to the tee where the feedwater is divided between the two steam generators (component 38 in Fig. 3). This section of the modeling takes the fluid discharge from the turbines and raises it to the temperature and pressure at which f: is delivered to the steam generators.
. -, , , , , . , _ ,_ _ . - - - - _ _ - - c ,
The condenser is modeled as a large tank with a very large wall area. The i thin valls have a high-thermal conductivity and a constant-temperature heat sink on the outside surface. This model fully condenses the incoming steam from either the turbine exhaust or the turbine-bypass system.
The hotwell is an even larger tank used for the collection and storage of the condensate. The lowest system pressure and temperature occur in this .
component. Cell 1 of the hotwell actually represents the volume and coolant inventory of the upper surge tank. It is included to reduce the complexity of the model while still providing an estimate of the available hotwell inventory.
The coclant supplied to the emergency feedwater system is taken from the hotwell/ upper eurge tank cos:bination.
The hotwell pump (component 51) includes the desineralizer/serator section of the feedwater train. The model accounts for the effects of this section by including additional frictional losses and heat addition to the coolant.
Each feedvater heater is modeled to achieve a feedwater temperature rise close to the design value for that heater. In addition, the model includes a time-dependent estimate of the feedwater-heater heat capacitance. The heaters are modeled with four heat-conduction nodes (Fig. 4) such that the first two cross the metal wall and the outer two model the secondary-side steas/ water mixture. The energy input from the extracted turbine steam is modeled by adding l a volumetric heat source to the middle ocde of the steam / water mixture. This heat source is controlled by a trip so that it can be ramped off fc11oving a turbine trip.
The two parallel MFW pumps of the actual plant are combined into a single pump for this model. This is a variable speed pump with the speed determined by the ICS. This pump will also be tripped off if any of a variety of setpoints are reached as described in the control section of this report (Sec. II.C).
The coolant flow from the feedwater train splits to provide flow to the steam generator conttol-valve section of each major loop. Because the loop-A
! and loop-B flow-control valves are identical, a discussion of the valves for only one loop 14 necessary.
- 2. Stean-Generator Control Valves. A schematic diagram of the control-valve arrangement is shown in Fig. 5. This schematic can be correlated to the two contsnl-valve sections of the secondary-side noding diagram of Fig. 3.
There are three check valves (valves 1, 2, and 3) in this arrangement to stop reverse coolant flow from the steam generator into any of the feedwater lines. ,
The emergency-feedwater valve (IFWV) is closed during steady-state operation, l
l ENERGY INPUT FROM TURBINE EXTRACTION STEAM q" ,
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l 1
eedwater Coolans g,[M.Y,r.-W/g., Flow )
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1: Emergency-Feedwater Check Valve 2: Realignment Check Valve 3: Main-Feedwater Check Valve 4, 5: Feedvater Realignment Valves EFW: Emergency-Feedwater Valve SUFCV: Startup Flow-Control Valve .
MFCV: Main Flow-Control' Valve Fig. 5.
Stean generator control-valve schematic.
- but opens and closes as necessary if EFW is demanded. (Refer to Section II.C.1 '
l for further description of the EFWV.)
The remaining four valves in Fig. 5 provide the major control for the MFW j flow. The valve area of the two flow-control valves is controlled by the ICS.
During steady-state operation, the start up flow-control valve (SUFCV) is 100% e et open and the main-flow-control valve (MFCV) is approximately 50% open. The action of these valves during a transient is generally difficult to predict because of the complexity of the ICS. In general, the MFCV closes following a reactor trip, and the SUFCV controls any feedwater flows below approximately 15%
of the steady-state flow. The action of the flow-control valves may be overridden by signals from the trip system (refer to Section II.C.1).
When one of the two realignment valves, valves 4 and 5, is open, the other must be closed. During steady state operation, or any time before the feedwater-realignment trip is hit, valve 4 is closed and valve 5 is open.
Following the realignment, valve 4 opena and the MFCV and valve 5 close.
There are several different flow combinations that can occur. During i steady-state operation, the flow splits along the parallel paths through the two flow-control valves, then rejoins to flow into the MFW header. The flow split is 85/15 with the larger flow through the MFCV. For flows less than 15% only a single path is open to the MFW header, and that is through the SUFCV, valve 5, and check-valve 3. If feedwater realignment occurs, the flow is from the feedwater train through the SUFCV, valve 4, check-valve 2, and into the EFW header. The EFW may also be running so that flows may be entering both SG headers, or the EFW may six with the MFW flow before they enter the EFW header.
If the steam generator is isolated, the MFCV, SUFCV, and EFWV are closed so that no flow can enter the steam generator. '
- 3. Emergency Feedwater. The EFW system is modeled so that it takes water from the hotwell and delivers it to the EFW7s of both steam generators. The EFW 4
pump is modeled as a composite of the turbine driven and motor driven pumps in the actual plant, and therefore has a large capacity. The EFW is delivered to a !
l tee (component 152, Fig. 3) that splits the flow between the two steam I
generators. If both EFWVs are open, the EFW flow is symmetrically split as long es the secondary pressure of the two steam generators is equal. If only one EFWV is open, the full flow available is delivered to that stese generator. The EFW flow stops if the hotwell inventory has been depleted.
- 4. Steam Lines. The steam line for steaa generator B is longer than the l steam generator A steam line. Other than this difference, the steam lines and l their valves are identical. A pressure boundary condition models the steam flow
}
l .
_ . . l
l . . . . . . . . .
-u-
! exiting the secondary into the turbine inlet. During steady-state operation the I
steam-nass flow is modeled to reenter the secondary as boundary condition inlets I to the condenser and the two heater drains. Following a turbine trip, the turbine-stop valves (TSVs) close the steam lines. If any pressure relief from j the closed steam lines is necessary, steam is released through the turbine-bypass valves into the condenser. For a normal reactor / turbine-trip -
! transient, the steam relief from the turbine-bypass valves is adequate so that modeling the main-steam safety valves is not necessary. Following a turbine
, trip, the secor.dary model is a closed loop unless a transient similar to a MSLB j
is modeled. For the MSLB, one of the turbine inlet boundary conditions is set to constant containment pressure and the corresponding turbine-stop valve is fixed open.
C. Control System The control system for the various components of the TRAC input modal is provided in two ways; with trips and with control blocks. Trips basically turn
. something on er off depending on certain conditions being mec. In th's model, control blocks use mathematical relations between system variables to adjust valve areas or pump speeds. The control blocks are used to model the B&W ICS.
- 1. Trips. Of the more than forty trips used in this model, only five are simple
~
trips. Simple trips have only a system variable as input, and their i
output is only used to control some component action. A summary . ! the simple trips is presented in Table III. The TBVs are actually controlle* by the.ICS at the Oconee-1 plant, but because they have a single setpoint for the transients calculated for this study, simple trip modeling is adequate. The only noticeable difference with this modeling appears in the secondary-side pressure plots for transients with steam-line pressure relief for extended periods. In-these cases, the plots appear somewhat saw-toothed because of the full opening or closing of the TBVs, whereas in the plant the ICS maintains a smoother pressure response by allowing partial valve openings.
The reactor trip system is presented in Fig. 6. As depicted, the signal output from the reactor trip is input to the TSVs, the condenser feed, the feedwater heaters, and the heater drains. Most cf these trips actually occur following a turbine trip, but in this model the turbine and reactor trips occur together. For all of the transients calculated for this study, the reactor trip j occurred immediately. If additional transients were to require the reacte trip to occur after other criteria had been r. :et , this model would have to be modified. The delays and rates are not ine*.uded in the informatior presented in Fig. 6, so these parameters will be discuased in the following paragraph.
.. ..e-1 TABLE III .
SIMPLE TRIPS l
Description System variable Setpoint (MPa)/Acticn TBV control- Steam-line pressure 7.064/ opens loops A and B 7.014/ closes Accumulator check- Check-valve pressure drop 0.14/ opens valves, both loops 0.05/ closes PORV control Pressurizer pressure 16.99/ opens
- 16.65/ closes
~
The reactor trip is -0.5 s from the beginning of the transient. To model i the insertion of control rods into the core, a negative reactivity insertion of
-0.0536 ok is added ~1.0 s af ter the trip. The decay heat is calculated using the American Nuclear Society (ANS) decay-heat constants that are internal to the code. The turbine-stop valves start closing simultaneously with the reactor trip and take ~1.0 s to fully close (except in the MSLB transient where the loop-A TSV is fixed open). The condenser feed trip occurs ~1.0 s af ter the reactor trip and the feed decays to zero over 5.0 s. The volumetric heat sources used to model the feedwater heaters and the feedwater added to the train.
l through the heater drains are tripped -0.5 s af ter the reactor trip and also take 5.0 s to decay to zero.
Each box in Fig. 6 is divided into four sections. These sections are for the trip identification number, the trip description, trip input, and trip output. The trip output either goes to a component or to another trip as its i input. The component numbers are enclosed in small circles and correspond to l
l one of the components in Figs.1 or 3. These trip system figures were developed as part of the TRAC-PF1 modeling work and are presented for clarity in understanding the interconnection between them. For further explanation of these figures, plesse refer to the trip system legend contained in Appendix A.
9
- 1. . . .. ...
17 I11 5II asector Trip Tsv - A EV. Tune 2860. E dos 0=1 o=I maa Tsv - 3 asse. E >oS 0=1 sell Condenser Food asse. E >os o=I a
sell Heaters Drains asse. E >os 4 o=1 Ig Fig. 6.
, , Reactor trip system. -
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~ -
Figure 7 presents the trip systen that controls 'the HPI, the Jour RCPs, and the MFW realignment valves. HPI . initist' ion occurs if the system pressure drops be' low 16.44 MPa, but is also cyciel 'on and off to keep the least cooled hot leg at 42 7 K subcooled (75, 12.5 F). For some transients reported, the HP1 subcooling monitor was not used. For those cases in which a particular trip
( is not wantede it is~ left in the system but given a sAtpoint that cannot be I
reached. All four RCPs also'use the HPI low =pressupe setpoint, but have a
, n .,
30.0 m delay from the time the setpoint is reached until the pumps trip. The A1 and El pumps are under separate trip control from the A2 and B2 pumps so that these pisaps (Al and B1) can be tucesd back on if the least cooled hot leg reachis 42 if "subcooling. For some of these transients, the pump subcooling 4 ss sonitorn is'also disabled..
. Folloking ' any'RCP crip. . the -MFW is realigned to flow into the steam
~
generatort'through the EFW header. This realignaant occurs by taking control of the MFCVs avsy. from the ICS and rasping them closed at a flow area fraction 1 e s
-o A A-
[ NPI PremL )
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( TE 443 / "
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HPI, RCP, and MFW realignment trip system.
closing rate of 25% per s, and by opening or closing the appropriate realignment valves over a 5.0 s period. Further information on this valving arrangement is described in Section II.B.2 of this report.
The trip system for the three pumps of the feedwater train is presented in Fig. 8. The hotwell pump trips only if the hotwell level falls below 0.1524 m.
l If this pump trips, then the other two aust also trip. The condensate booster pump trips if the hotwell pump has tripped, or if its suction pressure falls below 0.21 MPa. The MFW pump trips if any of the six trips feeding into it are actuated. These six trips include the two upstress pump trips, either SG 1evel greate-; than 9.27 m, a suctica pressure less than 1.72 MPa. or a discharge pressure greater than 8.89 MPs.
mg g * ** e n e- >
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l 1
The SG isolation trip system for both loops is presented in Fig. 9. This system is designed to ensure that both flow-control valves and the EFW valve close if steam generator isolation is required. The two reverse-value trips are necessary to make the three input values of the MFCV override trips compatible.
If one of the SG isolation trips is turned back off, then EFW can be delivered to that steam generator if it is required, such as in the MSLB transiant. .,
The controlling trips for the EFW system are presented in Fig.10. The trip conditions that demand EFW are at the top of the figure, and the conditions to throttle EW are at the bottom. EFW is demanded if the MFW discharge pressure drops below 5.271 MFa. The EFW demand opens the appropriate EFW valve at a fractional rate of 33% per s. If either EW valve starts to open, the EFW pump comes up to full speed in 4 s. The EFW valves close if the SG water level goes above 6.2 m, but reopens when the level drops below 6.0 m. The EFW valves
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Hotwell, condensate-booster, and main feedvater pumps trip system.
I \
( SV.001 ) [ "N' I
\' ' ' ( *nos. E >es )
N~Y 2711 3411 3311 EFW Valve-A SUFCY-A WFCV-A Centrol Dwerride Clamed Dverride Closed
- 3 FTPs. W >05 == -= 3419. E <03 =* 3319. E >05 [ **
l -:- : o.- o.-
( XV. Col ) [Beverse4002)
Isrf Velve i
t s ( 4019. E >CS }
t- yy 2722 3422 3322 EFT Velve-s SUTCV-B W KV-s Control Dwerride Claed Dverride Closed
- 3 2TP9 W dos *- -= 3429. E <c5 --* 3329. E >03 C**
l -t = 1 0 o -1 o . -1
~
Fig. 9.
SG isolation trip system.
also close if the hotwell level is less than 0.0254 m. If both EFW valves ~
close, the E W pump trips off.
- 2. Integrated Control System. The B&W ICS matches the feedwater flow with the power demand and maintains a constant steam-line pressure t.nd adequate
) superheat. The ICS can quickly respond to plant load demands while maintaining l smooth plant operating parameters. The block diagram of Fig.11 depicts how this control is accomplished. The ICS controls secondary pressure with the turbine valves, primary power with the control rods, and primary-to-secondary heat-transfer characteristics with the SG valves and MFW pump. Cross limits are sent back to the Integrated Master so that if some section is not perf,caing adequately, the other parts of the ICS balance the response to prevent any power, pressure, or feedvater flow mismatches.
e me.
e
. .= _. .
( .
G -
A bt."e 7,J &
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( CA -11N ) \ ( CA -2384 ) .
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j l Demand }
m 2719. E 208 )- --e( 2PB. E PCS h N'Y N~Y FF 1 2F22 EFW Valve-A EFW Valve-B Control Control
- i 27F8. M dos m- --e 27P9, M dos F*
- t == 1 Ett -1 I 300s. E >-OS :
0I f n-A Invel)
Op. Range ryo [>Steen-B tevell i NP4 Op. Range]
( EV. 011 ) l
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( EV. 0 n /
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- see agen teolatwn rigure
- Fig. 10.
T.aergency feedwater trip system.
The pecformance of the ICS aust be taken into account for accurate' simulation of most plant transients. In particular, the feedwater flow is under ICS control and does not have a simplistic response to most transient conditions.
The reactor is tripped instantly for all transients that were calculated in this study. This allows a great deal of simplification in modeling the ICS.
As is shown in Fig.11, the TBV section is modeled with simple trips (refer to Section II.C.1) and only the feedwater-control section needs to have detailed control-block modeling. If a transient were run with the reactor not tripped, i
ICS control of the TBVs would be necessary. Following a reactor / turbine trip, however, the turbine steam-flow control valves do not have to be modeled.
Auto Dispatch System l
Unit Imad Demand Integrated Master i I l l-Pressure Resetor Feedwater Cont ol Control Control l l l l~ l l Turbine Turbine Control Stgen A MFW nB Valves 7,@ Dr es Valves Pumps StNves V
- modeled with TRAC control blocks
- modeled with TRAC trips Fig. 11.
ICS organization.
The control-rod section of the ICS does not have to be modeled for these -
transients because of the early reactor trip nor is it needed for steady-state model operation because the reactor is at full power. If a transient were run where the rods were not inserted, this section of the ICS would have to be
- modified to allow correct reactor power control. Because of these simplifications, the only part of the Integrated Master modeled was the neutron-power cross-limiter. -
Figure 12 presents a block diagram of the feedwater-control section of the ICS as it is modeled for this study. The British Thermal Unit (BTU) limiter monitors some primary and secondary parameters to determine if enough feedwater is flowing. This determination is compared against the reactor power to reduce the feedwater flow if the power is dropping. The requested feedwater flow is then compared with the actual feedwater flow to get a feedwater-flow error.
l This error value will then open or close the MFCV unless the SC level has l reached either a high- or low-level limit. If limited, the valve action will then act to bring the level inside of the limit. The requested feedwater flow is also used along with the MFCV pressure drop to determine the MFW pump spee*h Each of these blocks is separately discussed in the following paragraphs. This discussion is limited to just the information necessary to describe the loop-A mo
model. The loop-B model is similar except that it uses several of the signals generated for the loop-A side to avoid duplication of coding. A complete .
listing of the flow diagram and mathematical representation of ,the TRAC-PF1 ICS model is included in Appendix A.
A diagram of the BTU-limiter control blocks is given in Fig. 13. Each of the blocks represents one signal manipulation. As depicted in Fig.13, there '~
are four system parameters input and one signal output. The letters in each block represent a particular control-block output and identifies the mathematical representation of the control blocks as they appear in Table IV.
Each of the block outputs is in volts and the limits are included with the equations where appropriate. The system parameters input are always in SI units. The steady-state output voltage for each of the blocks is given in the table. A BTU-limiter output value of 8 volts indicates the correct feedwater flow. For a value less than 8 volts, less flow is rueded, and for a value greater than 8 volts, more flow is requested.
The neutron power cross-limiter section of the TRAC-PF1 ICS model is presented in Fig. 14. This section has three input signals and one output voltage. Table V gives the equations and steady-state values that correspond to these control blocks. Blocks Al and B1 set up a 20% per-minute signal ramp DTU Limiter Neutron Power Cross Limiter rm L .f Feedwater MFCV Pressure Flow Drop Level MFW Pump Limiters Speed Control Feedwater Valves control Fig. 12.
Feedwater-cbntrol section of the ICS.
. . - l Hot lag A Hot lag A Stgen A Feedwater '~
Mass Flow Temperature Pressure Temperature O
I
=W
.a g i_
Fig. 13.
BTU-limiter control blocks.
TABLE IV BTU-LIMITER CONTROL-BLOCK EQUATIONS Steady-State Equation Output Voltage i A = 0.00204083
- RCFLOWA 18.0 B = -605.4459 + 1.04092
- RCTEMPA , -10.0 < B < 9.080 8.0 C = 82.549 - (1.16958e-05)
- SGPRESA , -1.0 < C < 9.080 8.0 D = -11.036 + 0.037260
- FWTEMP , -1.270 < D < 9.080 8.0 E = -16.0 + B + C + D , -10.0 < E < 12.0 8.0 l F = 0.55555 + 0.055555
- E 1.0 H = -10.0 + A
- F 8.0 af ter the reactor has tripped to ensure a feedwater runback of at least this rate. The control-block tables such as D1 and K1 use simple linear interpolation between points to complete the function values. All output block signals are in volts except for D1 (K) and J1 (watts). The DELT in block JL is Se
for time-step size. The final output from thfs section results in similar
- feedwater control as in the BTU-limiter section.
Figure 15 presents the feedwater-flow control blocks. The BTU-limiter voltage input comes from block R, and the neutron cross-limiter input from block i SG. The blocks depict the selection of one limiter signal and the comparison of that signal with the current feedwater flow. The equations that sake up this ,,
section of the ICS are presented in Table VI. All blocks in this table are j output in volts except for block SL, which is in kg/s. A negative output from
) this section tends to close the feedwater-control valves; a positive value opens them, and a zero value indicates no change is requested.
The steam generator level-limiter section of the TRAC-PF1 ICS model is presented in Fig.16. This section passes the feedwater-flow comparison signal straight through unless the high-or low-level setpoints are reached. The Resclor Power Reactor Trip Feedwate%
Time Temperature SP JL Al J1 B1 C1 D1 El-H G1 F1 11 K
SG Fig. 14.
Neutron power cross-limiter control blocks. l l
t
--.-e ~, -. ._.
TABLE Y ,
NEUTRON-POWER CROSS-LIMITER CONTROL-BLOCK EQUATIONS Equation Steady State A1 = TIME - (TIME OF REACTOR TRIP) 0.0 .
B1 = 1.0 - (0.2/60.0)
- A1 , 0.0 < B1 1.0 C1 = 18.0
- B1 18.0 D1 = f(C1) : C1 D1 460.0 0.0 204.0 0.562 240.0 3.6 320.0 5.4 356.0 9.36 402.0 18.0 460.0 21.42 483.0 El = -460.0 - D1 + 1.8
- TVTEMP 0.2 F1 = 1.0 + 0.0013
- El 1.0 i
G1 = F1
- C1 18.0 SP = POWER - 2568.0e6 0.0 First order lag of power with 4.5-s time constant JL = JL + ((SP - JL)/4.5) e DELT 0.0 J1 = 2568.0e6 - JL 2568.0x106 l
H1 = 1.6 + 14.4
- B1 16.0
~
Il = -1.0 * (H1 - 6.23053e-9
- J1) , -10.0 < Il < 10.0 0.0 K1 = f(II) : Il K1 0.0
-1C5 -1EO
- 0.5 0.0 0.5 0.0 10.5 10.0 SG = -10.0 + K1 + C1 8.0 high-level limit block (P1) is generally a large positive number until its setpoint is passed. As this setpoint is passed, P1 quickly becomes a large negative number that closes the feedwater-control valves. There are two low-level limit setpoints that can be used. ' The lower value (0.6096 m or 24 in.) is for normal operation. If a RCE trip has occurred, the higher value 4
. l Loop A BTU Neutron Power Feedwater Flow Limiter, H Cross Limiter SG . .
I n
m SL V
Fig. 15.
Feedwater-flow control blocks.
TABLE VI FEEDWATER-FLOW CONTROL-BLOCK EQUATIONS Equations Steady-State Value R = min (SC,5) 8.0 FWB = FWFLOWA - 680.4 0.0 ,
First order lag of feedwater flow with 1.0-s time constant FSL = FSL + ((FWB - FSL)/1.0)
- DELT 0.0 SL = FSL + 680.4 680.4 Si = 10.0 + R - 0.026455
- SL 0.0 (6.096 m or 240 in.) is used to enhance natural circulation by maintaining a higher steam generator level. When the low-level limit is reached, P2 changes l from a large negative value to a positive value that is passed on to T1. T1 is the final feedwater-flow error used by the valve-control section of the ICS.
l Table VII gives the equations and steady-state values that correspond to these l control blocks.
- - ,, ,,n,. , ,. ., ,. ,-, ,-.,g. , .. , , y,
._ _ _ - _ _ = ._ . - - . _ _ _ - _ _ _
The ICS section that adjusts the flow-control-valve area is presented in .
Fig. 17. This section delivers a flow area to the two flow-control valves cf this loop (Components 30 and 36) by using a proportional and integral controller in which the error signal is integrated and proportioned to determine the flow area. Two sets of constants are used for this controller depending on whether a
- the ICS is low-level limited or not. The equations describing this action are given in Table VIII. If low-level limited, larger values for the controller are used to speed the opening valve action.
The MFW pump-speed control is determined by signals from both loops. The resulting voltage obtained from the comparison between the BTU-limiter output and the neutron cross-limiter outout is used along with the minimum of the two MFCV pressure drops to obtain a pump speed. The control block diagram for this section of the ICS is presented in Fig.18, and the corresponding equations in i
Table IX. There is a constraint on the rate of change of pump speed built into TABLE VII LEVEL-LIMITER CONTROL-BLOCK EQUATIONS Equations Jteady-State Value
- operating level scale, 96 to 388 in (level in meters) l BL1 = f(ALEV) : ALEV HL1 -10.0 2M -15 5 9.855 10.0 P1 = -2.0 * (EL1 - 7.0) 34.0 l .
l Q1 = min (S1,P1) 0.0 l
- startup level scale, 0.0 to 250 in l LL1 = f(ALE)) : ALEV LLI -3.3 l
0.0 -10.0 6.350 10.0
- pumps tripped: 240 in = 6.096 a = 9.2 Y IF(PTRIP .EQ. 1) STP = 9.2 pumps running: 24 in = 0.61 a = -8.08 Y IF(PTRIP .NE. 1) STP = -8.08 -8.08 P2 = -2.0 * (LL1 - STP) -9.6 T1 = max (P2,Q1) 0.0
.. .. .-..s.
Feedwater Flow Stgen A RC Pump Trip Signal Comparison, S1 Water Level L3 L b a i&-
l
.I V
Fig. 16.
Level-limiter control blocks.
Feedwater Flow Low kvet Error, T1 Limit Signal. P2 h lCNST1H Q1) lCNST2}=--- .
fSUA lMrA i 1 Isural lurCal l I b [P to gmp Fig. 17.
Feedwater-valve adjustment control block.
~
TABLE VIII .
FEEDWATER-VALVE ADJUSTMENT CONTROL-BLOCK EQUATIONS Equatfg Steady-State Value
- if 72 < 0, low limit has not been hit IF(P2 .GE. 0.0) CNST1 = 0.12 0.1125 IF(P2 .LT. 0.0) CNST1 = 0.1125 IF(P2 .CE. 0.0) CNST2 = 2.4 0.9 IF(P2 .LT. 0.0) CNST2 = 0.9
- integrate, T1, is the last time-step value of signal T1 U1 = U1 + CNSTI * (T1 + T1 )/2.0
- DELT , -18.0 < U < 2.0 0.0 Ill = U1 + CNST2
- T1 0.0 i
Il = X11 + 8.0 , -10.0 < Il < 10.0 8.0 SUA = 64.1164 + 7.44164
- X1 , -10.0 < SUA < 10.0 10.0 SUFVA = 0.1
- SUA , 0.0 < SUFVA < 1.0 1.0 t
MTA = 0.5555
- X1 - 4.4444 , -10.0 < MFA < 10.0 0.0 MFCVA = 0.5 + 0.5
- MFA , 0.0 < MFCVA < 1.0 0.5 the TRAC-PF1 MFW pump model (27 rad /s/s) so that an additional constraint was not needed in the ICS model.
D. St.eady-State Calculation The primary-side steady-state operating conditions for the IRAC model are presented in Table X along with operating specifications from the Oconee-1 plant for comparison. The two primary-system loops and all four cold' legs had symmetric flows. The mass flowrates through the reactor pumps were controlled during the steady state by adjusting the pump speed with control blocks. During the transient calculations, the calculated pump speeds were held constant as long as the pumps were running. All steady-state primary-loop values compared well with available data with only slight discrepancies in the core and vessel
31 e
Loop A Limiter Loop B Limiter MFCV-A MFCV-B Comparison, R Comparison, BR AP AP 1 1 lDPAB] lDPBBl ..
1 4 lDPALl lCPBLl b A DPA DPBl b b FC -
FD M
FP s to Comp 49 Fig. 18.
MFW pump-speed control blocks.
pressure drops. In Table I, the coolant velocity and loop-flow lengths are added to show the relative time it takes coolant to travel through the loop from vessel exit to inlet.
Table XI presents a selection of secondary-side steady-state operating ,
conditions for this model. In general, the comparison is very t ,od with available Oconee-1 plant data. The fee:twater flowrates for the two loops were independently controlled by the ICS through control of the MFW pump speed and MFCV areas. The steam outlet flowrate was approximately 2% higher than the feedvater flow so that the steam generator inventory is slowly depleted with time. The approximate time it takes coolant to travel from the hotwell to a steam generator is estimated from the average velocity and pipe length to be over 11 min.
l l -
\
l I
/
TABLE II ,~
/ .
MFW-PUMP-SPEED CONTROLyBLOCK I EQUATIONS
/
l l
Equations j Steady-State Value DPAB = DELPA - 3.55e5
/ 0.0 .
First-order lag with a 1.0-s time codstant
- 40 psi limit on both sides /
l DPAL = DPAL + ((DPAB - DFAL)/1.0)/ DELT , -2.4E5 < DPAL 0.0
< 2.4E5 '
/
DPA = DPAL + 3.55E5 / 3 55E5 FA = 2.90074E-5
- DPA - 10.0 O.2975
! FE = 0.2
- ABSD D), -10.0 < FE < 10.0 0.0
- DELT , -10.0 < FF < 10.0 0.0 FG = FF + FE , -10.0 < FG < 10.0 0.0 FI = 0.5 * (R + BR) - FG 8.0 FP = f(FI) FI FP rad /s 523.6
-E0 370.4 0.0 392.8 6.0 460.0 10.0 586.43
- * * * - - .e- %. .. ,
3 TABLE I PRIMAZY-SIDE STEADY-STATE CONDITIONS Oconee-1 Parameter TRAC Model Nuclear Station Power (MW) 2568.0 2568.0 -
Coolant flowrate, total (kg/s) 17640.0 17640.0 Hot-leg temperature (K)* 589.56/589.33 589.3 Cold-leg temperature (K) 563.81/563.81/ 563.5 563.61/563.61 Primary pressure (MPa)-
(3 m below top of hot leg) 14.96/14.96 14. 96 Core pressure drop (MPa) 0.117 0.11 Vessel pressure drop ,MPa) 0.378 0.41 Pressurizer water level (m) 5.63 5.59 HPI coolant temperature (K) 283.2 305.4 Accumulator coolant t .mperature (K) 305.4 305.4 Hot-leg coolant velocity (m/s) ~19.5 -
Coolant flow path length (m)
(external to vessel) 63.5 -
- loop A/ loop B, or cold-leg A1/A2/B1/B2 O
e l
l l -
1
TABLE II '
SECONDARY-SIDE STEADY-STATE CONDITIONS Oconee-1 TRAC Model Nuclear Station , ,
Feedwater flow, loop A/ loop B (kg/s) 679.04/678.38 680.4 Feedvater temperature (K) 511.19 511.0 Steam outlet flow, loop A/ loop B (kg/s) 693.8/692.5 680.4 Steam outlet superheat (K) 17.7/17.03 33.3 Steam outlet pressure (MPa) 6.37/6.37 6.38 Steam pressure at turbine inlet (MPa) 6.2/6.2 6.2 Steam generator secondary inventory (kg) 1.711E4/1.715E4 ~1.77E4 Aspirator steam flow (kg/s) 95.4/95.3 -
MFW pump inlet pressure (MPa) 2.77 2.68 temperature (K) 461.0 462.0 Condensate booster pump inlet pressure (MPa) 8.0 6.9 temperature (K) 308.3 309.0 Hotwell pressure (MPs) 0.015 0.01 temperature (K) 305.6 305.9 inventory (kg) 5.31E5 5.31E5 Upper surge-tank inventory (kg) 2.69E5 2.98E5 Feedwater train average coolant ~1.5 -
velocity (m/s) -
Hotwell to SG flow length (m) 1004.3 -
MFCV area fraction (%) 48.60/48.15 50 Loop flow fraction (%) 85/85 85 89 0
9
III. TRAC TRANSIENT CALCULATIONS .
A. Oconee-3 Turbine Trip
- 1. Introduction. As a benchmark case for the Oconee-1 FIS study the l Oconee-3 turbine-trip and steam-generator-overfeed transient of March 14, 1980 l was simulated. The actual transient is documented in Ref. 4, and measured data i
is available for the first three minutes of the transiert.* The data available . .
include primary- and secondary-system pressures, hot- and cold-leg temperatures, pressurizer and steam generator water levels, and main-feedwater flow rates and supply temperatures.
In this transient, the plant was operating at 100% power when the Electro-
! Hydraulic Control system caused a turbine trip and subsequent reactor trip.
After the reactor trip, the ICS malfunctioned and caused an overfeed to the steam generators, resulting in an overcooling of the primary system. The steam-generator water levels increased above the expected levels until the main-feedwater pumps tripped on a steam generator high-water-level signal.
The transient was modeled using the TRAC-PF1 code and the basic TRAC model of the Oconee-1 PWR (Fig.1), with modifications to the steam lines to add main-steam safety valves (MSSVs). The TBVs were also modeled. The condensate-heater feedwater-train modeling and the ICS modeling was not included for this transient because these systems were not necessary. Instead, the measured main-feedwater flow rates and feedwater supply temperature given in Ref. 4 were specified as boundary input in the TRAC calculation.
The TRAC-calculated results compared very well with the Oconee-3 data and general trends and major peaks and dips in the data were predicted. The actual transient had slightly more overcooling than the TRAC calculation. The calculated primary pressure, hot- and cold-leg temperatures, and pressurizer water level were sligh'tly higher than the measured data. The major differences between calculated and measured results were in the staan generator secondary pressures and in the steam generator B water level. The calculated secondary pressures cycled between the TBV open and close setpoints, whereas the measured secondary pressures dropped and remained below the TBV setpoints for the transient period modeled.
The measured secondary-side water level of steam generator B was found to be inconsistent with the measured main-feedwater flow rate. The steam generator refilled at a much faster rate than could be accounted by just the main-
- Data not shown in this report because it is proprietary.
_,,,._,,,.__,__..,,,.--r -w,_ y,, ,
i feedwater flow, which indicates that auxiliary feedwater might have been inadvertently delivered to steam generator B during the transient. This ,
inconsistency was not mentioned in Ref. 4 nor was auxiliary-feedwater flow data given.
A more accurate transient might be calculated if additional information can be obtained about actual auxiliary-feedwater flow rates, TBV setpoints and operating characteristics, decay power, and HPI flow distribution. ,
- 2. Model Description and Assumptions. The basic TRAC model of the Oconee-1 PWR, Fig. 1, was used to model the Oconee-3 transient. The steam lines were modified to add the steam safety valves as shown in Fig.19. 'Neither the condensate-heater feedwater train nor the ICS was modeled. Instead, the measured feedwater flow rates, Fig. 20, and the feedwater temperature were specified as boundary input in the TRAC calculation. The standard ANS decay power, Fig. 21, was used and is automatically calculated by TRAC. The Oconee-3 measured reactor power is also shown in Fig. 21. The Oconee-3 power curve was used in a second TRAC calculation to determine the effect of reduced decay
- power, even though it did not include gamma-ray heating.
Table XII presents calculated and measured steady-state conditions. There is excellent agreement in the steady-state values except for a small difference in primary system pressure. Table XIII presents the MSSV and TBV setpoints used in the TRAC calculation. These are Final Safety Analysis Report (FSAR) values, as actual oconee-3 setpoints were not available.
Table XIV shows the sequence of events . that occurred in the Oconee-3 transient. The primary-coolant pumps did not trip in this transient. No auxiliary-feedwater flow was assumed in either steam generator. The HPI flow was assumed to be divided evenly among all four cold legs.
- 3. Transient Calculation. The TRAC-calculated transient compared very well with the actual plant transient. The actual transient had slightly more overcooling than the TRAC calculation. The calculated primary pressure, pressurizer water level, and hot- and cold-leg temperatures were slightly higher l than the measured values. In general, differences between calculated and measured values were consistent throughout the transient.
Figures 22 and 23 compare the calculated and measured primary pressure and pressurizer water level, respectively. The TRAC calculation showed a decrease in pressure to a minimum of 12.97 MPs (1881 psia) whereas the actual transient decreased to a slightly lower minimum. The calculated pressurizer water 1evel dropped about 3.04 m (10 ft).
l 4
S t eam Line Modeling Af t er Reactor Trip and Turbine S t op Valve Closurc D .
2aooo @
woc ID1L __ .5 i 4 1 5 i s is2 E U
b Main Steam .
Safety Volves
\fr om @.
si.om g 3'[
t @ 70 @ (ussy)
$1"",,,,g Line A I
@ 916 @ i7 0 @
@ iocaasi2 0 @
J90C 41Lli 21 Q 9 ICE 19:30 @
3 I 4 l 5 I 6 1 7 is4 E ,
\ fr **
s t .am
- / ' @ 3CpiaCE>t40 @ O' G olcr 8 LI"* 8 h 1 20 h 3 ss By ss Turbine I" 4 Bypass Volve 4 -
_ i O s8 hW7,0 @ '
5 1 6 187 L_1_X._2.._Js6aO @ l h
Turbine Bypass Valve Fig. 19.
Main st'eam safety-valve modeling for Oconee-3 transient.
t e
i I
a00 , ,
. rno '
Leap A 700- *
- Loop B -
800< -
OSO 7 7
) SCO-
}
~
6 900
$ 400- -
t 750 c
E 300 - E
$ 5
.00 goo. -
go. . 2SC 0 , . , 0 0 20 40 to 80 WO 90 MO #0 20 200 Time (s)
Fig. 20.
Measured main-feedwater flow rates.
no Calculated ANS decoy power
*- haecsw ed e decoy power (Does not
,, include gomma-roy booting) ,
t B
E
, gC- -
3 8
E .
. .0 . .
, 5 i
go. .
L 0 .
l 0 20 40 to a0 =0 no w0 se no 200 ,
Time (s)
Fig. 21.
TRAC-calculated decay pcwer and Oconee-3 measured thermal power.
a s: ,e. . . , .
l l
ta e' 23c0 Celculated
~ heeosured
- - - 22c0 1640,- .
2100
~ %
2 see'. .
- -2000 3 5
C 5 i s><': :n00 g
' 2
. .o t2e'. .
n00 twe' -
sco 0 20 40 60 80 10 0 uo M0 50 183 200 Time (s)
Fig. 22.
Calculated and measured primary pressures.
s - ->
25 Calculated 7- . - Iseosured -
6- .-20
=
w E S-O
" -5
}
3 e- .
u y
8 a s- . f . n f 2-1-
0 .
0 0 20 40 40 80 100 u0 14 0 to ISO 206 Time (s)
Fig. 23.
Calculated and measured pressuricar water levels.
e l
TABLE III INITIAL STEADY-STATE CONDITIONS ,
Parameter TRAC -
Oconee-3 Reactor power (100% power) 2568 MW Cold-leg temperature 563.0 K ..
(554 0F)
Hot-leg temperature 588.9 K l
(600 'F)
Primary-nass flow (each loop) 8820 kg/s (19445 lb/s)
Primary pressure 15.03 MPa (2180 psia)
Steam generator (each loop): 6.29 MPa Secondary Pressure (913 psia)
Main-feedwater flow 680.0 kg/s (1500 lb/s)
Fesdwater temperatu e 510.9 K I
(460 'F)
TABLE IIII MAIN-STEAM-SAFETY AND TURBINE-BYPASS VALVE SETPOINTS The following setpoints are FSAR values and were those used in time TRAC calculation. Actual Oconee-3 setpoints were not available.
- 1. Main-Steam-Safety Valve Setpoints
~ ~
I Pressure (MPa) Pressure (psig)
Open Close Open Close Bank 1 7.34 6.96 1050 995 Bank 2 7.408 7.029 1060 1005 Bank 3 7.512 7.133 1075 1020 Bank 4 7.615 7.236 1090 1035
- 2. Turbine Bypass-Valve Setpoints Pressure (MPa) Pressure (psig)
Open Close Open Close
! 7.067 6.998 1010 1000 m
y--->
aa - . ~ - . . . . . . . . . . .
TABLE XIV .
SEQUENCE OF EVENTS ,
The following times are actual transient times and are the event times used in the TRAC calculation.
.1d>
Event Time (s)
- 1. Turbine trip and reactor trip occurred. O
- 2. Operator assumed manual control of ICS 8
to reduce main feedwater.
- 3. HPI pump A started to assist pump B to maintain pressurizer water level. 30
- 4. Main feedwater pumps tripped. 103
- 5. End of plant data and calculation. 180 Figures 24 and 25 show the hot- and cold-leg temperatures of loop A and loop B, respectively. Temperatures in loop A were slightly lower than in loop B because of a higher main-feedwater flow rate in the loop A steam generator. The initial increase in cold-leg temperatures in the first 10 e of the transient
~
. resulted f rom the sudden reduction in steam generator heat removal caused by turbine stop-valve closure. Thereafter, the reduced reactor-thermal power allowed the cold-leg temperatures to decrease.
Figures 26 and 27 compare actual and calculated secondary pressures for loop A and loop B, respectively. The pressure peaks at about 6 s were also caused by turbine stop-valve closure. TRAC calculated lower peak pressures because of a modeling error in the location of the turbine-bypass valves. They were mistakenly located at the end of the turbine-bypass line rather than at the beginning. As a result, the calculated pressure peaks were lower because of the added volume of the turbine-bypass lines. Later in the transient, after about 40 s, the calculated secondary pressures were higher than the measured pressures. The calculated pressures cycled between the TBV open and close i
setpoints. The actual pressures decreased below the TBV setpoints and remained below the setpoints for the duration of the transient. The reason for this c
- difference is not clear. Possibly the Oconee-3 plant had different TBV I
setpoints and rate characteristics from those modeled in the calculation.
l 1
300 $20 Celevief ed bot-les tempe'eture
-- Calcutof ed cold-leg temperature
- -- - besoured no t= leg t emper of wre sec. -- Woosured cold-leg temperature -
n ~
5 D
- a. b 2 seo- -
2 o .
u0 o ' "
W 570- *
. , o w0 3
/
' s ~ _ --- - - - - .
3 u0 Ss:
0 20 40 60 50 10 0 00 WO 50 23 200 Time (s)
Fig. 24.
Calculated and measured hot- and cold-leg temperatures for loop A. .
~l ce,.uieted ,et-,og t - ereture
conculated cold-les temperature
~~--- u..eu edrne s-seg tempera t ure Soo< ------ Weesur ed cold-Ieg iemporoture -
e '*** m U b 2 m- -
2
- o we 370-I
~
T f 4
/ % MO
~
) " ~ . , , , , $
- ~ - - -
_ ~ _
540 550 0 20 40 60 80 80 0 00 l*O 10 0 100 200' Time (s) l ..
Fig. 25.
Calculated and measured hot- and cold-leg temperatures for loop B.
- ee . - - , ,
G
a .Ge. . .. . .. , .
7ht' ,
Calculat ed , ,q
?be'-, - -teessured -
. #00 7.4-t'-. -
- . epg a o
g v.2 *.. t- . .
E E 2 ,m .e .
- 20 5
- B I
$ s.t.c'. n.
i 4 6 c'-
- -'800 930 6.4 Co .-
8.2 C' . 900 0 20 40 60 80 10 0 20 WO WC 10 0 200 Time (s)
Fig. 26.
Calculated and measured steaa generator A secondary pressures.
1kC' Calculated . ,,
7ht'-. . .... ggeeSWred -
' 400 7,4.g s, ,
y 7.2.g 8 .'i .* *IO T E E 1ho*- .* *EO
- c. .
. =0 :
Seg'."I .-900 4.MD'- - 930 s.2.c' .00 0 20 40 M so 20 20 we EC 20 200 Time (e)
Fig. 27.
Calculated and measured steaa generator B secondary pressures.
e
- , - + - . , y
l Figures 28 and 29 show the secondary-side water levels for steam-generators A and B, respectively. There was excellent . agreement between ,-
calculated and measured water levels in steam generator A, but considerable ,
l difference in steam generator B. Subsequently it was found' that the measured i water level and n'ain-feedwater flow in steam generator B were inconsistent. The steam generator was refilling at a much faster rate than could be accounted for by just the main-feedwater flow. This indicates that auxiliary feedwater may have been inadvettently delivered to steam-generator B. This inconsistency was not mentioned in Ref. 4 nor was auxiliary-feedwater flow data given.
Table IV compares the calculated and maasured minimum pressures, temperatures, and pressurizer water level reached in the transient. Additiesal minimum-value results are ptesented for another TRAC calculation in which it was assumed the measured Oconee-3 thermal power (Fig. 21) to be the decay power.
The Oconee-3 measured thermal power did not include gamma-ray heating.
The results do show the sensitivity of the transient to decay power.
- 4. Summary. The Oconee-3 turbine-trip and steam generatar-overfeed transient of March 14, 1980 was modeled using the TRAC-PF1 code. The
~
calculational results compared very well with the measured data. Differences between calculated and measured results were minor and consistent thr.>ughout the transient. The actual transient had slightly more overcooling then the TRAC calculation. The calculated pressures, loop temperatures, and pressurizer water level were slightly higher than the measured values.
An inconsistency in the measured data given for steam generator B was found. The steam generator was refilling at a much faster rate than could be accounted for by just the main-feedwater flow. This indicates that auxiliary-feedwater flow may have been inadvertently delivered to steam-generator B in the actual transient. If
- it was assumed in the TRAC calculation that auxiliary feedwater was delivered to steam generator B, the calculation would have agreed better with the data. Other factors that could affect the degree of cvercooling are decay power, RPI flow, TBV setpoints, and valve-rate characteristics.
B. Main Steam-Line Break
- 1. Introduction and Summary. For this transient, the overcooling of the primary side of the plant is caused by a severe depressurization of the secondary side. The secondary-side depressurization is caused by a full double-ended steam-line break in one of the steam generators (SG A). The accident sequence begins with the break of a 34-in. steam line coincident with reactor and turbine trip from full reactor power. The main forcing function for the overcooling of the primary side is the delay by the operator in isolating 1
- 1 4 . j
. s,_s 1
- . . -. . x . -. ._. . - - . . .- -
~ - --
~,
+ ~ , ,
[ \ m_. -- .> g.
- 1 1 !
l
. i l
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, j
- A e Calculoted 3a 4 l
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, .~
. 'S s . -
0 .- - 3
, .0 to 40 60 '80 10 0 no WO 50 ISO 200
', s Time (s)
- Fig. 28.
Calculated and measured water levels in steaa generator A.
s ,-
% '% 95
. 'o
. --- c ievio,ed -
3, Moctured
,% g- .
, 25 E :::
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. . 0 0 20 40 80 80 10 0 90 WO to 100 200 TI,me (s) s i I* m '
,s,. _ Fig.~29. ;
Calcul'ated and measured water levels in
. ,. s
. , steam generator.b. + 1
-..u- ,
pg
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. '% , ', ~
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TABLE IV
. COMPARISON OF TRAC AND OCONEE-3 RESULTS
- l. With ANS Decay Power in TRAC Calculation TRAC Oconee-3 Minimum primary pressure 12 T MPa (1881 psia) .,
Minimum pressurizer water level 2.75 m (9.02 ft)
Minimum hot-leg temperature (Loop A) 561.5 K (551 0F)
Minimum cold-leg temperature (Loop A) 560.6 K (553 0F)
- 2. With Reduced Decay Power in TRAC Calculktion A TRAC calculation was performed using the Oconee-3 measured reactor power as a decay power curve input to TRAC. The Oconee-3 reactor power curve did not include gamma-ray heating. These results are presented to show the sensitivity of the transient to decay power.
Minimum primary pressure 12.50 MPa (1813 psia)
Minimum pressurizer water level 2.12 m (6.96 ft)
Minimum hot-leg temperature (Loop a) 557.8 K (544 0F)
Minimum cold-leg temperature (Loop A) 558.0 K (545 0F) .
the MFW and EFW_ to the affected steam generator, coupled with a- delay in throttling the HPI flow and restarting one RCP in each loop following attainment of adequate fluid subcooling in the primary system.
The base case analyzed (Case 1) had all of the ICS, protection, and emergency systems operate as designed. In Case 1, the operator is assumed to l isolate all feedwater to both stesa generators 600 e into the transient and then restores the unaffected steam generator at 900 s. Also, the operator restarts one RCP in each loop af ter attaining 42 K subcooling, and throttles the HPI to maintain 42 12.5 K fluid subcooling.
t
t Three parametric cases were analyzed in addition to the base case. Case 2 , ,
l was identical to the base case except the EFW system did not actuate as designed
) because of a modeling error in the input deck. In Case 3, in addition to the EFW system f ailing to actuate as designed, the RCPs never restarted af ter the l subcooling margin was reached because of input deck errors. In Case 4, the MFW l pump was tripped at 0.5 s and the subcooling monitor for restarting the RCPs was .
moved from the hot leg to the top of the core. Although these parametric cases were not specified by ORNL, they are still useful calculations because they give other possible scenarios that possibly could occur during a MSLB transient.
In terms of downconer fluid temperatures and primary systes repressurization considerations, the MSLB base esse was one of the most severe
]
j of all the specified ORNL transients. A relatively cold downconer liquid temperature of 405 K and hence a small margin against the NDT limit was calculated for the base-case MSLB transient. Repressurization of the primary system to the PORV secpoint (~16.9 MPa) was also calculated for the base case.
- 2. Model Description. The TRAC-PF1 input model for the Oconee-1 plant is described in Section II of this report, and the primary and secondary noding 1 diagrams are shown in Figs.1 and 3 respectively. For the MSLB calculations, the steam-line break was modeled in the SG A steam-line shown in Fig. 3 (component 68). The TSV (component 42) was fixed cpen, and all of the steam from SG A passed through the TSV to atmospheric pressure. In the unaffected steam-line (SG B) the TSV was closed, and the TBV system' operated as designed.
All of the other systems also operated as designed except for the parametric cases. The significant features and initial conditions for the MSLB calculations were:
- 1. Full reactor power.
- 2. Nominal temperatures and pressures in primary /aecondary.
- 3. Decay heat - 1.0 times ANS standard.
- 4. Reactor and turbine trips coincident with MSLS.
- 5. Operator fails to isolate feedvater to both SGs until 600 s. ,
. 7. RCPs restarted after 42 K subcooling reached.
- 8. HPI throttled to maintain 42 i 12.5 K subcooling.
O e
e
- - - - - -c., , , -
- 3. Results.
- a. Base Case (Case 1). Key events calculated during the transient for Case 1 are presented in Table XVI. The calculation was run long enough (7200 s) to determine whether or not the operators could recover the plant following the steam-line break. Figures 30 through '66 show plots of key system parameters calculated for the transient. Both short (0-900 s) and long (0-7200 s) .-
time-scale plots are presented to give a complete description of the system thermal-hydraulics.
l The transient was initiated by fixing the TSV in loop A open and modeling
, a break in the steam line downstream from the TSV (component 68 - Fig. 3). The TSV in loop B was closed at transient initiation and terminated condenser feed i from the turbine. The turbine and reactor were then tripped followed by a feedwater - heater flow / drain trip. At ~5 s the TBV in loop B opened af ter the TBV setpoint (7.06 MPa) was reached. The TBV in loop B continued to open and close until ~40 s, af ter which time the TBV remained closed until late in the transient (~5462 s). HPI initiation occurred at ~21 s af ter the primary system pressure had decreased to 10.44 MPa. At ~29 s the ICS detected a low-level limit in SG A and the EFW pump was started. The EW valve to SG A was opened, and EW flow was initiated. At 47 s the MFW pump tripped on low-suction pressure. The ICS detected a low water level in SG B at 48 s, and EFW flow was initiated af ter the EFW valve in loop B opened. At ~51 s, the RCPs were tripped (30 s af ter HPI initiation), and the feedwater realignment trip occurred (all feedvater directed through the EFW header). At ~53 s, the condensate-booster pump tripped on low-suction pressure. The SG B water level reached 50%
(operating range) at ~346 s, and the EFW valve to SG B was shut. All of the EFW flow was then directed to SG A. The 42 K liquid subcooling margin in all primary system loops was reached at ~526 s, and HPI was throttled. Also, RCPs Al and B1 were restarted at this time because the subcooling margin was sufficient. At ~530 s the primary system had depressurized to the accumulator tank setpoint (4.17 MPa), and the check valves downstream from the accumulators opened. At ~538 s the accumulator check valves closed. Per the transient specifications 3, the steam generator secondary sides were isolated at 600 s.
Because of reverse heat transfer in SG B (heat transfer from secondary to primary side), condensation caused the water level to increase to the high-level limit (90% operating range) at ~655 s. At 900 s, SG B was restored to allow the EFW and TBV systems to operate if needed. However, because the SG B level was at the 90% limit and the secondary pressure was low, these systems did not actuate until auch later in the transient. Following isolation of the steaa e b
- s---m ev- w --w---w----ww, - - - -
{ '
TABLE IVI SEQUENCE OF EVENTS ,
l EVENT TIME (s)
- 1. MSLB - loop A steam line 0.0 , , ,
- 2. Turbine had reactor trip; TSV 0.5 loop B closes
- 3. TBV loop 3 opens (setpoint 7.063 MPa) 5.0
- 4. HPI initiation (setpoint 10.44 MPa) 21.2
- 6. TBV loop B closes 39.9
- 7. MW pump trip on low 47.8 suction pressure
- 8. SG B low-level limit reached; 48.7 loop B E W flow initiated
- 10. Condensate booster pump trip (Iow 53.9 )
suction prassure) l
- 11. SG B level at 50%; loop B 346.7 E W Valve closed
- 13. Loop A, B accumulator setpoints 530.9 reached (setpoint 4.17 MPa)
- 14. Loop A, B accumulators off 537.9
- 15. SG A, B isolated; EW pump and 600.0 hoewell puno tripped off
- 16. SG B restored 900.0
- 17. PORV setpoint reached (setpoint 4678-7200.0 16.9 MPa) - PORV opens and closes for remainder of calculation
closes for remainder of calculation to maintain setpoint pressure
- 19. SG B level drops below 50% 6121-7200.0 operating range; E W initiated -
EW pump os/off for remainder of calculation to maintain level at 50%
- 20. , Calculation terminatM 7200.0 -
generators, the primary system began to repressurize. At ~4678 s, the PORV setpoint was reached (16.9 MPa), and the PORY cycled opened and closed for the remainder of the transient. The secondary side of SG B repressurized to the TBV setpoint (7.06 MPa) at ~5462 s, and the TBV opened and closed for the remainder of the transient to maintain the setpoint pressure. At ~6121 s, the SC B w' ster level dropped below 50% (operating range), and the EW system was activated to
maintain the level at 50%. The calculation was terminated at 7200 s, and the primary system was full of liquid. At this time, the decay power produced in the core was being removed through the unaffected steam generator (SG B).
The pressurizer pressure history is shown in Figs. 30 and 31. Initially, the primary-system pressure decreased rapidly because of the rapid .-
secondary-side blowdown in SG J. following the MSLB. The depressurization was terminated by ~100 s when natural circulation flows were established following pump coastdown and RCP trip at ~51 s. The primary system then began to repressurize slightly until the RCPs were restarted at ~526 s. The enhanced heat transfer through the steam generators, condensation of the steam in the loop B candy cane, and throttling of the HPI af ter 'the RCPs were restarted caused the primary-system pressure to decrease to the minimum value for the transient ( ~3. 5 MPa) . After the steam generators were isolated at 600 s, the primary system repressurire. i to the PORV setpoint (16.9 MPa) at ~4678 s. The PORV then cycled for tLe remainder of the transient to maintain the primary-system pressure at or below the PORV setpoint.
The pressurizer water level is shown in Figs. 32 and 33. The pressurizer completely emptied by ~40 , because of liquid contraction resulting from the severe overcooling of the primary system. Aftzt the HPI had been on for some time, the pressurizer began to slowly refill until the RCPs were restarted, and the HPI was throttled at ~526 s. Again, the resulting overcooling of the primary system caused the liquid to contract further, thus the pressurizer again emptied. Af ter the steam generators were isolated at 600 s, the primary eystem liquid expanded because of fluid heat-up and the pressurizer slowly refilled.
Downcomer liquid temperatures for the base case are presented in Figs. 34 and 35 at the top axial downcomer level just below the cold-leg inlet nozzles.
Because of the severe overcooling in the affected loop (loop A), asymmetrical liquid temperatures are calculated in the vessel downconer. Tha fluid temperatures in the downconer cells associated with the loop-A cold legs were l
calculated to be ~20 K r. older than the cells on the loop-3 side (Fig. 34). The minimum downconer fluid temperature calculated was ~405 K at ~526 s when the RCPs were restarted. While the RCrs were tripped off, the downconer fluid temperatures were affected by the vent-valve flow shown in Fig. 36. The warmer upper plenum fluid mixed with the colder downconer fluid during the time the RCPs were tripped. During the time the RCPs were operating, the vent-valves did not open because the pressure gradient was reversed.
- 4e v' + - " - ' '
-e w e
se . , , . . . . .
me- -
sose
,,, . veo
.goe -
uso Iee.. .
moo es. -
7so
- n. -
-sos se . . , , , . . .
e as ano soo ao soo eso soo soo sw MM Fig. 30.
Pressurizer pressure (0-900 s) - base case.
l m , , , , , , ,
. amo es. .
asse me" . -
acoe go. - ano go. -
.soo aso on. - -
moo es. -
V ~
soo l 2 .
l e soo zooo sooo moo soon esos vooo sees l MH Fig. 31.
Pressurizer pressure (0-7200 s) - base case.
- e. , , , , , , , ,
s- -
e a- .
a' . a g .
a- -
i
- s i.
e- - - -e
-t . . . . . . . .
o so ano ano ano see ese no ano see w W _.
Fig. 32.
Pressurizer water level (0-900 s) - base case.
u . . , , ,
a- ~
- e. .
- se 1 e- .
g t
N e- - . = l l
- a. .
- m .
3 .
- e. .
-a . .
am e moo sono anos sono seen mee snee MW Fig. 33.
Pressurizer water level (0 7200 s) - base case.
I 1
ese , , , , . . .
R1HZ ~ '
"] e226 216
- 236 E .. *246 .
i =256 see. *266 .
m- . _
m.
- m. .
375 , , , , , ,
aos neo aoo soo eso m soo ese e no MW Fig. 34.
Downconer liquid temperatures (0-900 s) at vessel axial level 6 (all azimuthal sectors) - base case.
m . . . . .
I
! eso. < 5 ; .1- _ _ _ e- . .
E eso. f R TH Z .
216 226 .
Ieso.ano.
- 236
=256 246 -
1 doo- ,
266 -
- m. .
f
- m. .
doo , , , , , .
e aos sooo sooo sooo sono sooo me esso MW Fig. 35.
Downconer liquid temperatures (0-7200 s) at vessel axial level 6 (all azimuthal sectors) - base case. .
a w
4 l
l l . - - . _ _ - _ _ - . . _ - - - . . _ _ _ _
The hot-leg liquid subcooling in each hot leg is shown in Fige.' 37 and 38. ,
Figures 39 and 40 show the hot-leg mass flows sad Fig. 41 shows the candy-cane void fractions. The subcooling margin calculated in loop A is significantly larger than in loop B for much of the transient because of the MSLB in loop A and the resulting enhanced heat transfer in the affected steam generator. The enhanced heat transfer caused higher natural-circulation flows in loop A - -
compared to loop B (Figs. 39 and 40) from ~150 a to the time the RCPs (Al and
- 31) were restarted (~526 s). The 42-K subcooling margin was reached in loop A at m175 s, but the RCPs could not be restarted until this margin was reached in all loops. The flow in loop B stagnated at ~150 s because the candy-cane in this loop reached saturation and voided (Fig. 41). Tne subcooling margin in loop B was not reached until ~526 s at which time the RCPs were restarted.
Af ter the RCPs were restarted, the void in the loop-B candy cane condensed and was swept out (Fig. 41) and the subcooling margin in both loops equalized. It should be noted that at the time the subcooling margin was reached in all loops
(~526 s) and the RCPs were restarted, the HPI was also throttled. The i subcooling margin never decreased below 42 K for the remainder of the transient, thus the HPI was never turned back on nor were the RCPs tripped again.
, As discussed in the preceding paragraph, the loop-B candy cane voided at
~150 s (Fig. 41) and the void was swept out af ter the RCPs were restarted. Even i
though the candy cane in loop B voided, which is the highest point in the primary system, the vessel did not void as shown in Fig. 42. Figure 42 shows the volume fraction of the upper plenum liquid, which is the region above the reactor core. From Fig. 42 it is seen that the vessel remained completely full of liquid for the entire transient.
Figures 43 through 46 show mass flows and liquid temperatures in the cold legs. Figures 43 and 45 show the cold-leg mass flows for loop A and loop B respectively. As discussed previously, before the RCPs were restarted significant natural-circulation flows were calculated in loop A because of enhanced heat transfer in the af fected steam generator. The loop-B flows prior to RCP restart were essentially stagnant as shown in Fig. 45 because the loop-B candy cane voided. When the RCPs (A1, B1) were restarted at ~526 s, essentially i steady-state flows were calculated in cold legs Al and B1, but the flows in cold legs A2 and B2 reversed (flow out from vessel) as shown in Figs. 43 and 45. For
( the remainder of the transient, the cold-leg flow directions and magnitudes l
remained essentially the same as shown in Figs. 43 and 45 af ter ~526 s. "The cold-leg fluid temperatures are shown in Figs. 44 and 46 for loops A and B l i !
l i .
T m -mwy,- - - g m. - - - w- - - - - - - - - w y c. --_ _ .
me . . . . . . . .
i
- m. -
.e - -
3es.
- 1. .
g .e. .
3 go. -
es-
- g. M
-40 e eso suo ano me soo eso s aos ese M(4 Fig. 36.
Total vent-valve flow into downcomer - base case.
as , , , . . . .
.se 1,00P A (5012) toor a psH) es-I es.
,,,,....=r.- -
se
- ij
! / ..
- a. ,.- - se ee
- m. \
.se i .\: l -.\::
1
~~
- e. .: -
. .e
. as
-as . . . . . . . .
e so son aos me neo eso no eso see MW Fig. 37.
Hot-leg liquid subcooling (0-900 s) - base case.
l G
t i
I
,. Woe , , , , , , , ,
. (peAd)LDOP A anoes Sees- - .%' - ' , -
3B000 a00,- 1 -
t
- 3000 t
I egeD- ; -
. . gge E 1 ages- t -
- t soso t
t t
Ne9- . 1, -
.ngo
. I
- e. .
.,,=.............-.I . 0 u
=300s l , . , , . , , ,
e no soo aos ao soo eso no soo eso NW Fig. 38.
Hot-leg liquid subcooling (0-7200 s) - base case.
so . , ,
gg. .
j LOOP A($000) . .ago LDOP e ps0 es.. .
. ,s
- n. -
g -
.. =
M ns j
Ise.
so- -
j me e- 1 l -
Q 1 i n T so- ! -
h f so -
so- : -
bI
.....; . .=
l
- e. .' . -o
-m . . , , . . .
e me zooo aooo moo sooo soon moo essa MW Fig. 39.
Hot-leg mass flows (0-900 s) - base case.
l l
E
anno . . . . .
WA 3eoso .
.oso. k=400P e .
soooo noso- -
-soon esso. . .
1 ~
1 l aseo. .
oooo l
l soou- ,
..oog
.L o.6... . o
?
-sm '
o eso sooo sooo aooo sooo sooo mo sooo MW Fig. 40. j Hot-leg mass flows (0-7200 s) - base case. 1
. u' , , , , , ,
$eed)LDOP A
- i. . ... 9=409P..e .
as- : : .
Ies- ! i .
ea- i j .
sa- : : .
i i
-e t . . . . . . . .
e me a, , aos ano soo oso m oso soo MW Fig. 41.
Candy-cane void fractions - base case.
9
u , , , , , , , ,
1 Ias-as- .
i se- .
I.ma-S
- e. .
-42 , , , , , , , ,
0 ' 50 20 age des See see m 800 000 MW Fig. 42.
Upper plenus liquid volume fraction - base case.
l i
msg - , . ,
eges- AI
~
W A2 -ages 8000- .
4000-* . 8000 M= .
.em I
~
3ges aOS- .
e- -
t . ..
i ! 4 i
-me. - i e = i: L
.! ,' ,!l' .!
.! --sese 3 33- i.1, ... ... ! s ... , ! s, . !.
333g ""8888 8 4 s s s s s m s m MW Fig. 43.
Loop-A cold-leg mass ficus - base case.
l i
---,c - ,-.r---.- w -
l l
srs , , . . . . .
. sao (psA4LDOP As ,
see. (m=Ninor A2 -
see sus- . . se 8 E
- - #3 a **
Isee- ..
as-ase see
. r., age as. .
. .aee ago, -
- ase ars , , , . .
m e se zoo aos ao soo eso see too MW Fig. 44.
,, Loop-A cold-leg liquid temperatures - base case.
, me. . . . . .
goog- r .
(sam @LDOP B2
. so enee- -
asse-- . sees ages- .
o .
tees 5.,
M- -
g
.aoas gas- .
- g. ..*.... . .
e
- 5 :;
-nee- i it i ; :~
i . ; t. :. _.j
!i i
- .:i ! -some
_ ,,,,_ t.-. .
-seos , , , , , , ,
m o me noe age me son ese 800 see MW Fig. 45.
Loop-B cold-leg mass flows - base case.
6
,--r--,---.a,- .-e ,--e.,,, -,..- e e,-.-
'a respectively. It is seen that the loop-A fluid temperatures were significently ,
colder than loop B during the time the RCPs were tripped (~50-526 s).
Detailed results of key system parameters in the steam generator are shown in Figs. 47 through 54. Figures 47 through 50 show det ails for the affected steam generator (SG A), and Figs. 51 through 54 show deedis for the unaffected steam generator (SG B). Figure 47 shows the secondary-side water mass in SG A, '~
and Fig. 48 shows the secondary-side pressure history. The secondary side depressurized to essentially atmospheric pressure by -85 s, and this time corresponded to the minimum water mass inventory. Figure 49 shows the resulting flow out the broken steam line during the course of the transient. After the secondary side of SG A had depressuri:ed sufficiently (~100 s), the EFW penetration increased, and the water inventory began increasing (Fig. 47) as less EFW was bypassed out the broken steam line.
Figure 50 presents some detailed plots to help explain the EFW penetration phenomena. The top plot in Fig. 50 compares the RAC-calculated vapor velocity at the EW injection point to the complete flooding curve predicted by the Wallis-Kutateladze correlation (K = 3.2) for various pressures in the SG A l secondary side. This plot shows that EFW penetration will not occur until the vapor velocity is less than 4 m/s, and this velocity is not reached in the TRAC calculation until the secondary side had depressurized to 4.5 MFa. The botton
- plot in Fig. 50 gives the TRAC calculated liquid-vapor velocity correlation at the EFW injection point location. This plot shows that EFW penetration as calculated by TRAC did not occur until the vapor velocity decreased to ~7.5 m/s, which closely agrees with the Wallis-Kutateladze correlation.
Figure 47 shows that the secondary-side water inventory remained
- relatively constant (~1000 kg) af ter EFW penetration occurred at ~85 s and did not change significantly until 450 s. As will be discussed later, the reason for the increasing inventory after ~350 s was because the EFW to SG B was terminated and all EFW was directed to SG A. The inventory decreased following RCP restart at ~526 s because of the enhanced heat transfer from forced convection on the primary side. Then, at 600 s the steam generators were isolated, and EFW was terminated. The water inventory in SG A then dec. eased to zero and remained empty for the remainder of the transient.
1 Figures 51 through 54 show some key parameter plots for SG B. Figures 51 and 52 show the secondary-side water inventory esiculated for SG B. At 48 s, the low-level limit was reached following TBV operation in SG B, and EFW was '
initiated. The water level continued to increase to the 50% operating range at which time the EFW valve to loop B was shut (~350 s). The SG B level remained 1
- u, ..- - -
see , , , , . . .
oso
, - s
-emem W sr oso sos-
~ m a =- '
E me
- m. - .
Isee-age.
\ :.
-me t :. .' . ,
1 . .
ase- -
- !t i m
t
- , *....,, aos en-
. -No
- O e N k S 5 O O *
- MW Fig. 46.
Loop-B cold-leg liquid temperatures - base case.
wooo , , . . . . .
. -Nm
- m. ~
, -anoos asse. -
-soooo 3 .so.. -
-.ooo g
esse-anoo-. #ee
.. . -e
. 4eso i
.soon
\
e O S S S S O M S
- MW yig. 47.
SC A secondcry-side water inventory - base case.
t l
~62-M , , , , , , ,
oo. .
.m '
so- .
I q. . T no Iso. so. . sos m.- . mo o..
-o
-m . . , , . ,
o no ao zoo soo soo eso no aos soo MW Fig. 48.
SG A secondary-side pressure - base case.
mos , , , , ,
moo. ., seo 1No. . 23oo moo. . 23Bo
.ao. .
. .so I I
soo- -
- tano I
! me. -
. soo ,
t i soo. -
. .mo
- e. -
o
-soo . . , ,
o no i son aoo ao soo eso no soo soo l
MW Fig. 49.
SG A steam-line flow - base case.
l l
4
.m. -
4
,v v
- 1. 7 .
l i
av I l
Ea>
5 T *h"' ,
E 86 l
.I - l Ano- mms-avamas '
g /***OT
- s s-4 e a s k 4 e- Pr==ur M R e-a 3
5*
a o s,,,,,g,,,,7(,,,,3a Fig. 50.
CCFL phenomena in affected steam generator (SG A) - base case.
senso ,
m5000
-90000 agege-~
~
. mas 9*
I y asese-g 3 .
3 m
3000- 3 eses
(
( .
i goes.
soos 3000 m
soo eso e me zoo aos aos soo aos inE W t
(
Fig. 51.
' SG B secondary-side water inventory (0-900 s) - base case.
i e, ,. g
soons , , , , ,
ease. . ,
- 90000 noee. .
mese .
Q 3 000-g . .
3 . .sosoo 3 m- -
messe.- . deseo so . .
3asse
. .aoo.
Sees , , , , ,
o woo sooo sooo sooo sooo sooo 7eco sooo
.. MN Fig. 52.
SG B secondary-side water inventory (0-7200 s) - base case.
se , , , , , , , ,
. .aos oe. .
7e- .
es. .
I SS-
.soo e-
~
- **o .
as- -
. .ao me.
-seo e- -
e , , , , , , , , e o so soo ano 4,u soo ese zoo eso see M(s)
Fig. 53.
SG B secondary-side pressure (0-900 s) -
base case.
l 9 *
. .~. .. -.
l essentially constant at the 50% level until the RCPs were restarted at ~526 s. ,
J After the RCPs were restarted, the SC B primary-side liquid temperatures decreased below the secondary-side temperatures and secondary-to primary heat transfer occurred. The steam in the top regions of the secondary side began to condense on the SG tubes, and the liquid level began to rise (Fig. 51). ne
The secondary-side liquid level rose until the SGs were isolated at 600 s.
level then remained at approximately 90% of the operating range until SG B was restored at 900 s. After 900 s, the water inventory slowly decreased (Fig. 52) because all of the primary-system energy removal occurred through SG B af ter At -6121 s the SG B secondary-side this time (no EW to SG A af ter 600 s).
level decreased to 50% of the operating range, and EW was initiated. For the remainder of the transient, the EW was controlled to maintain the 50%
cperating-range level.
The SG B secondary-side pressure history is shown in Figs. 53 and 54.
Initially the pressure was controlled at the TBV setpoint (7.063 MPa) following closure of the TSV in the loop-B steam line. The pressure then began to decrease af ter EW was initiated following the low-level limit trip at ~50 s.
The pressure continued to decrease because of cotidensation from the cold EFW until the 50% operating-range level was reached at ~350 s and EFW was terminated. he pressure remained essentially constant at ~3.5 MPa until the RCPs were restarted at ~526 s. After this time, as discussed previously, condensation of the steam in the secondary side on the SG tubes because of secondary-to primary heat transfer esused the pressure to decrease further to
~L.5 MPa. The pressure remained at ~1.5 MPa until ~900 s.
After SC B was restored at 900 s, the secondary-side pressure increased slowly (Fig. 54) until the TBV setpoint was reached at ~5462 s. The TBV system maintained the TBV setpoint pressure for the remainder of the transient.
Figures 55 through 61 show some details in other important components in '
the secondary side. The MFW pump speed is shown in Fig. 55 and the MFW liquid l
temperatures are shown in Fig. 51. The MFW pump tripped on low-suction pressure at -48 s. The loop-A MW liquid temperature (Fig. 56) essentially followed the saturati'on temperature corresponding to the SC A secondary-side pressure; thus this temperature was much cooler than for SG B. Figure 57 shows the MFW mass flows into each SG. It is seen that the MFW flow to SG A was much higher than to SG B because of the lower back pressure in SG A. The EFW flows to each SG are shown in Fig. 58. EFW was started at ~30 s to SG A and at ~50 s to SG B.
Again, because of the low secordary pressure in SG A, more EFW flow was directed to that steam generator. At ~350 s, the EFW valve to SG B was shat (50% level
l l
1 8e . , , , , ,
~
mes
- n. .
I I -
I ..
1 ..
..o.
I so- . 80 so.- . soo se , , , , , , ,
-me e isoo sooo anos moo sooo sooo me sooo
~
M (s)
Fig. 54.
SG B secondary-side pressure (0-7200 s) -
base case.
eee , , , , ,
.so SOS- .
I. '
I see< ' #8'
~
l I. .
. ..e i me- -
. -mee l e . . . .
e . se me se =
MW Fig. 55.
MFW pump speed - base case.
9,*
as - . .
. . me Smagoor s as one.
as me- .
.m, E m
me.
no Y I me.. .
me "388 go. -
us me . . . . .
ao we e se m es se me MW Fig. 56.
i MFW liquid temperature - base case.
l mes .
(sen$oor A . ages ses. .
gam *cor a see.
. . ,se m- .
2 . .
eso. ago j -
}
ses- .
.: mes me- : -
.i no ses- i -
- see ses- :
. ase
-i
- k...****~....~....{ . . . . . - -e
- e. .
-me . . .
as e se a se se me me MW Fig. 57.
MFW mass flows - base case.
we . . . . . . . .
( ad)r= r ase se- ( LOOPO -
see go. -
.mo Q.
4 ee. -
"5 no 4 es- ... : -
- . .no i
- g. ; -
g se *
/ .
t :
e :
so- :. -
e
. .(
g.J 'd 's............t o
-as ~ ~d e no zoo ano ao soo eso no eso ese M (a)
Fig. 58.
EW mass flows (0-900 s) - base case.
I l
wo . . . .
L (seAd) LDOP A No Se- M LDOP O -
3.
as- -
ase ee. .
Q me E D m- s so l m- s: .
se so-.! 3 jii g. -
! l!! !5 ~
- . :i : . :: -
.as -
-m e noe anos sooo moo sono sooo moo esse MW Fig. 59.
EW mass flows (0-7200 s) - base case.
os a O
m __ a- ______- _________-_ _____m__
w ., . - .. . . . . . .
se8ee asu . . . . . . . .
aess- __
- eeJK 333 E
. ,g,,
Ises4-ses.a-eesse M-aets-
. Am4ae see. a eE#8 s0La . . .
m e**
e es see see aos eso ese ese MW Fig. 60.
EW liquid temperature at pump discharge
- base case.
J ese . . . . . . . .
SOLS -LDOP A Dash - LDOP e
..J'*t ese 1
see. :
4
- ..se I.
}. .
l ..
en- I* -
j ase .
1 .
.ase
. 1 .! -
ses- 1 '!
i t : **
e see e sie sie see aie sie sie m sie ese MW Fig. 61.
l at injection
' EW liquid temperatures locations - base case.
in SG B reached), and all of the EW was directed to SG A. At 600 m the SGs ,
were isolated, and the EW pump was tripped off. Figure 59 shows that at
-6121 s , EW was initiated again to SG B af ter the level had dropped below 50%
of the operating range. The EW temperature at the pump discharge is shown in ~
Fig. 60, and the E W temperatures at each injection point are shown in Fig. 61.
Other key system parameters are shown in Figs 62 through 66. The HPI ~~
flows are shown in Figs. 62 and 63 for loops A and B respectively. HPI was initiated at ~21 s and was terminated at ~526 s af ter the subcooling margin in the primary system was reached. Figures 64 and 65 show the accumulator water levels and fluid volumes discharged into the primary system. The accumulator pressure setpoint of 4.17 MPa was reached at ~530 s and the accumulators operated for -8 s. Approximately 2 m3 of secumulator liquid was discharged into the primary system. Figure 66 shows the FORV mass-flow history. The PORV setpoint of 16.9 MPa was not reached until ~4678 s, and the PORV then cycled for the remainder of the transient to maintain the primary-system pressure at or below the setpoint.
- b. Parametric case - Case 2. This case was identical to the base etse (Case 1) except the EW system did not actuate as intended because of input modeling errors. All of the other systems functioned as designed in Case 2 including the subcooling monitoring system. The sequence of events calculated for Case 2 is given in Table XVII. The events that occurred during the first 54 s were approximately the same as those calculated for the base case (except that the EW system did not actuate). At -400 s, the RCPs were restarted after the 42 K subcooling margin was reached. Also at this time, the HPI was throttled. At 600 s the SGs were isolated, and at 900 s the calculation was terminated.
The primary-system pressure is shown in Fig. 67. The pressurizer water level is shown in Fig. 68. The minimum pressure calculated was ~5.0 MPa and occurred at ~175 s. Af ter natural-circulation flows were established (~150 s),
the primary system began to repressurize (Fig. 68), and the repressurization continued for the remainder of the transient. The slope of the pressure curve changed af ter the RCPs were restarted at ~400 s because of forced circulation and enhanced heat transfer through the unaffected steam generator (SG B). The pressurizer completely emptied by ~40 s and began to refill af ter the primary system began to heat up because of fluid expansion. After the HPI was throttled at -400 s, the pressurizer water level rise was terminated (Fig. 68). After the SGs were isolated at 600 s, the pressurizer water level began to increase again because of fluid expansion. :
. l i
l
~: ,
1
. l I
ass , , , , , , ,
so- (pend) 611. cop A2 -
ob. -
- g. -
g h
.ao 4 .
~
s- .
gg. .
e- -
o
-a.s , , , , , , . ,
o so zoo aos eso neo eso no aos soo MM Fig. 62..
Loop-A HPI flows - base case.
vs , , , , , ,
faer '"
M LDOP E2
-so
. . s.s - .
-5 4
m- . gr g
.so 12- -
l
.s I
s-. -
1s- .
e-- o
-a.s
--s e no ano ano me soo - oso no eso soo l
MW l Fig. 63.
Loop-B HPI flows - base case.
,ye
/
s 3
4 .-
-g*
Saso , , , , , , , asei LOOP A(50LD) !
4mo- LDOP e (tug .
.smo asoo- .
I am- .
- s.amsev Aygg- .
8 arss-s.ses s 31 rsoo- ,
- . amore t.............
4 AFB ' -
4.aos , , , , , , , , suseo o soo ano ano aco soo soo too soo soo MW Fig. 64.
Accumulator water levels - base case. !
u , , , , , , ,
~
LDOP A W) g- LDOP 3 W , , , , , , , , , , , , , , , , , , , , , , , , , , , , , , , , ,
r
- g u-
.se r s
Iu- -
m u- .
sa- -
o . o
-42 , . . , , . , ,
o no aos ano aos soo eso veo moo soo MW Fig. 65.
Accuriulator liquid volume discharged - base case.
1 w+
1 i
1 I
i
, m , , . - , , , ,
. .s 3- .
n ,
as- .
.ao 3 .. - . .. g h s- .
2 -
.m e- .
- i ;
l s
- s. .
r .
e- hahI L . .o
-e , , . . . ,
e one sooo sooo sooo seen sooo noo sees TihC @
Fig. 66.
PORV mass flow - base case.
so , , , , ,
- 2300 t i.o _
g j -.a i, f mos d- .
5 -=
l l- .
g
$ ~~ .
O 900 200 400 400 350 000 MD 400 000 M (9 Fig. 6'I.
Pressuriser pressure - Case 2.
~ ,q
~
4 m
S 4 S l
a , -i ,
% g
.r-
TABI.E IVII l i
MSI.58 (Case 2) i SEQUENCE OF EVENTS .
EVENT TIME (s) ;
1-14. Approximately some as base case 0-53.9
- 16. SC A B isolated 600.0
- 17. Calculation terminated 900.0 aE W pump never started
, Figure 69 shows the downconer fluid temperatures near the cold-leg connections in the vessel. The minimum downconer fluid temperature calculated i was -475 K and occurred when the RCPs were restarted at ~400 s. Asymmetrical fluid temperatures in the downconer were calculated in case 2 sis.ilar to the base case with the colder temparatures calculated for the loop-A side of the
- downconer.
Figure 70 shows the hot-leg asas flows for case 2 and Fig. 71 shows the candy-cane void fractions. Natural circulation flows were calculated after the RCPs coasted down at ~150 s. The natural-circulation flow in loop A was such higher than in loop B because of the enhanced heat transfer through the affected steam generator. At -400 s the RCPs (A1, B1) were restarted and forced circulation through the primary system was calculated. Figure 71 shows that the loop-B candy cane voided soon after loss of forced circulation in loop B
( ~150 s) . The candy cane remained voided until the RCPs were restarted. The loop-A candy cane voided for a brief period af ter the 'SCs were isolated at
-400 s (Fig. 71).
The cold-leg loop flows for loops A and B are shown in Fige. 72 sad 73 respectively. These flows were similar to those calculated for the base case, except that the loop-A flows were somewhat higher in the base case. The cold-les temperatures in each cold leg are shown in Figs. 74 and 75. The minimum liquid temperature calculated in loop A was 475 K, and the minimus j temperature in loop B was -400 K. It is interesting to note that in case 2 the ga ,
mm-r- , v- - , , - - , - - , , n
O
. . . . . . . . . . . . . . ~ . . ._.- _
i 4 , , , , , ,
1 s-s 1
j .. .
, I g h a-
- s j t- -
I
... ..o I
-t . . .
4 1
o no m aoo .co soo eso m soo soo i
TM M l Fig. 68.
Pressurizer water level - Case 2.
i i
i
- sao , ,
sn
~
l t RNZ j ,
wo2 216 -
- 226
.236 se g uo- ,1ss -
256 "
! 266 98o- ..g
.a soo. .
. m t
! 40s- -
m 4Ao . . . . .
aos aos eso nao eso m eso see e so MW Fig. 69.
Downconer liquid temperatures at vessel r
axial level 6 (all asinuthal sect. ors) - .
! Case 2.
t t
usee , , , , , , ,
$ seed)LDOP A 3sese
.o (8:eMcP e j _
{ sneee esco- -
eene . , .
1 1
e .
, r.....
I em l .. 1
\
sm- \ -. 00
\ . _. 1
- e. %, , ..
8 t.*
.aeoo . , . ,
-**e .
o no ano me eso soo een no soo coe M (s)
Fig. 70.
Hot-leg mass flows - Case 2.
r u , , ,
(noEdADOP A i-
/
r F'~ *=., 8 I ?
u- l -
1
/
Iu- I t
i i
o2-t f -
i l :
e.; - -
l l
42 . . , ,
e me ano see me neo see me see see M (s)
Fig. 71.
Candy-cane void fractions - Case 2.
i l
l 4
e
I Neo ' ' '
sees gggg, M Al- .
W A2 eseo.
.-eese .
esse.
asse-
,,,,, f g_ .
.aeee 5
seee-e.. .
.e
-mee. I au;,.,ug.
u!: lr _.
a:f -
uu.,
! l { h.,3 ..a e
=aceo . ,
no o no ae soo aeo eso eso eso see w (s)
Fig. 72.
Loop-A cold-les mass flows - Case 2.
neo . . . .
(eead)Loor es (dom-)tcoou ,
se,..
-'8888 does-R :ir 2 m- R
.eese .
hm e .
aeoa f
j m. -
- o. ..e k $' . .i !
l!-
-moo -
.a.oo-i*l $ li!
i.
- s i :i t
- J i --*
.t9 3..
""o se see aeo aoo ee. eso me eso see M (s)
Fig. 73.
Loop-B cold-leg mass flows - Case 2.
1 l f i
l l
l
m . . . . . , , ,
see geo- (paid)LDOP At
~
W)LD0P A2 se eno- .
l .nse g neo- l. .
see E . .
Igio.so- .
ee i so- .
ese- ..
me- .
ese- .
no e.
e e me ase ae. noe m m M (s)
Fig. 74.
Loop-A cold-leg liquid temperatures - Case 2.
"' (saad)1D0P B1 (seews s2 -~"
l see- .- ee.
E es - -" E
.-ee lsee-l \
on- ass .
\ .
S i l M es-. i- .
ase 45- .'38
- e. .
se m -M.
e 4 a a 4 a k m s ,
M (e)
Fig. 75.
Loop-B cold-leg liquid temperatures - Case 2.
0 0 4
s , ,. _ ._
loop-B cold-leg temperatures were colder than loop A; however this was not true in the base case. The reason for this is because the EFW system was operational in the base case and caused the loop-A cold-leg temperatures to be less than loop B.
Details of key stean-generator parameters are shown in Fiss. 76 through .
- 80. Figure 76 shows the secondary-side water inventory for EC A. Because the EFW systes did not work and because the MFW pump tripped at -40 s, the inventory in SG A was depleted by ~100 s and never recovered. The associated secondary-side pressure in SG A is shown in Fig. 77. By -45 s, SC A had depressurized to essentially atmospheric pressure. The steam-line flow out the broken steam line is shown in Fig. 78. Figure 79 shous the SC B secondary-side water inventory, and Fig. 80 shows the secondary-side pressure in SC B. Because of secondary-to primary side heat transfer in SG B, condensation of the steam caused the SG B water inventory to increase and the pressure to decrease. The condensation effect increased af ter the RCPs were restarted at 400 s because the heat-transfer rate from the secondary to the primary increased.
i
! Other systes parameters are shown in Figs. 81 and 82. The MFW pump was tripped at ~50 s en low-suction presscrs, and the feedveter flow decayed to zero by ~150 s. Figure 82 shcws the HPI flows into each cold leg. At 400 a the hPI ,
- ras throttled after adequate subcooling was reached in the Icops.
- c. Parametric case - Case 3. Parametric Case 3 was identical to Carr. 2 exc.ept that the RCPs did not restart as intended because of input errors.
Another significant difference between Case 3 and Case 2 is that the MF4 pump did not trip until ~330 e in Case J because of errors in the ICS modeling.
Because the MFW pump did not trip until late in the transient, this calculation can be considered as a MSLB with run-away main feedwater. The sequence of events calculated for Case 3 is giver. in Table XVIII. The events that occurred l
during the first 54 s were apgoximately the same as the base case except that the EFW system did not actuate, and the MFW pump did not trip until auch later.
[ At -J30 s, the MFW pump was tripped because of a high water level in SC B. At l -356 s the HPI was throttlad af ter the 42 K subcooling margin was reached. The l
RCPs failed to restart at this time. At 465 e the HPI was turned back on again because the 42 K subcooling margin was lost. At 600 e the SGs were isolated, and at 4 93 s the HPI was again throttled. The calculation was terminated at 1260 s.
The primary-systes pressure is shown in Fig. 83, and the pressuriser water level is shown in Fig. 84. The minimum pressure calculated was ~5.5 MPs and occurred at ~125 s. After natural-circulstion flows were established (~150 s),
-_ .-. . =. .- .-
wooo -
-30ece aseo. .-
meee so.o_ .
seoso
$ eeoo- -
g .
- o. .
_ ..o.co Neo- -
4 00 o- o
.aeoo , , , . . . ,
--dee o so ano Joe ese too Goo Me o0c ooo M (s)
Fig. 76.
SG A secondary-side water inventory - Case 2.
= . , , , , ,
eg . .'"
yy. .
- g. .
aos
.-mo W R .
Iao. m. ..soo h e- --o
.e .
o so ano ano eso soo soo me eso ese M (s) l Fig. 77.
l SC A secondary-side pressure - Case 2.
l 1
i I
i no . .
- 3eco Soo ,
soon noo.
.sooo ..
Soo-sooo 6 Mo-
.e
. 8 9 3
- r so. ...
me- see
. o
- ooo
=No . . ,
no o no soc ano .co soo soo ooo soo M (s)
Fig. 78.
SG A steam-line flow - Case 2.
>ooo - -
. w -. ,,
soooo.. -.a
'8 8 .
- a. ...oooo
.o o. -
1 1 oo.- .- * ** @
. ,_o m.
.:sooo nooo-
. .o sooo , ,
o me aos no .oo soo ooo me soo ese M (s)
Fig. 79.
SG B secondary-side water inventory - Case 2.
f G
88 , , , , ,
see so- -
m- .
no. .
f
..- ~ ,-no r i s soo So.- -eo
= , . , , . . .
m o so zoo aoo ao soo soo no eso soo M (s)
Fig. 80.
SG B secondary-side pressure - Case 2.
l emt , ,
i 3p3 .
l
- ^*o aa.'~ .
t f
a m ib seo- .
O
^
soo.- . aseo e-- -o
.ao . . , . . . , . ....
e so ano aos ao eco eso no aos ese l
M (s)
Fig. 81.
MFW pump speed - Case 2.
l 1
l l
as . . . . . . . .
(easd)LDOP A1(desf9LDOP A2
"~ h s:(Wtoop at
.=
es-. -
-m g,3
~
3 h u- h s- ..
gg. -
o-- e l
i
"" s s s s 4 s a s 6 m M (s)
Fig. 82.
HPI flows - Case 2.
no . .
.smo
- 8' ~- 2>Je
. W
.mo -
88' .
80-ee-de . , , . . .
'888 i e me me aos ses soon use was l
MW Fig. 83.
Pressurizer pressure - Case 3.
l l
l the primary system began to repressurize (Fig. 83), and the repressurization continued for the remainder of the transient. The slope of the pressure curve changed dramatically at 470 s and again at ~700 m because the HPI was turned on j at 465 e and throttled again at 4 93 s (Table XVIII). The primary system l repressurised more than in Case 2 because the RCPs were not restarted, and i
l natural-circulation flows existed for most of the transient. The pressurizer .
completely emptied by -30 s and began to refill at ~115 s after the primary system began to repressurize. The pressurizer water level decrease at ~350 s and subsequent increase at 4 60 s can be attributed to HPI . throttling (Table XVIII). The change in slope in pressurizer water level at ~700 s can I also be attributed to RPI throttling.
Figure 85 shows the downeeser fluid temperatures for Case 3. The minimum downconer fluid temperature calculated was 450 K and occurred at ~650 s. The downconer temperatures for Case 3 were colder than for Case 2 because of the run-away MFW flow. Asymmetrical temperatures were also calculated similar to the base case. The downconer fluid temperatures increased after ~650 s because the steam generators were isolated at 600 s.
Figure 86 shows the hot-leg mass flows fcr Case 3 and Fig. 87 shows the candy-cane void f ractions. Natural-circulation flows were calculated in both loops af ter the RCPs coasted down at ~150 s. As in Case 2 and the base case, the natural-circulation flow it. loop A van higher than in loop B. Natural-l circulation flows continued for the remainder of the transient because the RCFs never restarted. Figure 8' shows that the candy canes in both loops'pevar voided. The raason the candy canas never velded in Case 3 was becauce the NFW l
provided enough cooling to the loop-B SG to prevent any secondary-to primary heat transfer as occurred in the base case and Case 2.
The cold-leg loop flows for loops A and B are shown in Figs. 88 and 89,'
respectively. These flows are similar to Case 2 and the base case except that the RCPs were not restarted in Case 3. Figure 88 shows that the loop-A cold legs had natural-circulation flows up until the time the SGs were isolated.
These flows then decayed to approximately zero because of significant loss of ,
heat transfer through the steam generators. The temperatures in each cold leg are shown in Figs. 90 and 91. The minimum fluid temperature calculated in loop A was 435 K, and the minimum fluid temperature in loop B was 445 K.
Details of important SG parameters are shown in Figs. 92 through 97. The pressures, inventories, and steam-line flows for SG A were similar to Case 2 and will not be discussed further. However, the SG A secondary-side pressures (Fig. 93) and steam-line flows (Fig. 94) for Case 3 were higher than Case 2 e
u.,_., _
I 7 . .
I
[
s-s . .
[ 4-E
- m s- j 3
-a t.
e I l --e
-i use B00 400 000 We 2 00 o 20
- Fig. 84.
Pressurizer water level - Case 3.
- 0 FWZ .-m 216 ~
- 225 w k
.256 -
246 su g MO -
=256 I. .
26S .-m
.a a
.m aeg. .m 440 , . .
eso use neo o aos eso eso aos TnE W Fig. 85.
, Downconer liquid temperatures at vessel axial level 6 (azimuthal sectors) - Case -
3.-
e
--____m___
l u , . . . .
(p=E$DOP A I -
t-sa- -
gg. .
a4-aa- -
e
-as , , , , , ,
e see see ses ese see see sees TtE (a)
Fig. 86.
Hot-leg mass flows - Case 3.
l l
I a
(pep &DOP A 3eees Gesa9L00P e I
[ ,
sense
- mose i
seee- .
l -aeoo l g_. .
- - -..ee.
l e- A ..e l
i
-nose , , , , . .
--4ees a see 4ee aos ese see neo see Fig. 87.
Candy-cane void fractions - Case 3.
.- .a . _ _ _ _ _ _ ,. _.
esse , , .
sees
. $ sed)LDoP A1 ease.
@menor A2 .
/ .meos
. 00 *'
seos
- m. .see.
h h 888- . .se neo. -. sos e.
~m ..e
...see
.sco ,
see see neo was e see me see Fig. 88.
LooteA cold-Icg maas flows - Cas.2 3.
suco .
~
. (.wyw at .mw ,
4 pescar a2 .
j
.- . ee.
3000-
.ecco @
d l
h, m-. .meo Y -
moo-. . esse e.
k -.
.e
-mee , , , .
see seo noe e see me see see Fig. 89.
Loop-B cold-leg mass flows - Case 3.
r .
l
~
l
u E . E D (s.fu01.DOP A1 .geo eso- W A2 -
(
See- -
E E -
,,..- . me
- g. . ..
p\ .,
m i- f I
g
,1 3,a 33e i 250 450 800 see sees see teos Fig. 90.
Loop-A cold-leg liquid tesperatures -
Case 3.
M i . .
Gor.tvce m soo m- w fmu -
p .m no-
~#
8 g uo-. ..s.
sao-- . --me
=
- m. ,q . .
g I\ (*' l .a 8
~-
\ /
l -
see neo.
. \., p.- .
3
.ee , -
e ase me ese aos see neo i.ee Fig. 91.
Loop-B cold-leg liquid temperatures -
Cass 3.
e w m -- -
mmo .
+
.-3500
- g. l
. . some uneo-
. 2390.
2 2 .
. som
.unoo l
g.
, .-esso aseo.
so ..
. 0
-moo ,
ano ano ano es o ao so no ano Fig. 92.
EG A cecondary side water inver4 tory
' Caae 5.
n.. 3=- ,
m- ,
, ,- eso so.
no So-ooo i m.
I .
. 3eo 3
I ..so a.
-e o.
4 an me
.ao en .
Fig. 93.
SG A secondary-side pressure - Case 3. ,
e D
before 600 s because of the run-away MFW flow. Figure 95 shows the secondary-side water inventory calculated in SG B, and Fig. 96 shows the SG B secondary-side pressure. The ' run-away MFW filled SG B to the 90% operating
- range, and then the MFW pump was tripped (Fig. 97). He level then remained essentially constant for the remainder of the transient. The secondary-side pressure in SG B (Fig. 96) decreased because of condensation from the MFW until the MFW pump was tripped at ~330 s. After SG B was isolated at 600 s, the .
secondary pressure decreased further because of secondary-to primary side heat '
transfer.
- d. Parametric case-Case 4.
In parametric Case 4, the base case was recalculated to 2100 s with boundary condition and modeling changes. After the base case was run, information was provided by Duke Power that resulted in several clarifications in the location and operation of various instruments. This information rasulted in the following changes to the model:
- 1. The MFW pumps automatically trip at 0.5 s instead of 47 s (based on RELAP-5 calculation Ref. 5). This is because of uncertainties in the seasurement of the steam-generator liquid level during transient canditions (LP versus collapsed level).
- 2. T.t a !!PI throttling is based on subcooling at the exit of the core instead of the hot leg when the RCPs are not operating. HPI throttling fs based on subcoolleg in the het leg only when the RCPs are operating.
- 3. Le correct tastrument location for measuring subccoling for ROP restart is locatea 3.0 m belue the top of the candy-cene cecte-lit.e instead of the horizontal part of the hoc leg.
These changes did not dramatically affect the minimum downconer liquid temperature; it reached a minimum of J.20 K as opposed to ~405 K in the base
~
case. -
l The major events of the transient are presented in Table XIX. The transient was initiated by a double-ended guillotine break in the loop-A steamline. Both steam generators blev down momentarily, but the TSV quickly isolated the loop-B generator. At the same time, the reactor tripped, the MFW pumps were tripped, and the feedwater-heater drain tank flow was ramped.to zero.
l Closure of the TSV pressurised the loop-B steam generator and the TBV began cycling open and close to relieve the pressure. This lasted only a40 o because the EW flow into the steam space at the top of the generator lowered the pressure in SG !. . EFW flow began at ~11.5 s because of low liquid level in SG A.
gee
- 4
- t. .. . . .. . . . . . . ... .... . .
woo ,
-Noo
.go.
moo eso.
. gogo g
gee, s
.moe 4 y 8 s no- .@ ~.goe $
ago. k -..
o.
k ..o
. .a.
.ano o
ano ao eso sec zoo are woo Fig. 94.
SG A steam-line flow - Case 3.
i ecooo , ,
I eso0o-
.m Soooo-moooo
^. .
- 8 2 e, o..- .
i
, ,o,e anow-.
m.
onoso tooeo-
. moso moon..
sooo- .,
mooo .
eso isso o ago no goo ese moo Fig. 95.
SG B secondary-side water inventory -
Case 3.
I l
as ,
3 l
=- ..
- g. *-
.ees
}
es-
..e.o l e..
- s. Me eee m-. .een se nos e see me see ses neo see Fig. 96.
SG B secondary-side pressure - Case 3.
me . . .
no. f .
i ..
nee i
see.- .
I f see--
.-seeo i
i .e.
- .. .ee 1 e-. ..o
-ee , , .
e see me e.e see mes see use Fig. 97.
. MFW pump . peed - Case 3.
l .
The steamline break initially caused rapid overcooling and depressuri-I sation of the primary side. At 42.4 s, low primary-system pressure actuated ,
the BPI system and thirty seconds later, the RCPs tripped. The candy-cane in )
loop B voided af ter loss of forced circulation; thus, there was no natural circulation in loop B during the transient. Because the RCPs were tripped, instruments at the core exit were used to determine subcooling for RPI '~
throttling; condit. ions were met at ~302 s. Adequate subcooling for the restart of che RCPs was never met because the loop-B candy-cane was voided. At ~311 s, the liquid level in SG B reached 50% of the operating range and the EFW flow was rerouted to SC A.
At 600 s, all feedvater to SG A was terminated as specified by the ORNL event sequence. At this point, the overcooling transient was essentially over.
The decay heat began to repressurize and heat the primary side. SG A boiled dry TABLE IVIII MSLB (CASE 4) SEQUENCE OF EVENTS Event Time (s)
MSLB-locp A steaulics 0,0 Turbine and reactor 0.5 trip; MFW pump trip; TSV B closure TBV E cycling 5.4-40.0 EFW to both SGs 11.5 HPI initiated - 22.4 RCPs trip 52.4 Loop-B candy-cane 140 voided HPI throttled 302 EFW B terminated 311 SG A isolated 602 SG B restored 900 Calculation terminated 2100 .
e
t .
at about 1875 s. The calculation was terminated at 2100 s because no ,
significant differences from the base case were obtained including the minimum downcomer liquid temperature.
Plots comparing the system pressures and downcomer liquid temperatures for Case 4 and the base case are shown in Figs. 98-99. Differences in the system pressure cannot be seen until -302 s when the HPI was throttled in Case 4; the -
systea was no longer refilling in this case.
The repressurization rate was slower in Case 4 because the RCPs were not operating. In Case 4, flow stagnated in loop B and remained stagnant, and thus little energy was deposited from the secondary-side. In the base case, restart of the RCPh lead to significant secondary-to primary heat transfer and a faster repressurization rate. he downcomer liquid temperatures differ slightly for the two cases; the warming effect of terminating HPI flow ~225 s earlier in Case 4 is compensated by the cooling effect of not restarting the RCP3.
Plots illustrating key events on the secondary side are shown in Fig. 100-107. The break flow and SG-A pressure history (Figs. 100,101) indicate the rapid blowdown of SG A; by 80 s, the generator had almost depressurized to atmospheric pressure. Af ter blowdown, the break flow leveled off at ~250 kg/s until all feedwater flow was terminated at 600 s. Figure 102 gives the pressure for the secondary side of SG B. When the TSV closed, there was an initial pressurizetion, and the TBV cycled to relieve the pressure. EFW began at
~11.5 s and the cold 1.FW caused the pressure to decrease as a result of condensation on the secondary side of SG B.
The EFW flows to both generators are shown in Fig.103. When EFW to SG B was terminated, flow was rerouted to SC A. ORNL specified that EFW be terminated at 600 s; the total feedwater delivered to the generators is shown in Fig. 104 and 105. Even with the MFW pumps tripped at 0.5 s, NFW was still delivered by the condensate-booster and hotwell pumps.
The mass inventory in the tube-bundle region is shown in Figs.106 and 107. After SG A blows down, the feedwater flow equaled the break flow until the EFW to SG B was terminated. At this point, the mass in SG A increased until feedwater termination at 600 s; the generator boiled dry at ~1875 s. In SG B, the EFW and MFW filled the generator until ~400 s.
On the primary side, Figs. 108-113 depict the mass flows and temperatures in loops A and B. The RCPs operated for ~52.4 s and then coasted down. The initial overcooling of loop A increased the mass flow because of enhanced Heat transfer through SG A. Reverse heat transfer in SG B caused the flow directions to reverse for ~30 s before the loop-B flow stagnated. In the loop-A cold legs
. - . _ - _.. - _. -_. = . . -
ue , , . . . . .
. .mae tes - -
.some
$sse Maux so. Gand9 SW CM - ses I co i -
see
- use Ieh- -
use
..q as.
w .
~
see se . . . . . . .
e see ese ne see uno see see asse MW Fig. 90.
l .* System pressure for Case 4 and base case.
t soo , . . . . .
.ses see. .
.see ade- $st$ MRB4 -
g . M anca . . ,
.i .as ese- ~~,,
.e.
.se. .
m.
, -[ ./ .
see
< . .e . ass
(
'., . 2/ se me- .!
't see, see . . . . , . .
e ase see ne see see see see asse MS Fig. 99.
Downeomer liquid temperatures at vessel axial level 6 for Case 4 and base case.
3
-we
3 00 . . . . . . . .
see l
see. , -
ene
.mee see- -
4
.ses
. 1-. .
.mo I
ase. .
. .m0
,.. -N- -
. .e
. 3se
-ace , , . .
0 20 000 20 1000 ESO 900 930 3B00 aste MW Fig. 100.
Break flow for Case 4.
x . . . .
.00e m.. .
g." .
-20
- a. --**
- m. ..a0 I m. .
..soe E
... ..so L_
..e
-m . . , . . . , .
0 30 MG 20 1000 1830 900 930 2000 2300 MW Fig. 101.
- SG-A secondary pressure for Case 4.
l e.,
so . . , , . . . .
- ne's 3
m- -
.e g.
.gge
~
4
...so y lm. '
. -Me
- m. -
. ooo a.
m- . soo m .
1 .
o me eso no ao ano moo son moo me MW Fig. 102.
SG-B secondary pressure for Case 4.
wo , . . . v .
- (esos4 W A -2so no. timav9 m e -
-3eo so. .
me w-
- m. -
-me Q
f.,
So- -
c: : .ao 2 -
" i N m-
.f -.g.
m- ; -.a s
I
- o. : ..e O
woo nas noe suo mee amo i r 1 m i w i !
o me eso no M@
Fig. 103.
EFW flow (SC A and SG B) for Case 4.
\
1 -
l see . . . . . . . .
me
- 3. .
3, . .
~888 3,.. - g me d E .
E c .
.aos c I
mo.
m- -.
--O
--80 0 280 see ne 1000 ste M5 900 2 00 23te M(4 Fig. 104.
SG A total feedwater flow for Case 4.
no , . . . . .
-as as-. -.m
. ass i
. >p .me
- n. -
h -
-os h
g u. -
g -
2 .
-se 5 n- -
0- --O
-n -=4 t , . , .
e me see no neo nas not see anos mano MW Fig. 105.
SC B total feedwater flow for Case 4.
9 e
i mooo - . , , . . . .
reso oooo. .
esso ass-. .soon I
} *~ .
l t .noso 2 oooo- f l
--oooo g m.L j ti
- aooo- -.,,oo h
.o
~"* see sooo ano a ao eco no eso ew ok M (s)
Fig. 106.
SG-A tube-bundle-region mass inventory for Case 4.
anooo , , .
Nooo racco-
~
.ooooo m .soooo 2 ,oooo.
2
-aoooo l, m.
-soooo y E
~
noco- .
. ,some oooo-[ --esso l
. e o
moo ano eso soo anco m i
o asa soo - . an M@
i Fig. 107.
i SG-B tube-bundle-region mass inventory for Case 4.
l t
l l
e s
9 9
-100-a y a u a i a i acaso I (se64DOP A -asses sooo. : Gha h)LDOP a . , , ,
seos sooo- -
g I
g~. -
es.o g
- l moo-.! --4000 o.. r h a- ^
..o f
-moo , , .
--mos e me soo no woo nas moo seo sono a so MW Fig. 108.
Hot-leg A and B mass flows for Case 4.
e 4 i B I I
..so 588- -
-sro (peacCLDOP A
..i (p n)LDor a .us 2 i E wo-
.} -.o.
I 1
...so
'~
l* deo-r- ---y '_
l
-ase
- ' ~
.aeo den , , , , , , . ,
e ano eso me woo ano soo see anno sano MW Fig. 109.
i Hot-leg A and B liquid temperatures for Case 4.
I
.~ ,
_ . . . . . - - - ~ . . . .: - - - - - - - ~ - ~ ~ . - - - - -- - - - - ~ ~ ~ - -- - --
-101-d is. . . . . .
. -s 4
(pd$DOP Al mn
~.. .
- e
~
^ a o- -.
en- .
-am
. .m.
- I ' 7 sa m e n n m. rs. m.
i M@
Fig. 110.
Loop A cold-leg mass flows for Case 4.
m . , . . . .
sn - .
(polidADOP Al -
sas- pg -4 E E
- w ,,,.
2 e l es- -
3,
~
- s -s a.-
l ) '
) - -
\ -3
\ ll ~
.m
= s -.
M@
Fig. 111.
j Loop A cold-leg liquid temperatures for Case 4.
s .
w
- "1% , k G
5 mny,-- a e n - ,+
'- 102-sooo . , , . . . . . ;
.meos m e. .
(iwagoope1 . esso (Id owyLDoP M 3eo . . ..
5 --.
sn
..o 5 g
h 3 moo-. -. sono o-. -Q.n. ;w .:- - = - = - -.o
~~
.toco .
, o ao ano no soo nas noe see soon sono MW Fig. 112.
Loop B cold-leg mass flows for Case 4.
I l
ooo , . . . . . . .
.ooo IPs- -
. son
( agoopsi E
Wm ~
ano sas-.
E I.. es- -
f" ,,,
- -doo h' .J
- m. ,, _, / Ni
.sas
- m. .
.aeo doo , ; , , .
o ao soo no sooo ano moe soo sono aaso MW Fig. 113.
Loop B cold-leg liquid temperatures for Case 4.
4 ne
_ _ _ . . . . ._____..-.._____.i..
1
-103- :
. . j high natural-circulation flows kept the fluid mixed and the temperatures uniform. In the loop-B cold legs there was a small circular flow pattern.
Figure 114 indicates the total HPI flow for the transie'nt ; the flow was l decreased to zero when adequate subcooling was reached at the core exit.
Figures 115 - 116 show the total vent-valve flow and the average upper-plenum liquid temperature. The vent valves make a significant contribution to .
warming the downcome- lic.uid. The total vent-valve flow for Case 4 is similar to that calculated for the base case (Fig. 36) except that the vent valves were always operating in Case 4 (no RCP restart).
In conclusion, Case 4 gives a minimum downconer liquid temperature of approximately 420 K and indicates the system will repressurize similar to the base case. The chaeges made to the base-case calculation for Case 4 gave a more accurate prediction of the postulated accident at Oconee-1 and still indicate that this transient could pose a threat of PTS to the reactor vessel. However, these changes made no significant impact with regard to the overall conclusions stated for the base case, and the minimum downconer fluid temperature remained approximately the same (-405 K - base case, 420 K - Case 4).
- 4. Conclusions. The overcooling of the primary side of the Oconet-1 plant caused by a full double-ended steam-line break in one of the steam liaes was simulated with TRAC-PFl. The main forcing function for the overcooling was a delay by the operator in isolating the affected steam generator coupled with a delay in throttling the HPI flow. The base case analyzed had all plant
, protection and control systems operate as designed. The minimum downconer fluid i
temperature calculated for the base case was ~ 405 K. Repressurization of the primary system to the PORV setpoint was calculated for the base case following
< an initial depressurization to ~3.5 MPs. .
Three parametric cases were analyzed in addition to the base case. In the first two parametric cases (Cases 2 and 3), input and modeling errors prevented the EFW system from operating as designed. Case 3 had an additional input error i that prevented the RCPs from restartirg once adequate subcooling was achieved.
l For these parametric cases, the downconer fluid temperatures were considerably higher than the base case (Case 2 -475 K; case 3 ~450 K); thus greater l margins against PTS were calculated.
In the last parametric case (Case 4) other changes were made to the model to reflect information provided by Duke Power af ter the base case was run. In this case, the subcooling monitor was corrected and the MFW pump was tripped at 0.5 s. This case was run to 2100 s, and the minimum downconer fluid temperature
, . . - -.. - ~ ,- . - - - - . -.,,--re.,-g
-104-00 . . . . . . . . -es .
m- '
.0
.0 so- "
30 g .
1 g* .
~
.. 1 2 -
se z
,, .-as O
O B0 000 20 1000 She 900 SN0 2000 2330 M@
Fig. 114.
Tots 1 HPI flow for Case 4.
40 . . . . . . .
- m. .
me. .
- m. .
2se. -
i
- m. ,.
j m. -
~
..N .
0 s s m a s 6 m s m
%)
Fig. 115.
Total asas flow through vent valves for Case 4.
9
- - _ '_ _ _ _ _ _ --- - -- _ _ - _ _ _ - _ a_m___ _____ _
I
_ .. . . . . . . , . . . . . , _ . - ~ - - - ~ ~ ---.- ---.----- .-
l l
-105-ses , . . . . . .
. see ses- -
- No see ._
ee lase- . ,,
sue <
des es
, as me. -
. ase 88' ase e sie sie vie use noe see sie sees aus Thf M Fig. 116.
Upper plenum liquid temperature for Case 4.
was 420 K. None of the changes incorporated into Case 4 resulted in significant differences from the base case.
- 1. Introduction and Summary. This section of the report presents the Oconee-1 plant response to a small primary-system break (the failure of the PORV in the full-open position). The PORY was assumed to open at transient initiation and remain stuck-open for the remainder of the accident sequence.'
This event was followed by the reactor and turbine trips from full power. In addition to the PORV failure, the ICS failed to run back the main feedwater. As a result, the steam generators continued to fill until the MFW pumps were tripped c. a high SG 1evel signal. The PORV failure and the ICS failure were the only assumed system-related failures. The RCPs tripped 30 s after HPI
- initiation, and this was the only specified operator action.
TRAC calculated a minimum vessel downconer liquid temperature' of ~528 4 between 600 s and 700 e into the transient. The primary system was calculated to repressurize to ~11.5 MPa after 800 s.
l 2. Model Description and. Assumptions. A complet s description of the primary-side, secondary-side, and ICS modeling can be found in Section II. The j steady-state operating conditions are also presented in that section.
l
\
-106-TABLE XIX MSLB8 (CASE 3)
SEQUENCE OF EVENTS Event Time (s) ,,.
1-14. Approximately same as base case 0-53.9 except no MFW pump trip at 47.8 s
- 15. SG B level at 50% 197.0
- 18. HPI on (42 K subcooling margin lost) 465.0
- 19. SG A, B isolated 600.0 l
l 20. HPI off 693.0
- 21. Calculation terminated 1260.0 aEFW pump never started; RCPs nevar restarted b
Because of signal-variable errors in the ICS modeling the MFW pump did not trip until 331.0 s. This pump should have tripped on low-suction pressure sim,ilar to the base case.
The PORV LOCA specification containing the initial conditions, event sequence, and assumed failures is presented in Ref. 3. The TRAC transient event sequence for the PORV LOCA is prescnted in Table IX. To ensure a MFW pump trip on a high SG 1evel signal, the low-suction and high-discharge pressure trips that could prematurely trip the MFW pump were overridden. Also -the MFW pump speed was incressed to its rated maximum speed (595.8 rad /s) whereas the loop-A and -B MFCV flow areas were maintained at their steady-state operating values until the MFCV overriding trip at ~100 s. At this point, the SUFCVs were opened by the ICS to continue filling the SGs. The MFW pump maintained the maximum speed setting until the trip on high SG 1s vel.
S 4
-107-
- 3. Transient Calculation. Figures 117 and 118 present .*ie secondary-side ,
pressures for loops A and B, respectively. Immediately following the reactor and turbine trips, the TSV closures produced an increase in secondary-side pressure. The TBVs for both loops were repeatedly activated between ~4 s and
~100 s to relieve increases in secondary pressure. The differences in the pressure distributions between ~150 and ~250 e can be attributed to differences
. in NFW mass flows to each SG. Af ter the RCPs were tripped, the MFCV override trip closed the MFCVs, and the MFW was realigned to the upper header of the SGs.
The SUFCVs were opened fully by the ICS to continue the feed to the SG upper headers. The MFW aass flow and liquid temperatures for loops A and B are shown in Figs. 119 and 120, respectively. Figures 121 and 122 show the NFW (realignment) mass flows supplied to the SG upper header of both loops. The additional mass flow to the loop-A SG produced the lower loop-A secondary-side pressure by condensing a portion of the steam located in the upper levels of the SG, and caused the larger loop-A secondary-side inventory. The SG inventories for loops A and B are presented in Figs.123 and 124. The ETW pumps were not activated because the MFW pump was able to attain the high SG 1evel.
TABLE XX
_ PORV LOCA EVENT SEQUENCE Event Time (s)
- 1. PORV cpens 0.0
- 2. Turbine and Reactor trip 0.5 ,
- 3. Turbine-stop valves close 0.5
- 4. Secondary-side heater and heater-drain trip 1.1
- 5. Condenser feed from turbine trip 1.6
- 6. TBV-loop A opens for first time 4.4
- 7. TBV-loop B opens for first time 4.7
- 8. Condensate-booster pump trip on low-suction pressure 11.0
- 9. TBV-loops A and B open/close 16.2 '
- 10. HPI actuation on low primary-system pressure 70.3
- 11. TBV-loops A and B open/close 71.1
- 14. MFCV override trip 100.3
- 16. TBV-loop B opens for last time 330.0 l
- 17. Minimum primary pressure (~7.2 MPa) attained 550.0 !
- 18. Pressurizer water-solid 600.0 -
]
- 19. Maximum primary repressurization (~11.5 MPa) 850.0 4
- 20. End of calculation 1000.0 l i
=, . , . - - . . , - , - ~ . . - . . - . . . - , . a.-,-----_n-- w. ,.-n -, n ... , ,-- . . . , , . ,
-108-s s i s a a i a a
-M
- g. - - m75 I
1
-soso n- -
l -iozs 7 m- -
-sooo 1 u--
- s?$
un
- u. -
-sso 64-. --ers s2 . . . . . , . soo no o zoo soo 400 soo soo 7ao soo soo 1000 Time (s)
Fig. 117.
SG A secondary pressure.
so , , ,-
ys- _
-'"O 3_ -
-illo t
y_ N _
-meo 8 T 72- ') --8050 $
70-
/) --mao so- .-sso u.- . - eso <
I e4- . -sso M . , , , , , , M o
too zoo soo 40o soo soo 700 ano soo moo Time (s)
~
Fig. 118. ,
SG B secondary pressure.
s.
\
.s . -
-109-i soo . . . , , . . . -oso '
l p (soratoop A -
' 708- - noa j (dow9toor a .
i
)"
Goo-
-uso y m- -
-mo g
- m. .
l
-7so m soo-m 2
aco-
_ '.i
-soo C
3 me.- lf' .-250 o- - '- c- o
-mo . . , , , . , . .
moo -
o no zoo soo 4o0 soo soo roa soo soo Time (s)
Fig. 119.
Main-feedvater flow - loops A and B.
S'S -
- s2a
. (solid) Loof A po-(do m -
-se 333 g - -soo p v
- - 4to d
~
sas- ,,.*~~~
'..~ -4so W szo-
..- .m sis-i so-
.....~ -.
~8 so5 , . , , . .
o No 2co soo doo soo Goo 7oo Goo too looo Time (s)
Fig. 120. .
Main-feedvater liquid temperature - loops A and B.
-<s,- - '-w
-110-soo . . -m' i Celi m- 2
-.co3 Goo- -
-1B; Soo- . -
a h
-tocD doo- -
g -7M 300-y g
$ '"~ -Sod '
1 (
~
1 go. . -2M o- .
. o
-loo -. -
. -2:o STGDI TEC 2
-2co . . .
o roo aos soo soo moo Time (s)
Fig. 121.
Realignment mass flow - loop A (negative flow is into SG).
800 '
- 2iao CELL 2
soo- -
-tm
-won ga m- -
6
-ios3 -
E l 4o0-1 I
i y ~
.su o-- ^0
' siccw -
--so , 72*
o zoo 4o0 soo soo moo Time (s)
Fig. 122.
Realignment mass flow - loop B.(negative flow is into SG).
4 w - - - . - , - -, __ h
-111-scooo , . . . . . . , .
. .escoo m.
-.soooo
- m. .
g 3ecoo- ,
_3,,o g w
3 coco-g . ..cooo g N asooo.
.-aeooo oooo..
gooo. -
- -3Cooo icooo o tbo 2co 3bo 4do sbo abo 700 abo 950 900 Time (s)
Fig. 123.
Steam generator secondary inventory - loop A.
Snoo . . .
- -soooo 3sooo-
.__/
3:3oo.
. 7eono 3acoo.
^
o 6 z7soo.. -
.soooo $ -
$ Q
$ 25000-3
--50000 22soo--
so m .
4oooo esco-noon- .
l . 3oooo esco . . . . . . . . .
o too 2ao 300 doc soo eco 7ao too too looo Time (s)
Fig. 124.
Stean generator secondary inventory - loop B.
4
-112-Pressurizer pressure and water level are shown in Figs. 125 and 126, respectively. The primary-system pressure fell sharply until the HPI was activated on low primary-system pressure at ~70 s. The primary system continued to depressurize and reached a minimum pressure of ~7.2 MPa at ~550 s. At this time, the primary system began to repressurize, and by ~850 s reached ~11.5 MPa.
As soon as the HPI was initiated, the pressurizer began to refill because the -
HPI mass flow was sufficiently larger than the break mass flow. The break (PORV) mass flow and vapor fraction are shown in Figs. 127 and 128. The pressurizer remained voided until it refilled at ~500 s. Following this time, the pressurizer vapor fraction decreased rapidly and correspondingly the PORV mass flow increased. By ~1000 s the primary-system pressure was in equilibriuin.
Mass flow rates and liquid temperatures for the loop-A and -B hot legs are shown in Figs.129 and 130, respectively. The primary-system flows decreased following the RCP trip at ~100 s, and natural-circulation flows were soon established. TRAC calculated a minimum hot-leg temperature of ~552 K for both loops (Fig. 130). Loop-A and -B cold-leg mass flows are presented in Figs.131 and 132. The cold legs exhibited trends similar to the hot legs. The
! corresponding cold-leg liquid temperatures are presented in Figs.133 and 134.
Minimum cold-leg temperatures were calculated to be ~518 K at ~550 s for loops Al and A2, and ~525 K at 425 e for loops B1 and B2.
The void fractions for the loop-A and -B candy-canes and the upper plenum
_(level 8) of the vessel are shown in Figs. 135 and 136, respectively. No voiding occurred in the candy canes; however, the upper plenum voided slightly between 400 s and 450 s.
Figure 137 shows the downcomer liquid temperatures at the top axial level just below the cold-leg inlet nozzles. TRAC calculated a minimum downcomer liquid temperature of ~528 K. The system pressure at this minimum downconer liquid temperature was calculated to be ~7.2 MPs.
- 4. Summary. The Oconee-1 plant response to a small primary-system break (failure of the PORV in the full-open position) was calculated with TRAC-PFl.
In addition to the failure of the PORV, the ICS failed to run-back the MFW, which resulted in a MFW pump trip on a high SC level. The only specified operator actions included a RCP trip 30 s af ter HPI actuation. TRAbcalculated a minimum downcomer liquid temperature of ~528 K at 400 s into the transient.
The primary system was calculated to repressurize to ~l1.5 MPa af ter ~800 s.
- me
,.,m,_ , - - - - - _ - - - - - - - . _ - - - - - - , . -
-- ..... .... .-. . ~ ~ .~ , . . . . . - . . . . - . --
-113- .
no . . . . . . . . . .
-2200
- 8- --200o I
no-
-moo m.
. N
. --w0o m,,o. f I e.
-20o
.. s0-
-noo so-70- _ -30ao so , . , , . . . . .
o no too soo 4co soo soo 7oo soo soo moo
. Time (s)
Fig. 125.
Pressurizer pressure.
g 3 p . .
. . _-4o m.
-30 s-
/
oc a:
--20 le s.
4
-10
~
2-0 o . . . . . . . -
m moo soo o 10 0 200 300 400 500 Goo 700 Time (s)
Fig. 126. -
Pressuriser water level.
t I.
-114-so . . . . . . . -
-no So- .
.mo
.o. .
g -so .-
& t
- u. .
-so
" so- - *
-4o m* *.,4 o-- --o
-to ,
o no zoo soo 4o0 soo soo 700 aoo soo woo Time (s)
Fig. 127.
PORV mass flow.
t2 . .
1 -
o.s - -
5 g o.s - -
E os- -
l o.2 - -
o-
-o2 . . . , -
o no zoo soo 40o soo eco 7oo soo soo . mac l Time (s) l Fig. 128.
j PORV vapor fraction.
9
. ~ - .
~'
-115-300o . . - ' ' '
-ioeco
"' (schd)LDOP A1 W (doorQLDOP A2 sooo-
- m. ,
-Moo Ra moo-
_sooo f 4 -
moo.
.- noo ,
g 2000-~
h moo. ,
~
-3000 3o00-
-Soo soo-
--o
,o - -
-sco , g, g, . 3o ooo yoo soo no m Time (s)
Fig. 129.
Hot-leg mass flows - loops A and B.
Sooo ' '
- 2500
" (solid) LOOP B1 -
~C (doWQLDOP B2 me ..
3Soo- , -7500 2soo-
-sooo .
.- moo g 2 co-- "
~
2 - 500- , -3000 moo-
- - 15Co soo-
--o o.--
~5" o go soo soo do soo eco a em no em l
Time (s)
Fig. 130.
Bot-leg liquid temperatures - loops A and B.
-.y - ._ .-
-116-soooo , , , . . . . . . .
- 2ioco tooo- (solid) LOOP A .
F (dosh) LOOP B em. ,
-~ *
- moo- .
.nooo g , coo. .
3 f
y sooo-
-nooo g g
b coa.- --8000 d M
5 sooo- -
.m acoo- -
-sooo goo. . -
--o
-moo . . . , , . , , , ,
o too roo soo 4oo wo soo 7oo soo soo moo Time (s)
Fig. 131.
Cold-leg mass flows - loops Al and A2.
ses . . . . .
. .io (solid) LOOP A '
(dosh) LOOP B -soo ses- -
-seo g m- -
p
,, .seo -
sn- -
cm
.sm sm- -
-Soo r ses- -
sea.- g .-sao a 566- --soo sno.. .
.sso l s4s , , , , , , . . .
o too zoo ano 4ao soo soo too ar,o soo icoo Time (s)
Fig. 132.
Cold-leg mass flows - loops B1 and 52.
l l
I
-117-sao . . . . . . . . . -
- m. -
I
-ses
~
8 _ .uo 9
Sao-
. m Sao- -
3g i
33o.- --es i
-430 S2o- -
. .. -435 S2 . , , . . , , .
o too 2co 3co doo Soo 600 7o0 aco too tooo Time (s)
Fig. 133.
Cold-leg liquid temperatures - loops Al and A2.
sao , . . .
(sasid) LOOP 81 (dosh) LOOP B2 - 57o syn. 6 -
-566 b g. - b y
k ac Y
g Seo- -
- 525 Ho - -
-se s,, - ..
_ 4so sac . . . . . . . . .
o eo 2co ano 4co soo soo roo soo soo moo Time (s)
Fig. 134.
Cold-leg liquid temperatures - loops El and 32.
-118-1 . . . .
R TH Z 118 '
a.s - o128 -
138
+148 h o.. .
= 15 8 . , , ,
W 168 E
CA- -
0.2 -
o o
= = = . = = -
2co A
4oo soo soo moc TIME (s)
Fig. 135.
Candy-cane void fractions - loops A and B.
t2 (sor.o) LOOP A
,. (dosh)LDOP B ,
as- -
o.s - -
E os- -
. o.2 - -
0
-o.2 . . . . . . . . .
o no ano soo 4o0 soo soo no soo soo 1000 Time (s)
Fig. 136.
Vessel upper plenum void fractions - all azimuthal cells.
O e
.9,
_ _ _ _ _ _ _ _ - - _ _ _ _ _ _ _ - - _ .m__- --.-- __----2 _ _ _ _ _ _ _ _ _ _ _ _ _
. - ~ ~ ~ - -- - ---
-119-m , , ,
- m. R TH Z .
216 .,,, ,
- - o226 -
236 -s30 E 246
~ ~
.- 256 -s.o )
l*
- 266 l se0- - -
-s3s w-- \
-saa
\ -
-se 330
\l 4:,0. .:
-490 sas , , , ,
0 200 400 600 s00 1000 TWE (s)
Fig. 137.
Downcocer liquid temperatures at vessel axial level 6 (all azimuthal sectors).
D. Turbine Bypass-Valve Failures
- 1. One Bank of Two TBVs
- a. Introduction and Summary. For this study, the performance of the Oconee-1 plant following a secondary-side depressurization was predicted. The base case analyzed was the failure of one bank of TBVs to resent af ter initially opening following reactor and turbine trips from full power. Additional, failures assumed for the base case were failure of the level control in the affected steam generator, failure of the operator to restart the RCPs, and failure of the operator to throttle the RPI system. The lowest downconer liquid
, temperature ( 458 K) and hence the smallest margin against the NDT limit was
(
calculated for the base case. Repressurization of the primary system to the PORV setpoint was also predicted for the base case. In additional parametric cases that examined a reduced number. of failures, a greater margin against the NDT limit was calculated.
- b. Model Description. The primary-system model used for the TBV transients is shown in Fig. 1. On the secondary side, the main-steam lines from each steam generator to the TSVs are modeled. The turbine bypass lines lead to the condenser, which is input as a pipe with a constant-temperature heat sink.
The condensate collects in the hotwell. The hotwell and condensate-booster
, - - n + w
-120-pumps deliver the condensate to the feedwater heaters. The main-feed pumps then pump the feedwater to the steam generators.
Ihe significant features of the TBV failure transient are (one bank of two TBVs):
- 1. Reactor and turbine trips cause the TBVs to open. -
- 2. Failure of one bank of TBVs to close causes a secondary-side depressurization through the affected loop.
- 4. The operator does not restart the RCPs.
- 5. The operator does not throttle the RPI.
Two parametric cases were also calculated. The steam generator liquid-level coatrol in the affected loop operates correctly in Case 1. The steam generator liquid-level control also operates correctly in Case 2. In addition, operator actions to restart the RCPs and throttle the HPI are permitted if the primary system subcooling-sonitor trip points are exceeded.
- c. Results.
- i. Base Case. Table XXI presents the calculated event ti:nes for the base case. Following the reactor and turbine trips, the TSVs closed (0.5 s),
secondary pressures rose, and the TBVs opened for the first time at ~4 s. The secondary pressure peaked and then decreased permitting the loop-B TBV to reseat at ~45 s. However, the loop-A T3V failed to resent causing the secondary to depressurize at a f aster rate than loop B.
The pressurizer pressure is presented in Fig.138. The PORV opened at
- 1
~1037 s when its pressure setpoint of 16.99 MPa was exceeded. The PORV then i cycled for the remainder of the calculated transient to maintain the primary system pressure at or below the PORV setpoint.
The loop-A (affected loop) and loop-B secondary pressures are shown in Figs. 139 and 140, respectively. The stuck-open TBV in loop A caused that loop to depressurize both more rapidly than loop A and to a lower level. MFW flows for both loops A and B are shown in Figs.141 and 142, respectively. Per the problem specifications 3 the ICS failed to run back MFW to the loop-A SG and thus the flow did not decrease until the MFW trip on high liquid level in the loop-A SG occurred at ~61 s. The ICS ran back MFW flow to loop B by shutting the MCFV and allowing MFW flow only through the SUFCV. The small oscillatory flow in loop A was related to variations in the loop-A SG secondary-side liquid l
?
l
.. . ~. . . . . . . . . . . . . . . . . - - - , . . - . - . ~ . - - --..... .. .. ... . .-
-121-30 ,
m -
- m. / .
- -2350 f
.e sAst
/ PARAWETRIC 1 .
we- / 9 Pu* cmc 2 -
3
,s
.,/ ..eso f
./s**.
l me- ,
--900 e0-
\,j \ --
.e
. 00
~
.0 ,
-no
- 40 . ,
O N0 400 e00 e00 1000 000 #00 1500 M (s$ prusure (bors)
Fig. 138.
Pressurizer pressure.
e0 .
80- -
M' ggggg g *-900 PAmmCDec 2
,o. .
- e00 .
90- -
--tes l40--
so. .
'%.s. ' -- . -----
ees 30- ,.,%, . . . _
- n. - -
0 . , , , , . . e e see ese see see noe see mee use MW Fig. 139.
Steaa generator-secondary pressure - loop A. .
_ _ _ _ . ___m___.
-122-g 5 3 5 3 &
.no.
~~
saw Penaapec a .meo M- PWlaaCr,2 -
~#
e.- . .
%s g
so-. i
N '- -.m R
. en w-. \'s N g -.
So . , ,
"M e soo 4eo soo soo woo soo woo woo M (s)
Fig. 140.
Steam generator-secondary pressure - loop B.
soo _,
)
l N'. -soo
. aAu .
do PanautTmC1 -ene PARA 600C 2
~
sea. .
_ , h 3
2 .
..o. 3 m- .
.o.- . -me o..:-- A..A
-me . . , , , ,
so e eso ese see see eso moo see MW Fig. 141. .
Main-feedwater flow - loop A.
l
-m.
-123-TABLE XII l
- TBV EVENT SEQUENCE i
i Base Case Event Time-(s)
- 1. Turbine and reactor trip 0.5
- 2. Turbine-stop valves close 0.5
- 3. TBV loop A opens (fails to resent thereafter) 4.1
- 4. TBV loop B opens 4.3
, 6. HPI started following trip on low pressure 153.1
- 8. Feedwater realignment trip 183.0
- 9. Main-flow control valves overriding trips 183.0
- 10. EFW pump on 209.1
- 11. Loop-B E W valve shut on high SG liquid level 460.8
- 12. FORV opens 1036.7
- level. Loop-A and -B MFW liquid temperatures are presented in Fig. 143 and 144, respectively. An increase in the loop-B temperature followed the feedvater realignment trip at ~183 s. The temperature stabilized at the saturation temperature af ter the SG secondary was isolated, at 460 s. ,
EFW flows through loops A and B are shown in Fiss. 145 and 146, respectively. The loop-B flow decreased sharply at ~460 s when the loop-B EFW valve shut as the SG-B liquid level exceeded 6.2 m (240 in). A residual flow continued through the loop-B SUFCV until it shut under ICS action at ~600 s. A higher flow through loop A was predicted because TBV A was open producing a larger pressure drop potential for flow through loop A. Loop-A and -B EFW liquid temperatures are presented in Figs.147 and 148, respectively. A rapid rise in the loop-B fluid temperature followed closure of the EFW-B valve. A small flow of hotter fluid through the SUFCV produced the temperature rise.
The water inventories in the loop-A and -B SG secondaries are shown in Figs.149 and 150', respectively. The inventory in the loop-A SG rose before 40 s because the ICS failed to run back the NFW. The MFW trip occurred at
- - . . , , - . - - -- - ,- e, -, ,- n., ,..-,----.e , m + - --v-
-124-5 4 5 3 5 5 3
"'. ~. ,se eso. W BASE .
e=4 rasawnsc s . se
, PMAdTieC 2 ese- .
-See .
h m soo- .
.m l
,se. .
so. .-see
= -
e.
~ ; .
.e
-me -
e see me ese eso see see mee see MW Fig. 142.
Main-feedwater flow - loop B.
I I g sAsc PanaMEPuc 1 me em. PMAME11tc 2 .
s:
g , m
"~ N 8 . 'g
.m g .
- ' .'sg N '- -- s .-me I me.- me.
N.
.3,e 30 .
me me. .
.ses 6 . , , . . .
e zoo me ese ese see see use see NW Fig. 143.
Main-feedvater liquid temperatures - loop A.
1
-125-sm , , , ,
Sao- - -
ses- \, - -
= .
.. e g .. x. .
so-
\, -
o,o.- - .
... ...o soo-. \ -
mo o ano go a ooo woo ano woo soo w (s)
Fig. 144.
Main-feedwater :3.iquid temperatures - loop B.
soo ,
. coo 250- --
.soo 2ao- BASC -
. MRAMETRtc 1 .gg j mancmc2 g .
- a. , . . .
-3oo
- ~ ~ - -
__- ,r,~
, y ~~~~~ -
l so-. -.
J
.so .
ao
. o zoo soo ooo soo ano woo woo w (s)
Fig. 145. ,
Flow through emergency-feedwater header -
loop A.
l l
l w -
,, ,-.n . - - - -
i
-126-to , , . . .
.se to- ,
oo.. . sast -.es 8
PARAaC1RC 1
{ /
PARAAC18C 2 g
no.
a s
1 2
e V A -m h
.. 'N ./ V .
g h .
'..s .m 3 m.
Ni -
s g -m
, l i '\ _
._m
-m , . .
o zoo eso ooo soo woo nao :soo soo M@
t Fig. 146.
Flow through emergency-feedwater header -
loop B.
I W >
- - ' ' N ',7*-= ~
sos
~
~..
no.~ l
/
,r y
east 8 as.
/ / PARAnCIMC1 .-3. E
[ / PnRAtCRIC 2 -
l Iaoo._ . /
,.s ago.~ .
-so as-- -
Soo- -
as
~
e ano aos eso eso soo eso tion soo MW
! Fig. 147. .
l Liquid temperatures in the emergency-feedwater header - loop A.
S a m w- a - - -- - _ _ _ _ _ _ _ - _ - _ _ _ - . _ _ _ _ _ _ _ _ _ - . . _ - . _ _ _ _
l l . .. ... ;
I
-127-i s 4 as.- * -
- m
- ~ #.
8 , / fARMCmC 1
~'
S V) / FARMCNC 2 2 m-- l soo m- ,i -.m m- ik -
,e .m
/
soo- ; -
. :, -oo m- ,!{A -
J \ -wo sea. .
So m
o m 4co ooo ooo woo noo woo moo M (s)
Fig. 148.
~
Liquid temperatures in the emergency-feedwater header'- loop B.
ooooo
..oooo m aa. .
.. ,eA,- -moooo
^ ' -
sAst 6 -
- 88000-g /
f PARAutTmC 1 PARAaC18C 2
-.aom g$
5 ,/ j/ ***** I haccoo- / < .
/' / -oo m
, coco. l' .
-ooooo r
aoooo- -
. moo mooo , . .
o m 4no ooo see eco noo woo soo w(e)
Fig. 149.
Steaa generator-secondary inventory - loop A.
4
-128-
- 40 s, and the liquid inventory boiled off until the EW pump actuated at ,
4 10 s. With no liquid-level control, the steam generator secondary continued to fill and oscillations developed as the liquid level reached -12.4 m. The transient history of the loop-B SG was. very different. There was no initial increase in SC inventory because the ICS reduced flow. Closure of the loop-B SUFCV by ICS action at -400 a was evident.
~~
Mass flows through the primary-loop hot legs are shown in Figs.151 and 152, respectively. Following the RCP trips at ~183 s , the flows coasted down and natural circulation was established. A higher flow rate was established in loop A because of the higher SG temperature difference resulting from the failure of TBV A to reseat. Similar phenomena were observed in the cold legs as shown in Figs. 153 through 156, respectively. The corresponding cold-leg temperatures are presented in Figs. 157 through 160, respectively. Vapor fractions in the loop-A and -B candy-cane sections are shown in Figs.161 and 162, and it can be seen that no voiding occurred during the transient. The pressurizer water level is presented in Fig. 163.
Downconer liquid temperatures for the base case are presented in Fig.164 at the top axial downconer level (just below the cold-leg nozzles). At the end of the calculated transient (1500 s), the minimum temperature was -458 K.
ii. Parametric Case 1. A single specification was changed for this parametric study. The loop-A SG 1evel control following EFW activation was assumed to operate to maintain the secondary-liquid level at or below 6.2 m. It was assumed that the operator does not restart the RCPs and does not throttle the HPI. The event cequence for this case is presented in Table XXII. The event sequence was identical to the base case through event 10. At ~290 s the loop-A EFW valve shut on high SG-A secondary-liquid level. The loop-B EFW valve' shut on high SG-B secondary-liquid level as in the base case. However, the PORV opened 40 s early in Case 1 because of reduced heat transfer to SG A.
Results for Case 1 are presented in Figs.138 through 163 and may be coapared directly to the base case. The general trends of Case 1 were similar to the base case. The major differences appeared in the secondary side of loop A, and were caused by shutting the loop-A EFW valve on high SG liquid level at 490 s. The reduced EFW flow affected the primary side also. ~ Compared to the base case, the pressurizer pressure increased more rapidly to the PORY setpoint as shown in Fig.138. This was caused by reduced primary-to-secondary heat transfer associated with the reduced loop A steam generator-secondary water inventory (Fig. 149). The steaa generator-secondary pressure histories for loops A and B (Fige. 139 and 140) were similar because the level control l
l 1
l
-129- i J
soooo
-waeo .
.coe. -
- '. ^.sosso f .
2* <
/ M Earmc1 ~.nooo I 9-a g g
3 .moso. -
E j
ooooo l
o soooo. --ecco l/
~-
\.
.,oooo
.ooo .
o X2 doo ooo SM icoe dos teso 4 MM Fig. 150.
Steaa generator-secondary inventory - loop B.
mo , , .
sooo-m ,
-noos g Sooo sooo. PnAMdCTm: t (CM M*E 2 - ecco l
m- -
. oooo .
.h f3cos. __ - .
9- a
- 1 . sco-. .
e
-.. a Woo-. -m
- o. --o 4- i ..
e soo 4eo eso aos a nao isso aos 1 MM 1
.; Fig. 151.] l Eot-leg flow - loop A.. l 1
i j
)
..'t.,
I -
- i . <
l
-130-wee . . . . . . .
r .
u oo eseo.
~ ~8"'
i sau 9000- I P4RMCNC 1 -
I PnAMCIWC 2 aneo.
C 30.o f
mee. ,
.so.o 5 1ees- -- 30.
- o. ..e
-moo . , , . .
--soon )
r .
e aos .co soo soo moo soo woo sco e M (s) i Fig. 152.
Hot-leg flow - loop 3.
woo
- te
.ooo. , , _ . .
sau -"
3000- P4RMCloC 1 -
- PanMCac 2 . ,,,,
.p
.ooo. .
-M .
f m-.
.a l h I h
^ ' * * '. -
.m R iogo-. ' -
200.
(
... ..e
.io o .
--ases e aos em ese see moo soo woo neo MW i
, i I Fig. 153. *
! Cold-leg flow - loop A1.
f
-_ g
-131-I sooo g
.coe. .
. .oooo east ' ' '
30eo. .
maauttmC 1 -
M2 oooo
'***~ ~.me f
g .oo ..mo 5 m
h o.- -
e
. co . --me
. . . . . ._ . so i ..coes 300o .
o zoo eco ico ooo eco soo woo soo MM Fig. 154.
Cold-leg flow - loop A2.
. me teoso sooo- , ._.
noso usc sooo. masurtmei .
. PAucac2 . ,,,,
.oco-
-coco
(
f,a m e. .
. coco El
"$ sm-. '
-.. 1 i
i Woo- 8
--2eco o.. --e
. .aeo . . .
--me o aos eso one ese moo neo woo woo MW Fig. 155. ,
Cold-leg flow - loop 31.
e e
-132- ,
9 sete
~ sees e990- . ,
. sees sAst
,see- PaRAMETEC 1 . -
muum =C2 esse
~4400 h see- ... l e-- . ..
.l
.oes.- .-.mo
~
-me . . . .
o see ato see eeo see see use see MM Fig. 156.
Cold-les flow - loop B2.
ese , ,
00.
SPD- .
see ase-a .
mutAMCTWC t
' too g ses- mWhaCDec 2 .
E
's , ~ . s '- --- .-ee g<
Ises- el-ee- -
's . ee
..s,e
.,0. I No 3F5-no e sie see see see see see MW Fig. 157.
Cold-leg liquid temperatures - loop 51.
e
-133-
'\
l ao , , , , , , .
soo 95- -
880 Ste- gagg -
PWinesc .
g tas. PWimCBC 2 .
E s
' ' N s , * - -. -'**
lsee-~
g me.
dos g
W g .s 6-. --3ee syn. .
330 380 , .
o ase .a ese ese moc neo woo see MW Fig. 158.
Cold-leg liquid temperatures - loop B2.
See , ,
soo geo. .
-vo east see. PWtAhGRC 1 -
too g PWladflWC 2 se g ,
30
. .e0 soo-
\.s.%.
w._ --
me- -
- ' ~
3 440 , , , . .
, e ase me see ses nec see woo neo
( MW Fig. 159.
Cold-les liquid temperatures - loep A1. .
l l
-134-9 eso , , , ,
- o. .
.wo SG .,
soo- PWtamCRC l . -
M g PWLAdTitC 2 soo- -. s.:
330- . eso
.\
soo-ss
.~.N. w._ - - -
. .e.
s, .,
ao. ~..~ _.,
3a
. ,,a deo , , , , , .
e aso eso eso soo soo sco neo soo M@
Fig. 160.
Cold-leg liquid temperatures - loop A2.
om , ,
Glee- .
em PWeautTmc1 PWtmenec 2 mar- .
E a .
I I
An- . .
4 es- . .
l
- 4 80 , , , , .
e aos ao see ese see soo mes see MW Fig. 161.
Candy-cane vapor fraction - loop A.
i l
1
-135-sas , , . . . .
am. .
mast *~
PnRAaptC 1 PmRAncaC 2 ass- -
I amo -
l I .aan. -
- e. .
- -ses . . . . .
e me me see see neo noe men see MM Fig. 162.
Candy-cane vapor fraction - loop B.
a east PanandtTmc 1 m e- #9 PmRahCBC 2 .
m 8- .
n E
... ,,- ..a .
,/
/ a
- s. ,
I a ./
s- .
~
/'/ ( % . ,,,
e.
N. -' . .,
.s . . . .
e ao ao see soo woo non woo eso ;
MM weser w (o4 Fig. 163.
Pressuriser water level.
I
\
. . ~ . - - . .. -.
-136-g, e THETA = 1
- THETA = 2
~
. THETA = 3 550 = META = 4 .-
- THETA = 5 i h
- THETA = 6 -510wp
-@~
w w a: a:
S -470 $
l s 510 -
w 5
w D. O*
E -430#*
e- 490-
- g. - 390 l
450 , , , , , , , , , 350 0 150 300 450 000 750 900 1050 1200 1350 1500 Fig. 164.
Downconer liquid temperatures (base case) at vessel axial level 6 (all azimuthal sectors).
un ,
e THETA = 1
- EETA = 2 -564 I
- THETA = 3 560 -
,33g7, , 4
+ THETA = 5 F ' THETA = 6 - 524 wft- -
l
' w 540- u c: a:
a o 4 4 5
- c. s20-
-484g c.
E
- r N 500- ~444 480 , , , , , , , , ~, t 404 0 100 200 300 400 500 000 700 800 900 2000 1100 Fig. 165.
Downconer liquid temperatures (parametric -
case 1) at vessel axial level 6 (all azimuthal sectors).
O
- . .. .. l l
l
~
I l
-137-TABLE IIII ,
TBV EVENT SEQUENCE ,
Parametric Case 1 Event Time (s) t l 0-209.1 1-10. Same as base case
- 11. Loop-A EW valve shut on high SG liquid level 290.0
- 12. Loop-B EFW valve shut on high SC liquid level 460.8
- 13. POD opens 975.0 operated on both stesa generators. The closure of the loop A EFW valve limits the flow through the emergency-feedwater header as shown in Fig. 145. 'he remaining flow through the header comes through the loop-A SUFCV. The loop 4 energency-feedwater header flow (Fig. 146) decreased to zero shortly after
~700 s with the closure of the SUFCV by ICS action. Downconer liquid temperatures for Case 1 are presented in Fig.165 at the top exial downcomer level (just below the cold-leg nozzles). At the end of the calculated transient (1015 s), the minimum temperature was -482 K. The base case minimum temperature at the same time was -471 K. The slightly increased downeomer temperature for Case 1 was caused by the reduced heat transfer to the loop-A steaa generator secondary with its reduced liquid inventory.
iii. Parametric Case 2. The specifications for this case differed from the base case as follows: the SG-A secondary liquid-level control did not fail;,
restart of one RCP in each loop was permitted on attainment of 75 F subcooling; and throttling the HPI was permitted on attainment of 75 t 12.5 F subcooling.
The event sequence for this transient is presented in Table XXIII. Events 1-10 were identical to the base case. Cne RCP in each loop was restarted at ~383 s .
after a 30 s delay following the subcooled-sonitor trip. The HPI was throttled at -485 s after a second subcooled-monitor trip at 75 12.5 F subcooling.
Results for Case 2 are presented in Figs.138 through 163 and may be compared directly to the base case. Case 2 displayed significant differences from the base case. A major consequence of throttling the HPI was that primary-system repressurist. tion did not ocent and thus the FORY did not open. The absence of repressurization is seen in Fig. 133. Restart of the loop-Al and -El RCPs can be observed in Fige.153 and 155, respectively. Operation of the RCFs l
l
. _ . -. . , - , , , - . , n .. -
. ~ -..-- -. . -. .. . . . . . . _ . . . .. . . . . _ . . . . _ . . . . - _ _
L
-138- !
TABLE IIIII ,
TBV EVENT SEQUENCE Parametric Case 2 Event Time (s) 1-10. Same as base case 0-209.1
- 11. Loop-A EFW valve shuts 290.2
- 12. Restart RCPs in one loop after subcooling- 383.5 monitor trip
- 14. HPI throttled after subcooling-monitor trip 484.7 induced a reverse flow through the cold legs with non-operating pumps. The influence of. RCP restart on heat transfer to the loop-A steaa-generator secondary can be observed in Fig. 139. The pressure increased with RCP restart
(-383 s) and remained higher than the other cases for the . remainder of the calculated transient. The same influence can be seen in the loop-A steam-generator-secondary water inven;;uy (Fig. 149) as a marked reduction in the rate-of-inventory increase caused by increased evaporation of inventory with RCP operation. A different trend was observed in the loop-B steaa generator-secondary water inventory (Fig. 150) with the Case 2 inventory generally exceeding the base-case inventory. This suggests decreased energy transfer to the loop-B steam generator. The summed heat transfer to the A- and B-loop steam
~
generators is less than in the base case, and this is evident in the primary-system temperature. Downconer liquid temperatures for Case 2 are presented in Fig. 166 at the top axial downconer level (just below the cold-leg nozzles). At tha and of the calculated transient (1500 s), the minimum temperature was
-491 K. This was ~33 K higher than the base case. At 1015 s the minimum temperature was -499 K which compared with -482 K for Case 1 at the same time.
- d. Conclusions. The response of the Oconee-1 plant to a secondary-system depressurizatica transient was simulated using TRAC-PF1. The transient studied was failure of one bank of turbine bypass valves (two valves) to closa af ter initially opening following reactor and turbine trips. The base-case transient included additional - failures caused by failure of the level control in the affected steam generator, no operator restart of the reactor-coolant pumps, and
- - - ~-e- -m,,
l
-139- l l
I no operator throttling of the HPI system. A minimum liquid temperature in the '
downconer of ~458 K at 1500 s was calculated. If correct operation of the steam generator level control, operator restart of the reactor-coolant pumps, and throttling of the HPI flow were assumed, the pri.sary system did not repressurize, and a minimum downconer liquid temperature of ~491 K was calculated at 1500 s.
- 2. Two Banks of Two TBVs
- a. Introduction and Summary. This case differs from the prevf.ous case For this (Sec. III.D) by assuming two banks of TBVs fail instead of one bank.
study, the performance of the Oconee-1 plant following a secondary-side depressurization was predicted. The base case analyzed was the f ailure of two banks of TBVs to resent af cer initially opening following rer ct.or and turbine trips from full power. Additional failures assumed for the base case were failure of the level control in the affected steam generators, f ailure of the operator to restart the RCPs, and f ailure of the operator to throttle the HPI system. The lowest downcomer liquid temperature (~445 K) and hence the smallest margin against the NDT limit was again calculated for the base case.
Repressurization of the primary system to the PORV setpoint was also predicted Geo
~ -ST 3 l
- THETA = 1 J
- THETA = 3 640- . THETA = 3
- THETA = 4
- THtTA = 5 m
gsoo-
- Tam m . _g M W -
a: m
- ~ -362.3 E E
! -- - ug aso- .ma3 i
i
- i .
i . N 0 1000 2000 3000 4000 6000 scoo tooo l
l Fig. 166.
Downcomer liquid temperatures (parametric case 2) at vessel axial level 6 (all azimuthal sectors).
t
! l 1
-140-l l for the base case. For 'the parametric cases examined, a reduced number of failures were taken, and a greater margin against the NDT limit was calculated.
The same model used for the TBV transients described, previously -
(Sec. III.D) was also used for this study. The significant features of the TBV l failure transient are (two banks of two TBVs):
1
- 1. Reactor and turbine trips cause the TBVs to open.
- 2. Failure of two banks of TBVs to close causes a secondary-side depressurizatica through both loops.
- 4. The operator does not restart the RCPs.
- 5. The operator does not throttle the HPI.
Two parametric cases were also calculated. The steaa generator liquid-level controls in the affected loops operate correctly in Case 1. The steam generator liquid-level controls also operate correctly in Case 2. In addition, operator actions to restart the RCPs and throttle the HPI are permitted if the primary system subcooling-monitor trip points are exceeded.
I b. Results.
l
! i. Base Case. Table XXIV presents the calculated event times for the base case. Following the reactor and turbine trips, the TSVs closed (0.5 s),
secondary pressures rost, and the TBVs opened for the first time at -4 s. The secondary pressure peaked and then decreased, but both banks of TEVs (four valves; two on each line) failed to resest. Continued flow through the TBVs resulted in a secondary-side depressurization.
The pressurizer pressure is presented in Fig.167. The PORY opened at
~1175 s when its pressure setpoint was exceeded. The PORV then cycled for the remainder of the calculated transient to maintain the primary-system pressure at or below the PORY setpoint.
The loop-A and -B secondary pressures are shown in Figs.168 and 169 respectively. The depressurization characteristics for the two loops were nearly identical. MFW flows for both loops are shown in Figs.170 and 171, respectively. Per the problem specifications 3 the ICS failed to run back MFW to th. cr-a and thus the flows did not decrease until the MFW tripped on high SG B aix , level at ~91 s. A higher mass flow was predicted for loop B prior to tripping the MFW pump because the MFCV area under ICS control was -4% greater D
-141- -
I
- I am . . . . . .
1
. asso t
/ '
l . .
=- - '
l a me-9M mutadtfisc 1 l -
/ .sooo mutedE"IloC 2 / .-
s' I me-me.
,./
/ - '8e
-noe -
l nso ee.
. .mos
- n. i
~
_ ,. s "- -ne
- m. . .-
soo
. m o
zoo ao eso aos moo noo a aos MM Fig. 167.
Pressurizer pressure.
So . . . . . . .
.nso es- -
3 .
BE .
m, -mee mutMdEffeC 2 se- -
I se-
.eso m.- - -*oo as- -
.mo
- m. N. .
. . N .'.. -- . .... ... ~ _ % , .
e- _
e . . . . . . . o 1
e me me eso aos mee noe mee see i .
MW Fig. 168.. .
Steam generator-secondary pressure - loop A. '
9 - ,.
, ~. . .. . . .. . - _ . . .--. .
.. j i
-142-g u 5 8 3 8 3
. see 3 -
m- .
as .
-see gg, PasteCTsuc 2 .
n- -
I eg.
ene 1'.
4
- ~***
as- -
- m. -
..N**** . % .%
, , - - . . . . ,,--.. % .se, e- -
e . . . . , . . e e ase a ese ese noe ese wee neo lhE W j Fig. 169.
l Steaa generator-secondary pressure - loop B.
me , ,
.mee see- -
g nee ses- PaftmEfac t -
. Pastusmc2 ,,,,
4 me. .
g
. see S see. .
Ges ase- ,
..e se- . -
ano
~-
e- -
- ---- w ~. - e 1 -me '
--see e aos me ses ese noe see wee use ftE W Fig. 170.
Main-feenvater flow ~ loop A. -
l l
e 4
-~~.
-143-TABLE K11V TBV EVENT SEQUENCE Base Case Event Time (s)
- 1. Turbine and reactor trip 0.5
- 2. Turbine stop valves close 0.5
- 3. Loop-A TBV opens (f ails to resent thereafter) 4.1
- 4. Loop-B TBV opens (fails to reseat thereaf ter) 4.3
- 5. HPI begins following trip on low systes pressure 87.5
- 8. Feedwater realignment trip 117.4
- 9. Main-flow control valves trip 117.4
- 10. Energency feedwater pump on 147.0
- 12. PORV opens 1175.7 than in Ic >p A. Loop-A and -B MFW temperatures are presented in Fig.172 and 173, respectively.
EFW flows through loops A and B are shown in Figs.174 and 175. EFW flows were initiated at ~150 s. Because the EFW 1evel-control system failed, the EFW flow continued until the end of the calculated transient. Loop-A and -B EFW temperatures are presented in Figs. 176 and 177. Before ~150 s stagnant conditions prevailed, .and the liquid temperatures were near the initial conditions. The EFW flow induced a flow of liquid through the SUFCVs that mixed
[ with the EFW before entering the SCs. The temperature rise beginning near 200 s reflects mixing of these two flows.
The water inventories in the loop-A and -B SG secondaries are shown in Figs. 178 and 179. Although the fill characteristics were similar SC B filled more rapidly before the MFW trip because, as previously discussed, the loop-B MFCV opened to a larger flow area by the ICS. With no liquid level control, the SG secondaries continued to fill and oscillations developed as the liquid 1 9vel reached ~12.'4 m (top of the generators).
l .
l
l -144-(
l I
- me . . . , , . .
- L wee o a. -
g ase see- do Pamtmc -
i .
m2 , ,,,
g age. .
See 5, 33 -
c eee E i
see- -
no- .
ase
- e. . '
.A,,^ -
_w . .e
-me e see ese see see see oss tese neo MW Fig. 171.
Main-feedwater flow - loop B.
ano , , , , , , ,
i so- -
me see- PnAAndniaC 1 -
g -
_ PnAA&ETNC 2 ase as-
- E l '88 t %A., .
I i gg .
se.
., m , m
.ag ase
% .m ass. .
3e es. .
.am des , , , , , , ,
,e ase aos ese eso use soo sees see MW Fig. 172.
Main-feedvater liquid temperatures - Icop A.
e G
+40*
i i
-145-l me , , , , , .
.mo on- ,
.ee -
aos- PutAdliac1 .
MWIMGRC 2 4ae g -
E
.e- .
- s, .
w.
N.
I ese me- - s. .
.,_ . g es-
% h. .
me-3ee me. a4 48e , ,
see e see me eso ese see see mee MW Fig. 173.
Main-feedwater liquid temperatures - loop 3.
see , , , , ,
.see see-
.e BASE mutmouc1 .
es- 2
.p. ew ~.... .. SL u'~ .
g l me-
.see g
so- ,,
j e- m I i
- O , , , ,
see see mee see e see me ese see its W Fig. 174.
Emergency-feedwater flow - loop A. .
l l
. .. . .- . . . . . -- . ~.
-146-l l
me . . . . . . .
see see. .
l as
' i aAE -
i mutmamec1 M,
.ng*W""%2 =s. - *
- g .
.see
$ es. E c .
ase 2 1
1
.e I
- e. J . e
-Go , . . . . .
,e see me eso ese use see sese mes MW Fig. 175.
Emergency-feedwater flow - loop B.
as . . . . . . .
gee. .
es.
F(~ m
. me g .e . .
. ..e E
see l *as- *'
. ~
O so- f \. N -
-ase g
ase-ase l5
.as ass- -
- y. me ase- -
es ss, aos me ese see eso ses seen ese MW Fig. 176.
Emergency-feedwater liquid temperatures - loop A.
_.y. ,
l s . w.
l I
-147- l 1
= .......
. g "
as.
- N ,,,- -- % .aes
\ ms
~
g "
g,
~~
E
.e.
N ..
I I 'N ~ .
I m.
- i
- g. 153
,ee. .
.e m"8 ,
4 M M ein sie isso M ieio ines MW Fig. 177.
Emergency-feedwater liquid temperatures - loop B.
- essee . , , , , , , ,
i aAst i
- Pansw1 mci .,eoeee
, m as. PnA4WTisc 2 .
/ -insess asene. /. '
.. I' .
1 2 m-
/vW,- .
- o .,oooon l- - !
/ .
.eeeen g g
asses. , .
~
/ eenee sense. .
.neee mese e see sie sie me sie eie m een aio ,ee, ,no wp Fig. 178.
. Steam generator-secondary inventory - loop A.
, - . . _ , , _ - . _ , , , , . - _,,_._ ,,, .,..an,,
-148-Mass flows through the primary-loop hot legs are shown in Figs.180 and 181. Following the RCP trips at -117 s, the flows coasted, down and natural circulation flows through both loops were established. Similar phenomena vote observed in the cold legs sa shown in Figs. 182 and 185. The corresponding loop cold-les temperatures are presented in Figs.186 through 189, respectively. -
Vapor fractions in the loop-A and -B candy-cane sections are shown in Fige.190 and 191, and it can be seen that no voiding occurred during the transient. The pressurizer water level is presented in Fig.192.
Downconer liquid temperatures for the base case are presented in Fig.193 at the top axial downeomer level (just below the cold-les nozzles). At the end of the calculated transient (1320 s), the minimum temperstare was ~445 K. The minimum downconer temperature for the base-case failure of one bank of two TBVs was -465 K at 1320 s. The lower temperature predicted for the base-case failure of two banks of two TBVs was the result of enhanced heat transfer to two, as compared to one, steam generators.
ii. Parametric Case 1. A single specification was changed for this parametric study. The loop-A and -B SG 1evel controls following EW activation were assumed to operate to maintain the secondary-liquid level at or below 6.2 m. It was assumed that the operator does not restart the RCPs and does not throttle the HPI. The event sequence for this case is presented in Table XXV.
1 The event sequence was identical to the base case through event 9. At -147 a the loop-B EW valve shut on high SG-B secondary-liquid level. The loop-A EW valve shut on high SG-A secondary-liquid level at -373 s. The POEV opened ~13 s early in the Case 1 because of reduced heat transfer to the two SGs.
l TABLE XIV TBV EVENT SEQUENCE Parametric Case 1 Eveg Time (s) 1-9. Same as base case 0-117.4
- 10. Loop-B EW valve shut on high SC B liquid level 147.0
- 11. Loop-A EW valve shut on high SG A liquid level 372.6 ,
- 12. PORY opens 1062.1 1
l i
-149-
- I esoee , . . . . . , , ,
sASE pammouc i . moose esose- ,ca 4 PnAmpeC 2
. -messo Sesse- ,
. [ \ .wsome ..
g moes- g .. . . ,
,g /
.aosso asaso- L
. . moose f
~
. .soooo aeose- -
. .aosso soone- -
. .maae noeo -r . . . .
ano soo ao soo soo no soo saa neo veo e me MW . ,
~~
Fig. 179.
Steam generator-secondary inventory - loop B.
1ees , , , , , , ,
~
, sees.
l .
.eese
. . i BASE . naso I PRAMETMC 1 tese- I PnAMETMC 2 -
.meos
- m. .
. seos -
, , ~ . ,
. soos asos- ,
. .mee mes- . ,
. esse
^
e , , . . . . . e e age me ses ese see see sees see MW fig. 180.
Bot-leg flow - loop A.
,c
-150-
- U v v s i s a v
( [ . teese sese. .
. r g .nese sese. PutAdiac 1 -
. mtAmmc2 ,,,,, , , ,
$ asse. - g g . .sese g
, g
.soes c mes- ' -
l ese- . ,
.asas 1
i __ _ _ _ ,
! e. . . .e
-mee e age me eso see use neo ises aos MW Fig. 181.
Bot-leg flow - loop B.
mes , , . . . . .
. - '"888 g m sese- e nutAdmuc t -
. MtAduc2 , ,,,,
/
as . . g ooen g ,
m- -
, oose h I -. .
e
._ i mes- . -
. nase
- e. . . .e
' ~
-Wee . , , , , , .
q aos me see ese neo see use see MN
! Fig. 182.
Cold-les flow - loop A1.
-151- .
esse . . . . . . .
mese f ..
ease.
- wJo .
Em -
asse- PWtAEDec 1
- P8JtAnfimC 2 este asse- ,
- 1. .
..se i I .. .
. .e i
- e . -asse
-meo
- ..eeee
-asse . . . . . . .
see e ase me ese eso see ese mes MW Fig. 183.
Cold-leg flow - loop A2.
mes . . . . . . .
- - -- weso
,~. -
gg asse sese- . mutAmtinc t -
N2 sees moe.
. 2 8
.e g m- -
esse H- . .mee e
ese- -
- e. .
-asse
-see
~
see
- .e ase see ese see eso use use
, MW l
t Fig. 184.
Cold-leg flov - loop 31.
D
-152-eene . . , , , , ,
~
f .
aee. .
. esse em -
noe. Futucac t
- PMthcluc 2 sees
'~
nee- ,
5._. .
. sees 3
i s. .
. .e i
. . -seos
-mes
. .. ease
. 4ee , , . . . . ,
see e see me ese see use see mee 1mK @
l Fig. 185.
Cold-leg flow - loop B2.
~
eso , . .
.see
.see
. east .
PWIAMCTIMC 1
- ~"'
E -- -
E
.e o Me<
N ..
sea.
- -ene
.' \
dee. %._ .
,' % .. ' ~ ' w .. -
. ass ese-
-ase 4e0-se ese -
e see ese see see see use mee see MW Fig. 186.
Cold-leg liquid temperature - loop A1.
I i .
l e
f i
-153-eso , . . . . . .
sen
~
tes-ses 88 sASt -
PnRAmpuC 1 -
l
$ g, .
E ggg.
. - ato S .
N.,N-\'N--
-aso
~~ , m .. -ses e,s.
. b*.4 ggg. *
. ss e aso aio soo see moo aos we moo MW ..
Fig. 187.
Cold-leg liquid temperature - loop A2.
. eco . . .
ses aso-seo
,, . east .
PARA &OhC 1 PAAAtGBC 2 . ses b w. -
E l .
aeo ete geg. -
s -e w.._.~~m.
s.~'
- %~ ,
. me a o. -
- se "o u 4 m n m m m m M@
Fig. 188.
Cold-leg liquid temperature - loop Bl.
-154-em . . . . .
.e. . . ,
east PWtancac 1 I PWtanceC 2 I_ .
I .e. .
4.s6 -
1
-ase . . . . . . .
e ase me ese ese see use sees see MN Fig. 189.
Cold-leg liquid temperature - loop B2.
cae . . . . , , .
Lee- .
east PWeautfac 1 PWlascoc 2 am- .
See . .
I -em- -
-ass-
-48e , , , , , . .
e aos me see ese see see mes see MW Fig. 190.
Candy-cane vapor fraction - loop A.
~~
l
. I
~~
j
-155-I e
mas . , , . . . .
su.
< east i Peanosc 1 1 PMAnceC2 mes-l l
a.es
- I .
i
.su - .
=&as , , ,
e asu ao see ano see see mee see M@
Fig. 191.
Candy-cane vapor fraction - loop B.
s . . . . , , ,
. .m e- .
- 3Ag M PnAmdETItc 1 e- . .m I g e.
. .m .
m e 4
g
. .e s
['%,%. .e e- .
'd ~ . .e
. ..e
-a ,
ses me ese see see ses ses see lhE 64 Fig. 192.
Pressurizcr water level.
\
-156- I Results for Case 1 are presented in Fige.167 through 192 and may be compared directly to the base case. The general trends of Case 1 were similar to the base case. He major differences appeared in the secondary sides of loops A and B and were caused by shutting the EFW valves on high SG liquid levels. The reduced EFW flow affected the primary side also. Compared to the base case, the pressuriser pressure increased more rapidly to the PORY setpoint .-
as shown in Fig.167. This was caused by reduced primary-to-secondary heat transfer associated with reduced loop-A and -B steaa generator secondary inventories (Figs. 178 and 179).
Downconer liquid temperatures for Case 1 are presented in Fig.194 at the top axial downconer level (just below the cold-leg nozzles). At the end of the calculated transient (1062 s), the minimum temperature was 465 K. The base case minimum temperature at the same time war ~453 K.
iii. Parametric Case 2. The specifications for this case differed from the base case as follows: the SG A and B secondary liquid-level controls did not fail, restart of one RCP in each loop was permitted on attainment of 75 F subcooling, and throttling the HPI was permitted on attainment of 75 12.5 F subcooling. The event sequence for this transient is presented in Table XXVI.
Events 1-9 were identical to the base case. The HP1 was throttled at ~421 e after a subcooled-monitor trip at 75 12.5 F. One RCP in each loop was restarted at ~517 s after a 30-s delay following a second subcooled-monitor trip.
Results for Case 2 are presented in Figs.167 through 192 and say be compared directly to the base case. Case 2 displayed significant differences from the base case. A major consequence of restarting the RCPs and throttling the HPI was that primary-system repressurization did not occur, and thus the.
PORV did not open. The absence of repressurization is seen in Fig. 167.
Restart of the loop-Al and -B1 RCPs can be observed in Figs.182 and 184, respectively. Operation of the RCPs induced a reverse flow through the cold legs with non-operating pumps. The influence of RCP restatt on heat transfer to the loop-A and -B steam generator secondaries can be observed in Figs.168 and 169, respectively. The pressure increased with RCP restart and continued to increase to the end of the calculated transient. The increase of 'the primary systmt pressure (Fig. 167) was terminated after the HPI was throttled at ~420 s.
The priar.ry pressure decayed rapidly to the accumulator setpoint of 4.168 MPa at
~565 s. Water at 305 K was then injected into the primary for ~20 s. The primary pressure then increased until it was turned around by the increased heat I transfer to the secondaries by RCP operation. The accumulators again discharged 1
l
e i
-157-l TMu mI ,
I TSV EVENT SEQUENCE l
Parametric Case 2 1
'~
Event -
Time (s) 1-9. Same as base case 0-117.4
- 10. Loop-B E N valve shut on high SG B liquid level 147.0
- 12. HPI turned off 421.0
- 13. Restart 7. cps Al and B1 after subcooled 517.0 monitor trip
- 14. Loop-A accumulat.. r begins 'di; charging 565.5
- 15. Loop-B accumulator begins discharging 565.5 near the end of the calculated transient, thereby increasing the rate of downconer temperature . decrease. This was the caly TBV transient that experienced accumulator discharge and it markedly influenced the extrapolated downeozer. temperatures at 1200 s.
Downconer liquid temperatures for Case 2 are presented in Fig.195 at the i top axid downconer level (just below the cold-les nozzles). At the end of the calculated transient (1500 s), the minimum temperature was 467 K. At 1062 a, the minimum temperature was 477 K, which compares with 465 K for Case 1 and 453 K for the base case at the same tf e -
- c. Conclusions. The respe-m sf .i a, Oci.;.ee-1 plant to a secondary-system l depressurization transient was Q:uty tatug TRAC-PF1. The ' transient studied was failure of tuc banks of tuebine-bypass valves (four valves; two on each steam line) to close after initially opening fol.1owing , reactor and turbine trips. ' The baae-casa transient included additic .a1 failures caused by failure
'of the level control in the affected steam generator, no opeystor restart of the reactor-coolant pumps, asd no operator throttling of the Ell mystem. A minimum.
liquid temperature in the downcoeer of 495lK at 1020 s was calculated. If correct operation of the steam generator level control, operator restart of the reactor coolant puap's,3rA ' throttling of the HPI flow were ' assumed, the primary
$ 4}. .
- 't 4 <
-158a se0 eTHEE4 u 1 *672 a 1
- THETA = %
500- ,9ygg, , ,
. THETA = 4 -532 l _0 . THem = .
w 4<
. tHEm = .
-492 w F '
as 680- $
-452 E
a.
M- E a.
- 412 400-400- - 372
~ ..
M
- 440 . . . . . . . . . 332 f 0 15 0 300 460 000 750 900 1060 1200 1360 i .
Fig. 193.
Downconer liquid temperatures (base case) at vessel axial level 6 (all azimuthal sectors).
Se0
. THETA = 1 -S72
< a THETA = 2 000 -
+ THETA = 3
= THETA = 4 "E o 540-
- THETA = 5
$8
- THETA = .
w -492 w$*
$680- $
g .
~452 4 6600- E
- a. a.
5
- 412 e.
400-
-' ~
440 . . . . . . . . 332 0 ISO 300 480 000 750 900 1060 1300 1350 l
Fig. 194.
Downconer liquid temperatures (parametric
.. case 1) -
vessel axial level 6 (all azimuthal sectors).
l l
l l . ..
-159-500 e THETA = 1 -572 J
l
- THETA = 2
. THETA = 3
. THETA = 4 -532 540-
- THETA = 6 h
- ~
- THETA = 4
-492*h Sao- $
-452 800- E a.
>g. - 412 4eo. - 372 440 . . . . . . . . . 332 0
15 0 300 450 000 750 900 1050 1200 1350 1500 Fig. 195.
Downcomer liquid temperatures (parametric Case 2) - vessel axial level 6 (all azimuthal sectors).
system did not repressurize, and a minimum downcomer liquid temperature of
-472 f.was calculated at 1320 s.
E. Hot-Leg Break LOCAs
- 1. 2 in. Break
- a. Introduction and Summary. This report presents the Oconee-1 plant response to a 2-in. break in the surge line midway between the pressurizer and the riser of the candy cane. Following the initiation of the break, the reactor' end turbine trip from full power. Reactor decay heat was specified as 1.0 times the ANS standard. For this transient calculation, the ICS and all key system components were assumed to function correctly. The only specified operator l
action was the RCPs trip 30 s after HPI actuation. Two cases involving HPI throttling to system subcooling were investigated. One calculation investigated the ~ effects of HPI throttling to 42 2 12.5 K subcooling and. the other i investigated the effects of no HPI throttling. The throttled HPI case was not run because the subcooling margin was never achieved.
TRAC calculated a relative minimum in temperature at approximately 1000 s into the transient; the pressure at 1000 s was -6.2 MPs. The temperature and pressure were both decreasing at the end of the calculation at 3760 e and had values of -450 K and ~2.1 MPs, respectively.
I I
-160-Because the reactor should not have been tripped until the low pressure recetor trip setpoint was reached, and because the calculation was not run to 7200 s, it is recommended that the RELAF5 calculations be ,used for the ORNL study.
- b. Model Description and Assumptions. A corplete description of the primary-side, secondary-side, and ICS modeling can be found in Section II. The . .
steady-state operating conditions are also presented in that section.
The 2-in. break LOCA specification containing the initial conditions, event sequence, and assumed failures is presented in Ref. 3. The TRAC transient event sequence is presented in Table IXVII. Because the ICS and major system components were specified to function correctly, it was not necessary to make any overriding assumptions.
- c. Transient Calculation. Figures 196 and 197 present the SG secondary side pressures for loops A and B. Following the initiation of the break, the reactor and turbine tripped from full power, the TSVs closed causing a temporary increase in secondary-side pressure. The TBVs for both loops were activated
-4.2 s to relieve the initial secondary-side pressure increase. The initial relief of secondary-side pressure caused a sudden drop in the SG secondary-side inventories shown in Figs.198 and 199. As the SG secondary-side inventories continued to decrease, the RCPs tripped (30 s af ter HPI actuation) and the MFW TABLE XXVII HOT-LEG BREAK LOCA - 2 IN. BREAK SEQUENCE OF EVENTS Event Time (s) .
- 1. Break opens 0.0
- 2. Reactor and turbine trips 0.5
- 3. TSVs close (both loops) 0.5
- 4. TBVs open (both loops) -4.2
- 5. HPI actuation 43.1
- 6. TBVs open/close (both loops) 51.0
,7 . RCPs trip 71.0
- 8. MFW rcalignment 73.1
- 9. TBVs open/close (both loops) 75.7
- 10. Vent valves open ~100
- 11. ICS closes SUFCVs -350
- 12. Candy canes remain voided ~500
- 13. Minimum downr.omer temperature (-470 K; 6.2 MPa) ~750 *
- 14. Loop esci11ations begin ~1200
- 15. Accumulator injection begins ~1750
- 16. End of calculation 3670 9
=
mr wm w w -w ., ~ - _ _ _ . , , . _ _
-161-9 so . .
- WO ,
gg. .
i 500 n.
. see T
- w. .
3
--20 I
190--
Me m.- .
n . . . . . . .
'3D*
O 900 1000 300 2000 3500 J000 Je00 4000 MM
.2 Fig. 196.
i Steam generator secondary-side pressure - loop A.
t u 6 s
- -noo 00- -
500
- n. .
I.- .
-tes l
1 40-~ -'88 y-- --80 ao . . , , . , ,
see i o see see soo sooo see asse - anos esse j
MW ,
1 Fig. 197.
Steam generator secondary-side pressure - loop 3.
i l
e
- - . _ , . . 4 . - -
-162- .
soooo . .
. -messe Mfee- -
. eees .. !
e,ees. -
l y .
-neeee g sesos. -
- -Wooes 40000- -
B . -e@oon 5
..soes
.ees.. . e.5 toooo . . . . . . .
O see teen 1eco 2eco 3eco 3o00 ases deco M (s)
Fig. 198.
Steam generator secondary-side inventory - loop A.
enoso .
. -15000
.eoogo
,,,,o.- .-usen 2 . -noose a -
M. ~
. .seoso
... ..seeme
. -7ecce 3m,00 -
-seems seese . . . . . . .
--^eene -
e see neo mes sees noe 3ees meo aseo M (s)
Fig. 199.
Steam generator secondary-side inventory - loop 3.
.9 - _ , _ . _ ._ _
-163-was realigned to the EFW he,ader. The loop-A and -B MFW mass flows and liquid temperatures are presented in Figs. 200 and 201. The realigned mass flows reduced the secondary pressures about 0.7 MPa between ~100 s and ~187 s and increased the SG inventories. The ICS continued to supply MFW to the SG upper header until ~350 s. At this time, the SUFCVs were closed by the ICS based on the 3G inventory. Realigned mass flows and liquid temperatures are shown in Figs. 202 and 203, respectively. The EFW pump was not activated in this -
transient as it was not needed.
'Ihe pressurizer pressure and water level are presented in Figs. 204 and 205 respectively. The primary system depressurized rapidly until HPI actuation at -43 s ar.3 remained above ~6.0 MPa until ~1200 s. From 1200 s to the end of the calculation the pressure decreased steadily to ~2.1 Mrs. The pressurizer water level also dropped rapidly, and was zero by ~50 s. Figures 206 and 207 present the break mass flow and void fraction. The break mass flow was greater than ~100 kg/s for ~1000 e into the transient until the primary slowly voided.
The mass flow decreased to ~70 kg/s as the void fraction increased to -0.8. The candy-cane void fractions in Fig. 208 show that the loops did not refill in the course of the calculation.
Mass flows for the loops A and B cold-legs are shown in Figs. 209 and 210, respectively. As the primary-system flows began to decrease following the RCPs trip at ~73 s, the HP1 fluid began to flow toward the vessel and fell into the dqwncomer as shown in the cold-leg mass flows and temperature profiles. Figures 211 and 212 present the cold-leg liquid te:nperatures for loops A and B. 'Ih e hot-leg mass flow and liquid temperatures shown in Figs. 213 and 214 reflect the cold-leg response to the HPI. The mass flow and corresponding liquid temperature fluctuations that occurred af ter ~1000 s will be discussed later.
TRAC calculated a minimum cold-leg temperature of ~420 K and ~440 K for loops A" and B, respectively. The minimum hot-leg temperature was calculated to be
-495 K at the end of the calculation (3670 s).
An important feature of the B&W PWR design is the vant valves located around the upper plenum of the vessel (level 7 in the TRAC model) that provide the upper plenum region access to the downcomer. Af ter th'e RCPs have tripped, the vent valves are capable of providing a source of hot fluid for _ mixing with cold HPI fluid that may flow toward the downcomer during these stagnant periods in the cold legs. Between ~?OO s and ~1000 s of this calculation HPI fluid did flow toward the downconer and was mixed with the vent-valve flow in the downcomer at the cold-leg junction.
s 8
--+"- wm w - - - - , _ _ _ _ , , , .,
-164-eso . . . -eno (pmod)LDoP A m-. (pse$cor s eso oog. .
- -Me _
gog. -
. .so 3 3 f
m soo- -
.m f
-soo aso. -
no- -'"
l l .__
o-- , 7 .h^rpe , fb -
o o . . . .
o soo woo eso sooo moo sooo sooo ao M (s)
Fig. 200.
MFW mass flows - loops A and B.
i soo . . i (posiOLoor A soo- (p egoopa -
. ., . sos
- m. .- .. , .
E .
s'
.= E m- '. -
\ -
. '.. -.ao see..
l . .
so- .,,7. , - -"
1 soo.. . ., -.uo me-. -
.as doo . . . .
me o son neo soo sooo sono aseo mes w (s)
Fig. 201.
MFW liquid temperatures - loops A and B.
O 6
0
1 l -
-165-M y i
. .m 6dse9L00* B e ,
e
, 7
= = =N ,
. 00
.co h.
-m.
h
-m
.e.. .
-oo. . . .we n
s
..re
~** me e eso neo eso mo moo aseo meo MM Fig. 202.
Emergency / realigned mass flows - loope A and B.
sao , ,
pg. (sedd)LDoP A .
(eseQLoor a see. I ..sso see. .
E m -
- m. .
i sno- .. l l . .
5 . ..
soo- y *.u w .
.o _- -as So , , , , , , .
e eso see eso sees aos aseo asse mee MW Fig. 203. l Emergency / realigned liquid temperatures - loops A and B. ,
i l
6 e - -- ,
- ., , - -- w,e - ,
-166- ,
t 5 3 (seed) LOW A se- 6songuxe s -
esensee me. .
g so- -
h so- -
- e.
e- -
- oneensa
~
3p. -
o . . . . . .
emooooo e eso woo eso sooo see aeos anos sooo M (s) l Fig. 204.
! Pressurizer pressure.
e ,
- s. -
s 4 -
2 g
, s. --n -
I- -
.o I
- g. -
- e. # 4 h -.o
e see see ein nose mee aseo asse 4ees
[ MW Fig. 205.
Pressurizer water level. .
w.
-167- .
soo , , ,
" ~
- g. .
So- -
r -3eo m- -
t',
-~ -
g O --
"~ 5 i #-. I
,"TWjW ;.i
)
- m. .
o-- --o
"" o soo woo soo sono moo mo anoo sooo M (*)
Fig. 206.
Break mass flow.
u , , , , , , ,
- s. .
f
-- I u l l -
- l \ f
- l an- -
- e. .
e s s s s s s s m M 64 Fig. 207.
Break void fraction.
e
-168-u , , , ,
(seM)LDOP A
.. (do*'koo" 8 .
o.o-l -
I t. !!
! !E
- o. - ! !!
It o- -
=o.t . , , .
o soo woo eco sooo asco aooo moo amo M (s) l Fig. 208.
l Candy-cane void fractions - loops A and B.
sooo .
(seed)LDOP A1 - 4 0.
6 (do*)toor A2
.oo . .
oooo 3eco- -
g oooo (
6 8 .
i soo-. -
.nooo i
I o.. l _
-o
, ,4 Qk --anoo l
. moo . . . . , , .
o eso nos eso sooo asco asse sono aseo M (e)
Fig. 209.
Cold-leg mass flows - loops Al and A2.
l l
i l . . . ,
l l
-169- ,
.i sono , , . . . . .
- (paAd)l.DOP 91 -sees W st m e. -
-seso h am- -
.,,,, h I neo-.
.asse i
e.. -
e
-- M
.m .
me me aos o see see eco sono neo w (s)
Fig. 210.
Cold-leg mass flows - loops B1 and 32.
soo .
(psed) LOOP A1 sm- M)LDOP A2
-ses so-. -
gog.. --aae es.- .
W.
z m
_ s
. 1 . .
es- -
. ase o m m m & M S M M MW Fig. 211.
Cold-leg liquid temperatures - loops Al and A2. ,
, + , . . , - - , ,
- - - - ,, , ,e
-170-one , ,
(aand)Lotr 91 -se son.
mm -
-soo .
b w.
E
\{ ,,- -
l.
l s...
es-J y g
g l --
I
(
~
l me- 1 3o
- m. .
3.
e ao soo nao sooo asoo aseo moo . coo M (s)
Fig. 212.
Cold-leg liquid temperatures - loops B1 and B2.
. woo 8
(desh)toop e sooo- -. ,,oo sooo- -
E .nson h m-me l
"' ~
. mo,
- o. nn... ..o o soo moo soo sooo asco aooo moo moo M (s)
Fig. 213.
Hot-leg mass flows - loops A and B.
_____ .. _=
-171-Downconer liquid temperatures at the top axial downconer level (just below the cold-leg nozzles) are presented in Fig. 215. TRAC calculated a minlaum downconer liquid temperature of -470 K at approximately 1000 s. This was followed by heating until 1800 s. At 1800 s, additional cooling 'provided by increased HPI flow as the syster. prasure decreased and accumulator flow once again decreased the liquid temperatures. At the end of the calculation, the downconer temperature was approximately 450 K.
From the results of a previous calculation in which the vent valves were accidently isolated from the downconer (input error), the importance of the vent valves for this particular transient was determined. Figure 216 presents an azimuthal comparison of selected downconer liquid temperatures at axial level 6 i
for both calculations. When the vent valves were modeled properly, TRAC calculated downconer liquid temperatures that were at least 25 K varmer. From the FTS viewpoint, the vent valves were an asset in maintaining " warmer" downconer liquid temperatures. The total positive vent-valve mass flow is shown
~
in Fig. 217.
- d. Analysis of the Loop-flow Oscillations. The loop oscillations that were calculated in the primary system were initiated in the loop B cold-legs.
Similar oscillations have been calculated in other small-break transients for B&W plants.6 The oscillations began after the liquid levels in the primary system decreased to the cold-leg / hot-leg elevations. At ~1100 s, cool HPI fluid dropped into the loop seal from the loop B2 cold leg, and as a result produced a sustained (for ~200 s) positive mass flow (toward the vessel) in loop B1 and a negative mass flow (away from the vessel) in loop B2. Between ~1200 s and
~1250 s a similar occurrence happened in loop A. Cool HPI fluid from the loop Al cold leg dropped into the loop eal in loop A. Immediately following this, occurrence, the loops A and B sass flows and liquid temperatures came in phase and began oscillating. Initially, the loop B oscillations were regular with a period of ~17 s and an amplitude of -600 kg/s. The loop A oscillations were not initially regular.
In order to explain the loop oscillations, several case studies were conducted. These cases included: turn the HPI off before the oscillations begin, turn the core power off with the SG heat trcnsfer on, turn the SG heat transfer off with the core power on, close the break, and renode the middle-to-upper levels of the vessel and SGs. The oscillations presented in I this report were calculated with the renoded vessel and SGs.
- - - ,, - - - - -, -- , ,,..e.,--. . _ , , - _ - . , , , , , , _ , , - - - - - , , , , , - - , , , . . , , , , - - , - - - , , , , - , .
-172- ,
000 ,
(puAd)LD(P A 800 -
.00 Os400P a .
.we
~~
s00-g ,,
l0- 50-
"'8
.ao
.00 430 330 300 2000 25'00 3000 J500 4000 0 800 1000 M (s)
Fig. 214.
Hot-leg liquid temperatures - loops A and B.
I m , , ,
R TH I ,,
s00- 216 -
.226 -.g
- - .236 g . 246 .se Sd- =256 266
- ase
'}l are-
. . 's .. .
900-8 -
I ."i.
g m. f
. i.d
. -39 0 the 1000 900 2000 2500 3000 3e00 4000 ftC (s) ,
Fig. 215.
Downconer liquid temperatures - vessel axial level 6 (all azimuthal sectors).
S
l
-173-see . . . . . . . .
. een as- -
. see
^ t nov.v.umo csoung
"" A v.v.nasc' pm0 3, -me 8 -
E .- 1 m- .
i % lg,i :..e 1
I,,s. . . ..se i(ii i%,3.,.9i,'
h Ps.A IM
- g. gh
- y. .g uV4l .
. ass
- m. .
.see
-
- des . . . . . . .
e see eso ese eso see see sees sees see IllE W Fig. 216.
Downconer liquid temperature comparison for 2-in. break case (vent valves versus no vent valves).
isee . . . . . .
see. .
mes- -
- 3. .
3 -. ,
r N m- g bQ 1
l e. .) .
e see see see sees sees asse asse asse N
Fig. 217.
Total v ut valve asse flow.
i - _ _ _ _ _ _ _ _ . _ _ _ - _ - - - - - - - . - - - - - :- - - - - .-- , - - -
-174-Briefly, the results of the parametric-case studies indicated that the HPI coupled with secondary-to primary heat transfer in the SGs were the forcing .
functions that caused the oscillations to persist once initiated. The oscillations were manometer-type and the columns of liquid that oscillated included the legs of the loop seals, lower half of the SGs , and the vessel.
Af ter the HPI dropped into the loop seal, the elevation head in the loop seal .
increased because of the denser HPI liquid. This increased elevation head pushed the column of liquid in the loop seal down and isp the lower-half of the steam generator. Because of reverse SG heat transfer (secondary to primary),
this additional liquid was heated thuc changing the effective elevation head in the SG. Thus, because of these changing elevation heads in the loop seals and steam generators, the oscillations persisted.
- e. Summary. The Oconee-1 plant response to a 2-in. break in the surge line was calculated using TRAC-PFl. For this small-break transient, the ICS and all key system components were assumed to function correctly. Also, the operators were assumed to trip the RCPs 30 s after HPI actuation. TRAC calculated a miniaua vessel downconer liquid temperature of ~470K at 1000 s.
The primary system pressure at this minimum liquid temperature was calculated to be -6.2 MPs. HPI flow and accumulator injection reduced the temperature at the end of the calculation at 3670 s to ~450 K.
The calculated ~ minimum downconer liquid temperatures never approached the current NDT value of Oconee-1 for two reasons; (a) vent valve flow sizing with the fluid in the downconer region, and (2) calculated loop oscillations.
However, if the calculation were continued the LPI system will actuate as a result of the depressurization. The addition of LPI would probably result in downcomer liquid temperatures that may approach or exceed the current NDT of the Oconee-1 plant. However, this transient may not be important in terms of PTS because the primary system pressure will be quite low when the downconer liquid tiesperature f alls below the NDT limit.
- 2. 4 in. Break
- a. Introduction and Summary.
This report presents the Oconee-1 plant response to a 4-in. break in the surge line midway between the pressuriser and the riser of the . candy cane.
Following the initiation of the break, the reactor and turbine trip from full power. Reactor decay heat was specified as 1.0 times the ANS standard. In this transient, the ICS and all key components are assumed to function correctly.
The only specified operator action was the RCPs trip 30 s af ter HPI actuation.
Two cases involving HPI throttling to system subcooling were to be investigated
l l
-175-for this transient. At, the end of the base-case calculation (~1433 s), the subcooling condition had not been achieved; therefore, only one calculation was .
required.
TRAC calculated a minimum vessel downcomer liquid tempe'rature of ~350 K; the primary system pressure at this minimum temperature was ~1.0 MPs.
- b. Model Description and Assumptions. A complete description of the j primary-side, secondary-side, and ICS modeling can be found in Section II. The - -
steady-state operating conditions are also presented in that section. 1 The 4-in. break specification containing the initial conditions, event sequence, and assumed failures is presented in Ref. 3. The TRAC transient event sequence for this transient is presented in Table XXVIII. Because the ICS and major system components were specified to function correctly, it was at necessary to make any overriding assumptions.
- c. Transient Calculation.
Following the initiation of the break, the reactor and turbine tripped from full power and the TSVs closed, causing an increase in secondary-side pressure. Between ~17 a and ~92 s, the TBVs for both loops functioned normally and relieved the increases in secondary-side pressure. The SG secondary-side pressures for loops A and B are shown in Figs. 218 and 219, respectively. At
~300 s the secondary pressures in both SGs remained momentarily constant just below 4.0 MPa following the closure of the SUFCVs based on the increasing secondary-side inventories. Af ter 400 s the loop A secondary-side pressure decreased at a much faster rate as the primary cooled the secondary. This heat-transfer mechanism lowered the loop-A secondary pressure and produced a larger loop-A secondary-side inventory as a result of coadensation. Figures 220 and l 221 show the loop A and B SG secondary-side inventories. l 1
Loop A and B MFW mass flows and liquid temperatures are shown in Figs. 222-and 223, respectively. The SUFCVs continued to deliver feedwater to the lower SG header until 47 s when the MFW was realigned to the SG upper header. The MFCVs were closed by the ICS ~10 s into the transient. Figures 224 and 225 show the loop A and B realigned mass flows and liquid temperatures. The EFW pump was not activated in this transient as it was not needed.
Pressurizer pressure and water level are shown in Figs. 226 and 227, respectively. The primary system pressure fell very rapidly until 'the HPI was I
actuated at ~ 17 s. The system pressure was slaost stabilized af ter ~800 s at l 4.0 MPa as the HPI mass flow approached the break mass flow. The pressurizer emptied immediately and was never refilled because of the relatively large break. Figures 228 and 229 present the break mass flow and void fraction.
I w - - - - - - - - w. w w- -_ - - - - - - . - ,,. .- ~_ ,--gp-.,.--yn-y-.g - - . u--
-176-
,o , , ,
-n00 '
80-so-. .
,,nn '
I .0 800
-800 140-~
- m. .
-400 30- .
-200 e- .
0 0 0 250 450 600 aN M N TIME (s)
Fig. 218.
Steam generator secondary-side pressure - loop A+
es _ , . - ' ' '
b .nco nk .
7s- g .
3 ,4 '
- 00 1,- ~
. 00 I eo. .
I gg. -
.-300 s0-
-700 46-
-600 40- -
35 g 4 , a 6 O O O "
11WE (s)
Fig. 219.
Steam generator secondary-side pressure - loop 3+
e G
9
. - - - - - - - - - - +- - _ _ - _ - _ _ . _ - . _ _ _ _ . _ _ _ - _ -
-177-soooo . . . . . . .
9
- m. ,
-Soooo ggg. -
. -tescoo
- m. -
m.
- --9o0oo g , , ,
E -
2 e 200o- ,
,,m M-
. .. coo.
Booo-
-45o00
- m. .
- m. -
-3oooo toooo . . . , , . .
o 2ao 4o0 Goo Soo looo Goo Woo 16 0 o TIME (s)
Fig. 220.
.. Steam generator secondary-side inventory - loop A.
scooo . . . . . .
. .iosooo
- m. -
-- M*00 m.-
$ "- . .noco $
m m- -
l 5 m.
.soooo
.-45o00 m.-
Booo-
-socoo
- toooo , , . . .
o soo 40o soo soo ioco noo woo woo 11ME (s)
Fig. 221.
Steam generator secondary-side inventory - loop B.
9
, - - _ m .
-178-e00 , , , , , , , -me (sarid)i.oop A i m- . (e. goop e
- 300 ooo.
. -uso 300-
. - 2 00
.pso 300-2 -500 300-20- /.
- --250
~
,. r _.
- q. ,
- . II ..o
-mo , , , , . ,
0 200 400 600 8C0 1000 120 0 6460 1600 THE (s)
Fig. 222.
MFu flows - loops A and B.
MO , , , , , , ,
. (.ond)toop A,....-. ~j 300 330- .t.oor s '. ...... ,,....
,.~.. ..,,
. ., 43o 320- -
E E se- --*eo s00-- --440 neo-. -
.ao 4g0 .
.ago gro . .
-340 480 , , , , , ,
0 200 400 600 800 1000 1200 1400 1000 TWE (s)
Fig. 223.
MFW liquid temperatures - loops A and B.
90
-179-so . . . .
(soruQLDOP A o-I 8 -
-e q g -
. se 2o.
~ ~D Q
5 6 .
--es
...e
.ao. . ---as
. --am 30
--se
-50 . . , . . . .
0 200 400 600 800 900 C00 WOO $800 TlWE (s)
Fig. 224.
Realigned mass flows - loops f. and B.
soo . . . . . .
. .eco (solicQLOOP A (dosh)LDOP B 550 -
.sco E .co.
.... ..~-.....
E
-doO me. .
. .soo doo-
-2o0 uo. -
k
-no 300 , , . . .
0 200 40o eco soO io00 n00 WOO isoo TIME (s)
Fig. 225.
Realigned liquid temperatures - loops A and B. .
D
-180- i l
m , , , , , , .
. .roo *,
wo. -
- neo
- So- -
-soo I 20-oo.
-noo
- -900 to-
- a. - . -ooo 20-- . -300 o , . . . . .
o o zoo 4co soo soo woo coo woo moo TWC (s)
Fig. 226.
Pressurizer pressure.
e , , , , , , ,
s- -
-s 4 .
E E 3 . -m a: .
-5 9
-A A --o o- ' ^^
-t . . . . . . ,
m m o a m m e a e TWE (s)
Fig. 227. I Pressurizer water level.
O l
-181- ,
l o00 , , , , , , , -D5o no- .
l aco- -
)
-ano i e soo- -
- 8 00
400-f i "ND no- -
soo. .
h d (
o- --O
-m0 . . . , , .
0 200 400 600 e00 1000 1200 WOO 1600 TWE (s)
Fig. 228.
Break mass flow.
u , , , , , , ,
a j
me. l .
as- ,
ea. .
e.a -
e- ,
-e.2 , , , , , , .
e aos see ese ese see see wee see .
TWE W Fig. 229.
Break void fraction.
e
- -182-TABLE XXVIII HOT-LEG BREAK LOCA - 4 IN. BREAK EVENT SEQUENCE BASE CASE ,
Event Time (s)
- 1. Break opens 0.0
~
- 2. Turbine and reactor trip 0.5
- 3. Turbine-stop valves close 0.5
- 4. Secondary-side heater and heater-drain trip 1.0
- 5. Condenser feed from turbine trip 1.5
- 6. TBV loop A opens for first time 4.4
- 7. TBV loop B opens for first time 4.8
- 8. RPI system actuation on low primary system pressure 16.8
- 9. TBV loops A and B open/close
- 12. MFCV override trip 46.8
- 13. TBV loops A and B open/elose 92.1
- 14. Candy-canes void 125.0
- 15. ICS closes SUFCVs 300.0
- 16. Accumulator injection loop A (first time) 54C 7
- 17. Accumulator injection loop B (first time) 541.0
- 18. Accumulacor injections loop 678.6
- 19. Accumulator injection loop A 726.1
- 20. Accumulator injection loop B 784.5
- 21. Accumulator injection ceases loop A 828.7
- 22. Accumulator injection loop A 921.4
- 23. Accumulator injection loop B 925.7
- 24. Accumulator injection ceases loop B 947 .5
- 25. Accumulator injection ceases loop A 947.7
- 26. Accumulator injections (both loops) ~1100.0
- 27. LPI system actuation on low primary systes pressure ~1236.0
- 28. End of calculation ~1400.0 Initially, the break mass flow was quite ' large (> 450 kg/s) until ~200 s.
Following this time, the upper regions of the primary became significantly void causing the break mass flow to decrease. The candy-canes for both loops were completely voided by ~125 s as shown in Fig. 230.
Mass flows and liquid temperatures for the loop-A and -B hot legs are !
shown in Figs. 231 and 232, respectively. The primary system flows decreased to approximately sero following . the RCPs trip. The loop-B hot les flow became stagnant as the candy-cana voided; however, the loop-A hot leg did not stagnate and continued to feed the break with vapor. Bot-leg liquid . temperatures decreased in accordance with the primary system depressurization. TRAC
=- -w --* _ _ . _ . _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ - _ _ _ _ . _ _ _ _ _ _ _ _ _ _ . _ _ . _ _ _ _ _ _ . _ _ _ _ _ . _ _ . . _ _ _ _ _ _ _ _ _ _ _ _ _ _ . . . _ _ _ _
. - . . . - - . . 2. . . . . .
-183-t 12 , , , . . . .
(polid)LDOP A
,. (Moo's ,- .
o.e -
a er o.s .
o.4 -
a2-o- -
-o.2 , . . . . . .
o 2co 4o0 son nos eco noo woo moo M (s)
Fig. 230.
Candy-cane void fractions - loops A and B.
moco . . . . . .
.ziooo (solid) LOOP A sooo.
? We . . ,soo
. .wooo sooo- -
.eseo
"~ ~
m . -me y h 3 anoo. .
.m o- -
=a%'A -
-o
._ woo
-2 coo , , , . . . . .
o 2co 4ao coo soo moo woo woo seco M (s)
Fig.1 231.
- Hot-leg mass flows - loops A and B.
\
lN
\
-184-calculated a minimum hot-leg temperature of ~455 K. The loop -A and -B cold-leg mass flows are presented in Figs. 233 and 234, and exhibited trends similar to *
! the hot legs. The corresponding cold-leg liquid temperatures are shown in Figs. 235 and 236. Minimum cold-leg temperatures (before LPI actu.ation) were calculated to be ~430 K at -650 s for loop Al and ~430 K at ~600 s for loop A2.
During the stagnant period in the cold-legs (between ~400 and ~700 s), RPI flow into the downcomer along with accumulator injection at ~540 s and a decreasing ..
vent valve vapor mass flow rate were responsible for the calculated minimum temperatures. The vent valves opened at ~50 s immediately following the RCP trip. The vent valve total mass flow is presented in Fig. 237. The loop B cold-leg liquid temperatures had trends similar to those found in loop A but with the minimums (before LPI actuation) occurring at times cotresponding to loop-B accumulator injection. TRAC calculated minimum cold-leg temperatures of
~425 K at ~770 s for loop B1 and -445 K at ~950 s for loop B2. The initiation
! of the low pressure injection (LPI) system at ~1240 s on low primary system pressure dropped the cold-leg liquid temperatures near or below -400 K. The LPI system injected directly into the vessel downcomer in axial level 7.
As identified in Sec. I, transients involving system repressurization and overcooling have been identified as events that cculd damage and possibly cause the failure cf a PWR vessel. Thus, the key PTS parameters are pressure and liquid temperature in the vessel downcomer region around the veld locations.
For this particular transient, the primary concern is thermal shock because the primary systee will not repressurize. In this plant, the welds located in vessel level 6 in the TRAC model are the veld locations that are considered important for this study. TRAC calculated a sinimum downconer liquid temperature of ~350 K in level 6. The system pressure at this minimum downconer liquid temperature was calculated to be ~1.0 MPa. The minimum temperature.
values calculated for downconer level 6 (between ~540 s and ~1200 s) directly correspond to accumulator injections. The LPI system actuation produces the minimum temperatures af ter ~1200 s. Downcomer liquid temperatures at the top axial downcomer level (just belo9 the cold-leg nozzles) for each azimuthal segment are presented in Fig. 238.
- d. Summary. The Oconee-1 plant response to a 4-in. break in the surge line was calculated using TRAC-PF1. For this transient, the ICS and all key components were assumed to function correctly. Also, the operators were assumed to trip the RCPs 30 s af ter HPI actuation. TRAC calculated a minimum downconer j I
liquid tempere:ure of ~350 K; the system pressure at this minimum temperature was ~1.0 MPa.
-185-3 g a T I I
-se0 (sek)t00P A ,
(dosh,4.DOP S - -
See- ,
sm ,
500- ,p,
- E
-Se .
IS40.. ,,
500-
.-440
-40
- 410 480-
, -300 3so 440 . , , . . .
0 200 dC3 600 A00 900 U00 1400 1000 M (s)
Fig. 232.
, Hot-leg liquid temperatures - loops A and B.
3000 . , , . . . .
f
- -10000 (solicf)LDOP A1 4000-
. -so00 3o00-
- -sou0 1- .
_0 I .
I m00- .
-2000 i
o-- = #.
- y. , % [ --O /
(, -
~
--8
-io00 . . . . , ,
o 200 .0c e00 soo n.7 - o00 w00 =00 M (s)
- 1. / \
Fig. 23,3 Cold-leg, mass flows - foops Al and A2.
, 1 t ' ,
j/
I i . ,.
-*I ..*' j j
+ ,.: _
y .,
1 4 l
-t 2 g.
>e <- ,.
O A f, f% _.
p
> .o
- 1 i ' j
- t
._.,s .i .
l I
-186-
- sooo . . . . . . .
(polid)LDOP 91 ~""' .
.. ) @eWQLDOP B2 .
-sooo ,
aooo-
. emo 1 1
& sooo. -
I 1
y_ -
. ooo g
- o. . -u .:? . --o
, ,.. i . -2000
--dooo
-tooo , , , , , -
o zoo soo eco soo eco noo woo 500 M (s)
Fig. 234.
Cold-leg mass flows - loops B1 and 52.
ooo . . . . . . .
-Soo (solid) LOOP A1
"~ ,
@odQLOOP A2 .,
sea. -
-soo 8 I'8' a, .i E
- k. -
-mo L'
ISoo-.
cs-- I I -"
8.-
(! j l';' ..-
9 1 -
.. g V' 4eo . :: 4 . . -
seo 5 l -
4:s- .
- --3oo i
doo- -
-29o m- -
360 , , ,
o too 4o0 Sco ono moo noo woo moo M (s)
Fig. 235.
l l Cold-leg liquid temperatures - loops Al and A2.
{
I
-187-5 E I B B I I
-80.
m- Id # " -
(dosoor a2 .
soo- -
. sos g m. -
E
- -ao
.o0 .
l !g g I as-go_
2/jij h 1
- -#0 33o at;,- .
l -
3oo
, no- _
- r. -
373 .
-200 350 , , , . .
o 200 400 soo soo coo noo uso c30 1WE (s)
Fig. 236.
Cold-leg liquid temperatures - loops El and 32.
use , , , , , , ,
eso. .
1 .. .
me- . -
I
%.)a.uw,f( .
.. J .
e 4 4 4 4 h 5 h m lbd%$
Fig. 237.
Total positive vent-valve vapor mass flow.
-188-Adequate fluid mixing between the vent-valve fluid and the egid-leg (HPI) fluid in the downconer at the cold-les junction maintained the downconer liquid .
temperatures above -450 K. However, the actuation of the 1.PI systen at ~1:40 e ,
dropped the downconer temperatures very rapidly and below the current NDT value
(-365 K) for the Oconee-1 plant. Even though the downconer liquid temperatures were calculated to be below the current NDT for Oconee-1, this calculation could not be considered a significant PTS transient because repressurization did not occur.
F. Rancho-Seco Type Transient (SC Dryout Followed by EFW Overfeed)
- 1. Introduction and Summary. The thermal-hydraulic response of the Oconee-1 plant to a Rancho-Seco type overcooling transient, that is, steam generator (SG) dryout followed by emergency-feedwater overfeed, has been analyzed. The accident sequence began as a loss-of-sain feedwatet transient (MFW puaps trip). The EFW-control valves failed to open on demand but were subsequently manually opened by the operator. Also, the RCPs remained on during the transient, and the EFW to the SCs was not terminated until 4200 s. The primary system repressurization was limited to ~13.8 MPs as a result of the operator throttling the RPI system.
800 - . , , , ,
-sce
$$Q- .
_w i N E m. .
k %T, .. - +
~
e0- -
it TH Z g -
216 -30e g m. e226 ,
a236 g ,e
. +246 P 20e
- m. 256 .
j 266
-me 300 . , ,
0 200 400 000 soo 1000 C00 #00 400 TWC (i)
Fig. 238.
Downcomer liquid temperatures - vessel axial level 6 (all azimuthal sectors).
l
-189- i 1
. TRAC calculated a minimum downconer liquid temperature of J52 0K at 4200 s. Repressurization of the primary system to ~13.8 MPa was also calculated.
- 2. Model Description and Assumptions. A complete description of the primary systes, secondary system, and ICS modeling can be found in 'Section II.
The steady-etate operating conditions are presented in that section also. The SG dryout followed by EFW overfeed accident specification containing the assumed plant initial conditions and postulated event sequence is presented in Table XXIX. The calculated event sequence is presented in Table XXX. In order TABLE XXIX RANCHO-SECO TYPE TRANSIENT INITIAL CONDITIONS AND POSTULATED EVENT SEQUENCE Initial Conditions:
- 1. Reactor at 100% power
- 2. Nominal temperatures and pressures
- 3. Decay heat: 1.0 ANS standard
- 4. Pressurizer spray / heaters function as designed Postulated Sequence of Events:
- 1. MFW pumps trip
- 2. Turbine trip (TSVs close)
- 4. Reactor trip (on high pressure)
- 5. Both EFW control valves fall close
- 6. SG dryout
- 7. PORY (primary) function as designed
- 8. Turbine bypass system operates as designed Operator lully opens EFW control valves at 9 minutes 9.
- 10. HPI activates on low pressure all. Accumulator and LPI function as designed
- 12. Operator fully opens EFW control valves at 9 minutes
- 13. Operator limits pressurization to 13.8 MPa by throttling HPI flow
- 14. EFW flow terminated at 70 minutes
- 16. Operator throttles HPI to maintain 42 K subcooling after SGs are restored to proper level "May be phenomenologically dependent , ;
-190-TABLE IKK RANCHO-SECO TYPE TRANSIENT ,
SEQUENCE OF EVENTS .
Event Time (s) t
- 1. NW pumps trip, MFCVs and SUFCVs close 0.0
- 2. TSVS close (both loops) 0.0
- 3. Reactor trips on high pressure 4.4
- 4. TBVs actuated -4.6
- 5. FORV actuated 226.4
- 6. E W initiated to both SGs 540.0
- 7. HPI actuated on low pressure ~738.0
- 8. HPI throttled to limit repressurization -1255.0
- 9. E W terminated to both SGs 4200.0
- 10. Minimum vessel downconer liquid temperature (-452 K) calculated ~4200.0
- 11. Calculation terminated 4300.0 to ensure the correct plant response to the postulated sequence of events, significant portione of the TRAC ICS model were used.
- 3. Results. Following the loss of MW and coincident turbine trip, the i reactor tripped from full power on high RCS pressure at -4.4 s. The FORV j functioned properly between -226 e and ~550 e and relieved the pressure increase
- in the primary systen as the SGs dried out. The primary system pressure, 1
pressuriser water level, and PORV mass flow rate are presented in Figs. 239, 240, and 241, respectively. On the secondary side, the TBVs functioned properly and relieved secondary-side pressure increases that occurred also during the SG dry-out period. Figures 242 and 243 present the secondary-side pressures for l loops A and B, respectively. At 540 s, the E W valves were opened fully and E W l
began to refill the SGs. EW mass flow rates and liquid temperatures at the point of injection are presented in Fiss. 244 and 245, respectively. The SG inventories for loops A and B are shown in Figs. 246 and 247, respectively. As the EW was injected into the SGs, the secondary-side inventories began to recover and the primary system began to depressurise. Cold-leg liquid temperatures for loops A and B are presented in Figs. 248 and , 249, while Fig. 250 presents the hot-les liquid temperatures for both loops.
As a result of the depressurisation, the RF1 system was actuated at ~738 s l on low pressure. At ~1255 s, the HPI system flow was throttled to limit RCS l repressurisation to -13.8 MPs as specified in the event sequence. The *RF1 l system flow was continually throttled to limit systen repressurisation
-= a
- - -n, - -
,- ,,w
-191-so , , . . . . . . -moo
,o. -
. mes
- mos I. ** -
_ :::::::::/
~
'**o*
I ..
l gg. -
- moo go. -
. . - mos me. -
moo so , , . , , . . .
o soo moo soo sooo asco 200o anco moo aos MW Fig. 239.
Primary system pressure, n , , , . . . . .
- g. -
,. - so a- -
5 1 v. , - E N ..
- = N
_w -
I s- -
.s 5
- a. -
3 - a
- g. -
l . s 1
m e soo noo soo sono asso ao'co anos mac MN .
Fig. 240.
Pressurizer water level.
e l
l w . ,- w ,
-192-v.s , , , , , , , ,
. as -
s- -
as, gg . .
.a
$ m- - 2 4 as d -
74- -
s E N . g f s- -
,, 3 u- -
.s e- - - - o i
1
-u , , , , . . . ,
a soo moo moo soon noo 2000 anoo moc moo M (s) l Fig. 241.
PORV mass flow.
oc , , , , , . ,
l
~
au 3 .
l
! l M- .
i . .
es. .
.. . .e
- n. -
too
- a. -
- aee
- e. .
1 l as- - .
30 , . , , , - , ,
e aos som use asse aseo sooo aseo moo a:ae M (4 Fig. 242.
SG-A secondary-side pressure.
e
l
-193-I so . . . . . . . .
see Ma ,i
. l
- a. M "
. .oo l I so , .
soo 1SS-eso 1
80- ~
-Sto
- g. "
soo
- a. "
o Soo lobo abo 2eho MM M M M TihtC (s)
Fig. 243.
SG-B secondary-side pressure.
wo . - , . . . . .
i 35 Shf) LDOP A f.
gnosn)toop a : :
90- : : . -ma . wgow
, , . . . , , , , . . .. . . . . * * *
- i so
- 8818 "
So- i .
~
g-e- ' '"
,, 4 m
) ..
n
,, _ -- e
-ai
~" o soo isso moo sooo m anoo zoo m anoo w 60 Fig. 244.
l EFW mass flows.
I l 1 l
-194-oso . . . . . . . .
. DAS4 - LOOP A seo< -*- 80 *
- 100' 8 ,-
g m-ee. .
ano-aso-I -e : : : _ : : .,
aso , ,
o nos moo saa 200o asco sooo anoo ar,oo 44o w (e) .
Fig. 245.
EFW liquid temperatures at injection point.
ococo .
- .psooo ococo.
nooo- ,.
ooooo. .nosso m.
. k g ..oooo g 3 4oooo.
3 5 . , coo E .
] soooo-sonno l
l soooo.
- moco wooo.
- e e-
=Wooo eso sooo sono anos o soo nos eso sooo asco M is)
Fig. 246.
SG-A secondary-side inventory.
n -- --
-195-soooo , , , , , , , ,
j esses- .
woese ,
mese- - -
.m Goeoo-
-enese
- m. .
I g secon gj ..
3 . .
3 E noso E l 2egag. -
l sesos neceo-
,,,o . . moon
- e. - o
-scw , , . ,
o son eco soo 200o stoo mor> asco 4o00 4 00 TME (s)
Fig. 247.
SG-B secondary-side inventory.
eco , , ,
-c ,
kid)LDCP A1 ese son- (poenMxr A2 -
suo sec. .,
.neo E E seo- . .
soo. . neo a .ee eso. .
g
-ee 5 neo. .
aso ano. .
ase 1 1 sooo aseo sooo aseo asco aos i i o soo uno ano ThE(q)
Fig. 248.
Loop A cold-leg liouid temperatures.
l
-196-one . , , , , , ,
g pg oso '
eso- (seWQL.DoPm2 suo goo. .
- neo 3 E 1 o- -
gen. . ano en oog. .
1 -me 4eo. .
soo me.1 .
aeo 440 . . . . . . . .
o ano noo soo sooo asco sooo anco 4o00 moo ThdE (s)
Fig. 249.
Loop B cold-leg liquid temperatures.
ooc , , , , , , , ,
M@A soo soo- $seWQl.DoP B so soo- -
oso a '
E seo. -
3.
go. . aos soo- N~
~
mo me.
\ ' .
-en
. soo ano. .
'%, soo eso , , , , , . ,
o soo moo moo sooo sono soon anoo esoo ese MM Fig. 250.
Hot-leg liquid temperatures.
-197-throughout the transient as indicated by the loops A and B HPI mass flows shown in Figs. 251 and 252. The RCPs were not tripped following the actuation of the '
HPI system, and continued to operate as specified in the event sequence. Cold-leg mass flows for loops A and 2 are presented in Figs. 253 and 254, respectively. Figure 255 presents the hot-les mass flows for both loops.
The continued operation of the RCPs provided forced convective heat ,,
transfer on the primary side that assisted in the rapid cool down of t'he primary system as indicated by the cold-leg and hot-leg liquid temperature profiles between ~540 s and ~3500 s.
At ~3500 s into the transient, the loop A secondary side (SG and steam lines) had been completely filled with EFW and began to repressurize as shown in the SG secondary-side pressure profiles. Also, the loop B secondary side was calculated to repressurize ~200 s later than the loop A side. Both secondary sides were repressurized to the TBV setpoints as EFW continued to feed the system. As a result ci the secondary-side repressurization, the primary side began to cool at a slower rate. At 4200 s, the EFW was terminated to both SGs (as specified). It was st this point in the transient that TRAC calculated a minimum vessel downcomer liquid temperature of ~452 K. The system pressure at this calculated minimum downconer liquid temperature was ~13.8 MPa. Vessel downcomer liquid temperat are profiles for all six azimuthal sectors at axial level 6 (at the weld loce-ions) are shown in Fig. 256.
~
- 4. Conclusions. The overcooling of the primary side of the Ocenee-1 plant caused by a SG dryovt followed by EW overfeed (Rancht-Seco type transient) was simulated with TRAC-PF1. The TRAC simulation calculated most of the plant response and occurrences as outlined in the postulated sequence of events. A minimum vessel downcoment fluid temperature of ~452 K was calculated, at 4200 s. Repressurization of the primary system to ~13.8 MPa was also calculated.
IV. CONCLUSIONS AND RECO.0(ENDATIONS The response of the Oconee-1 plant for several overcooling transients has been predicted using 11 TAC-PF1. The complete plant including the primary and secondary sides wa modeled se that accurate predictions of system thermal-hydraulic conditions could be made. The plant control and protection systems were also modeled in sufficient detail to simulate actual plant response during these postulated overcooling tr.asients. The results of these calculations are to be used for PTS analyses at ORNL.
-198 - .
. v.s , , , , , . , ,
,, NLOOP A1($oW$1EP A2 , ,
.www
- as as- ' -
'n ~ -
a WUsw
- as i
- n. -
Q au 8 i; p .. ,
7.s - . .
.s .
4 j- s 3 s- l l/t -
,, 3 I' '
l ,
l l
2A- l 1
s 3
- o. ,
=
- J
- LsJilt_._ A .__. ..
-2.s . . . . . . . .
--s o soo moo noo sooo asoo sooo anoo ao00 4 00 <
MW j Fig. 251.
Loop A HPI mass flows.
w . , , , , ,
(#4 LDoe st(seWQ p m2 u- ,
as a wisw e- "
- #18" u
. l .ao e- -
' 8 l
.. . I' l lt b4 .
g n li -
- ,h g
N -
L
- t k \
g.,
a-
' > - -s l l l 1
e: : .i 4 -
- - - w -: - , - - - s e ano neo een sooo anoo acon anoo aseo moe l MW 1
i Fig. 252.
Loop B BPI mass flows.
98
-199-seo , , . . . , , .
$orsd)LDOP A1 since .
asse- W A2 -
sano asse- -
. ioeco O
deos-1eeoo 5 me. - -
tasso =
.seco -
moo- -
.anoo m.o. -
asco stoo , , , , ,
-seoo o soo moo moo sooo noo anon anco moo ano M (s)
Fig. 253.
Loop A cold-leg mass flowr.
seo . , , . . . .
bond)LDOF E1 -
-smaa j nano- 6see mz -
l . ioso oooo- -
sesoo Moo-
- k esoe &
f moo. -
aneco 5
d saco- ,
43eo- -
Stoo me. . -
-osos atoo , , , . .
~****
~~
e soo sooo moo anoe anoa sono anoo sooo aos l
M (s) l Fig. 254.
Loop B cold-leg mass flows.
-200-sono . , , . . . . .
maae
@id)LDOP A '
,,,,.. Gnaenors -
assee mese. assee
. q
. seco e, esee. ,. ~~
I I
m-seos f - s moo-ensos seco
" Moo , . . .
anoo anoo sooo ::co moo moo e soo moo moo MW Fig. 255.
Hot-leg mass flows.
aes , , , . , , , ,
j 1-R TH Z see wo- / 216 -
226 asef 236 -
see
.. E 246
- 256 -
n.
266 Ieso- *4 mo-
- .ae .
soo.
8 mo-
- ast me. - ass me , . . . . .
e sao neo son aseo asoo anos snee moo a wW Fig. 256.
Downconer liquid temperatures-vessel axial level 6 (all azimuthal sectors).
e
-9e
-201-Several overcooling transients were analyzed. The transients calculated I included a MSLB with a delay in isolating the affected steam generator, a snall-break PORV LOCA with failure of the ICS to throttle MFW flow and RCP trip, and TBV transients with SG overfeed. An actual planc transient '(Oconee-3 turbine trip) was also simulated by TRAC to compare with actual plant data. Two small hot-leg breaks were also analyzed to investigste the effects of vent-valve flows on downcomer fluid mixing. Finally, a Rancho-Seco type transient was investigated.
The results of the calculations indicate that some overcooling was obtained in all of the cases analyzed as evidenced by highly subcooled liquid temperatures in the downconer. The most severe transient in terms of overcooling and system repressurization was the TBV transients. For the TBV case (two banks of TBVs), the minimum calculated downcomer fluid temperature was
-350 K, and the primary system repressurized. The least severe transient was the PORV-LOCA transient, which had a predicted minimum downcomer fluid temperature of ~528 K. The final NDT temperature for Oconee-1 is ~365 K af ter 32 effective full power years of operation.
It is recommended that other calculations be performed to fully address the Oconee-1 PTS concern. Specifically, other operation actions shoeld be considered to fully cover the spectra of overcooling scenarios. In the case of the small-break LOCAs, other break sizes and locations should be investigated.
Additional failures of the ICS and protection systems should also be analyzed to see if more severe overcooling transients could occur. For example, a MSLB calculation with run-away MFW flow and all other plant systems operating would possibly lead to a more severe overcooling transient.
REFERENCES
- 1. Safety Code Development Group, " TRAC-PF1: An Advanced Best-Estimate Computer Program for Pressurized Water Reactor Analysis," Los Alamos National Laboratory report (to be published).
- 2. R. C. Kryter, et. al., " Evaluation of Pressurized Thermal Shock," Oak Ridge National Laboratory report ORNL IM-8072, NUREG/CR-2083 (October 1981).
- 3. J. D. White, " List of Oconee-1 Tranzients for Thermal-Hydraulic Calculations," Oak Ridge National Laboratory letter, (December 1982).
- 4. " Transient Assessment Report for Oconee Nuclear Station Unit III Reactor Trip of March 14, 1980," Duke Power Company report (No date).
-202-
- 5. C. D. Platcher, *RELAP $ Thermal-Hydraulic Analysis of PreL arized Thermal Shock Sequences for the Oconee-1 Pressurized Water Reactor," ,
Idaho National Engineering Laboratory report EGG-NSMD-6343 (July 1983).
- 6. J. R. Ireland and R. J. Henninger, " Analyses of B&W Small-Break LOCA TRAC Calculations," Los Alamos National Laboratory report LA-UR-82-3294 (November 1982).
e t
l t
l - .
1 1
O
-203-APPENDIX A OCONEE ICS CONTROLI.Eit FOR LOOP A (all signals input are in SI units)
(initialization of valves is for steady state only) ..
(Letters indicate boxes in the previous figure)
BTU LIKITER A = 0.00204083
- RCFLOWA B = -605.4459 + 1.04092
- RCTEMPA , -10.0 < B < 9.080 C = 82.549 - (1.16958e-05)
- SGPRESA , -1.0 < C < 9.080 D = -11.036 + 0.037250
- FWTEMP , -1.270 < D < 9.080 E = -16.0 + B + C + D , -10.0 < E < 12.0 F = 0.55555 + 0.055555
- E R = -10.0 + A
- F
- hold initial value for H until 10 s have passed IF(TIME .LT. 10.0) R = 8.0
" TOP" OF LAYOUT
- -20%/ min ramp after trip -
A1 = TIME - (TIME OF REACTOR TRIP) 31 = 1.0 - (0.2/60.0)
- A1 , 0.0 < B1
- f.w. demand function C1 = 18.0
- B1 l
DI = f(C1) : C1 D1 03 2N.0 0.562 240.0 3.6 320.0 5.4 356.0 t 9.36 402.0
- l 18.0 460.0 l 21.42 483.0
-204- .
f.w. temperature compensation .
21 = -460.0 - D1 + 1.8
- FWTEMP F1 = 1.0 + 0.0013
- El C1 = F1
- C1 ~~
hold initial value of C1 for 10 s IF(TIME .LT. 10.0) C1 = 18.0
- neutron power cross limiter
- initialize signal IF(TIME .LT. 10.0) POWER = 2568.0E6
- bias signal back to zero SP = POWER - 2568.0E6 1st order lag of power with 4.5 s time constant JL = JL + ((SP - JL)/4.5)
- DELT
- remove bias J1 = JL - 2568.0E6 R1 = 1.6 + 14.4
- 31 Il = -1.0 * (R1 - 6.23053E-9
- J1) , -10.0 < Il < 10.0 K1= f(II) : Il K1
-1E5 -10""6
- 0.5 0.0
- 05 0.0 10.5 10.0 sua f.w. temperature and power limiters SC = -10.0 + K1 + C1 take the smallest value - SC or R R = min (SC,H) initialise signal IF(TIME .LT. 10.0) FWFLOWA = 680.4
- bias signal to zero t -_ . - - - _ _ - - - - - - _ . - - _ _ - _ _ - _ _ _ _ _ _ _ - _ _ _ _ . _ - - . _ - _ . - _ _ . _ _ _ _ . - _ _
-205-FWB = FWFLOWA - 680.4
- 1st ceder lag of f.w. flow error with 1.0 s time constant, loop A FSL = FSL + ((FWB - FSL)/1.0)
- DELT
- remove bias SL = FSL + 680.4
- S1 = 10.0 + 1 - 0.026455
- SL l
STCEN OPERATING LEVEL LIMITERS l
- high level limiter, Loop A
- operating level scale, 96 to 388 in (level in meters)
HL1 = f(ALEV) t ALEV RL1 2 T5T -10.0 9.855 10.0 P1 - -2.0 * (HL1 - 7.0)
- take the smallest between signals S1 and F1 Q1 = min (St.P1)
- low level limiter, Loop A
- startup level scale. 0.0 to 250 in LLI = f(ALEV) : ALEV LL1 0.0 -10.0 6.350 10.0
- decide which setpoint to use dependint on pump trip -
- pumps running 241n = 0.61m = -8.08v IF(PTRIP .NE. 1) STP = -8.08
- low level error function P2 = -2.0 * (LL1 - STP)
- take the largest signal - P2, Q1 T1 = max (P2,Q1) a
-206-
" BOTTOM END" OF FLOW CONTROL
) choose which constants to use depending on whether the STCEN l
- is low level limited or not
- if P2 < 0, low limit has not been hit
- IF(P2 .CE. 0.0) CNSTI = 0.12 IF(P2 .LT. 0.0) CNST1 = 0.1125 IF(F2 .CE. 0.0) CNST2 = 2.4 IF(P2 .LT. 0.0) CNST2 = 0.9 integrate, 71, is the last timestep value of signal T1 U1 = U1 + CNST1 * (T1 + T1,)/2.0
- DELT , -18.0 < U < 2.0 X11 = U1 + CNST2
- T1 Il = Ill + 8.0 , -10.0 < XI < 10.0 !
startup control valve function, Loop A SUA = 64.1164 + 7.44164
- X1 , -10.0 < SUA < 10.0
' normalized flow area for SUFV-Loop A
- this signal sent to valve SUFVA = 0.1
- SUA , 0.0 < SUFVA < 1.0 Main flow control valve function, Loop A MFA = 0.5555
- X1 - 4.4444 , -10.0 < MFA < 10.0
' normalized flow area for MFCV-Loop
- this signal sent to valve A
MFCVA = 0.5 + 0.5
- MFA , 0.0 < MFCVA < 1.0 R
e 4
_ _ _ _ ___._______ _ _ l _ ___ _ _ _.._________ _ ___
-207-i Oconee ICS Controller for Loop B_
(balance of signals come from Loop A sectione)
BTU LIMITER BA = 0.002U4083
- RCFLOWB BB = -605.4459 + 1.04092
- RCTEMPB , -10.0 < BB < 9.080 BC = 82.549 - (1.16958e-05)
- BE BH = -10.0 + BA
- BF
- hold value of BH at 8.0 until 10 s passes IF(TIME .LT. 10.0) BH = 8.0 l
" TOP" 0F LAYOUT SECTION
- take the smallest value - SC or BH BR = min (SC,BR)
- initialize signal IF(TIME .LT. 10.0) FWFLOWB = 680.4
- bias signal to zero FWC = FWFLOWB - 680.4 .
- DELT
- remove bias l BSL = FBL + 680.4 BS1 = 10.0 + BR - 0.026455
- BSL
-208-
- STCEN OPERATING LEVEL LIMITERS ,
- high level limiter, Loop 3
- operating level scale, 96 to 388 in (level in meters)
- BHL1 = f(BLEV) : BLEY BRL1 M5 10.0 -
2.438 -10.0 BF1 = -2.0 * (BEL 1 - 7.0)
- take the smallest between signals BSI and BP1 BQ1 = min (BSI BPI)
- low level limiter, Loop B
- startup level scale. 0.0 to 250 in BLLI = f(BLEV) : BLEY BLLI 6M M 0.0 -10.0
- low level error function BP2 = -2.0 * (BLL1 - STP)
- take the largest signal - BF2, BQ1 BT1 = max (BP2,BQ1)
"BOTIOM END" 0F FLOW CONTROL
- choose which constants to use depending on whether STCEN B
- is low level limited or not
- if BP2 < 0, low limit has not been hit IF(BF2 .CE. 0.0) BCNST1 = 0.12 IF(BP2 .LT. 0.0) BCNST1 = 0.1125 IF(BF2 .CE. 0.0) BCNST2 = 2.4 IF(BP2 .LT. 0.0) BCNST2 = 0.9
- integrate, BT1, is the last timestep value of signal BT1 301 = BU1 + BCNST1 * (BT1 + BT1,)/2.0
- DELT , -18.0 ( BU1 < 2.0 .
Bill = BU1 + BCNST2
- BT1 -
BK1 = BX11 + 8.0 , -10.0 < BX1 < 10.0
- startup control valve function, Loop B SUB = 64.1164 + 7.44164
- EX1 , -10.0 < SUB < 10.0
-209- -
normalized flow area for SUPV-Loop B -
- this signal sent to valve i SUFVB = 0.1
- SUB , 0.0 < SUFVB < 1.0 -
Main flow control valve function. Loop B HFB = 0.5555
- BX1 - 4.4444 , -10.0 < MTB < 10.0 -
- normalized flow area for MFCV-Loop B
- this signal sent to valve MFCVB = 0.5 + 0.5
- MFB , 0.0 < MFCVB < 1.0 PEEDPUMP CONTROL DELPA is the pressure drop for the Loop A MFCV, DELPB for Loop B initialize signal IF(TIME .LT. 10.0) DELPA = 3.55E5
- bias signal to zero DPAB = DELPA - 3.55E5
!st order lag of DELPA with a 1.0 s time constant
- 40 pai limit on both sides DPAL = DPAL + ((DPAB - DPAL)/1.0)
- DELT , -2.44e5 < DFAL < 2.4E5
- remove bias DFA = DPAL + 3.55E5 FA = 2.90074e-5
- DPA - 10.0 -
- initialize signal IF(TIME .LT. 10.0) DELPB = 3.55ES
- bias signal to zero DPBB = DELPB - 3.55E5 t
1et order lag of DELPB with a 1.0 s time constant
- 40 pai licit on both sides DPBL = DFBL + (DFBB - DPBL)/1.0)
- DELT , -2.44e5 < DFBL < 2.44E5
w . , . -
, -210-i
- remove bias .
DPB = DP4L + 3.55E5 FB = 2.90074e-5
- DFB - 10.0 ,
- integrate once per timestep, FE,is last timestep value of signal FE FF = FF + 0.2333 * (FE + FE,)/2.0
- signal R is from " top of layout" section, Loop A
- signal BR would be the identical signal from Loop B FI - 0.5 * (R + BR) - FG
- FP in required pump speed (523.6 rad /see = 5000 rpe)
-E0 3T6.4 0.0 392.8 6.0 460.0 10.0 586.43
- internal limit on rate of change of pump speed is set to 27 rad /see per see 9
e L _ _ _ , _ _
RC Mow A RC Temp A SG Pres A FW Temp - RC Flow B RC Temp B SG Pres B
] ] c 1
'C -, ex em _
$ Yrip 4 ":
e Power FW Flow
=% ~
D S en B Level eb o can O!G il @
si iEi ,__ mg a h
~
FW Mow A "h E EE=
[
- h.
u RC
$N'ax b Pi T h B r8
[bk--t 8 "I"3 : E iE to dmp
,r e tp mp uns
..{3g: m De A Delp B
@ s].
gn. n$rnm i,,
8 m'a .
YU
~
h gY to (g' mp to Comp to mp N3 TRAC-PFI ICS HODEL FOR OCONEE-1 t
-212-Trip System Legend (for S,ection II.C) .
INPUT -
I 1 2 ) 2 3 -
3
( }-
4 4 e
Can Also IP' to ,
comnonent i that fe Affected by this Trip
- 1. the trip #
- 2. description of trip .
- 3. input to trip, four kinds available:
S.V. E signal variable input
., T.S.E. I trip signal expression, a mathematical operation of S.V.'s C.B. I a control block i s any leading i means it is a trip controlled trip defined in the input deck by this f. Following a trip cont. t. f, the condition that the input must meet to change the trip set status is indicated:
[Isummation H I product
- 4. trip output (the trip ISET value) can only be -1 I on-reverse 0 I off .
+1 a on-forward s
e oe
'/. l i
. -213- '
, AFFENDIX B ,
i 4
,4 kITRAP0TJ.TiONS
/,+
x
<f ..
1.
MSLB TRANSIENT Because the; MSLB calculation was run t.o 7200 s., no extraglation of the results is necerstry. The pressuriser pressure, downconer IJguid temperatures, andheattransferesefficientsareshowninFfkr.B-1throughB-3respectively.
< 2 The uncertainties in the MSLB calculatfor.] are categorised as follows..
- 1. Steam generator (SG) secondary-side water level -
TRA0 used a collapsed-leval calculation to approxf?. ate the AP, measurement!.
~
- 2. Main feedwater'(MN). pws; trip - becat ce collapsed liquid Icyc1 used, MFW pump tripped later than if AP were used.
- 3. HPI throttling - core exit temperatures should have been used af ter RCPs dtripped - TRAC used hot-leg temperatures. HPI should 'have been throttled at ~275 e instatd of. -525 s e.s calculated.
- 4. RCP, restart >- 41 K subcooling margin must be, reached 'in all loop g r 4 locations before RCFs restarted - TRAC only used one ' location (hot leg);therefore,RCPssh$uldnothavebeenrestartedat~525s.
The first two uncertainties regarding the liquid-level ~ calculation do not appear to affect the primary-side overesoling calculated by-TRAC. The reason for ' this is because tbe flow into the affected SG prich -to' the feedwater realignment trip" (-50 s) is only through the start up flow control valve (main flow control valves closed because of reactor trip) which limits thejflow to
~15% of norma.\. So, even if ~ the main feed pumps are running r t.he ' flow is limited to aproximately thejease rate as if they are not running. Therefore, the primary-side overcooling rate during this period (0-50 s) is essentially independent of whether or not the main feed pumps are reunics. Also, the emergency feedwater pump ; operation during <
?
the first ~45 s of . the' transient has no effect on the primary-side overcooling because all of this liquid is bypass e;d ,
cut the break (refer to Section III.B of report for further detsP,s). '
4* /g -
The effect of throttling the HPI on primary-system overecoling at '~275 s
- ,e j instead of '~525 e is expected to be small. This is becaate' the overcooling caused by the energy remov.1 ;thEough thei affected SG ia auch greater than'the l'. ;
s ~ cooling provided 1,3 the HFI. / Therefore, the effects of HPI ~ throttling in the MSLB transient ara'h511eved to bejinsignificant for this time period.
e 3, e 'n+
- A ,
- __ _ _ _ ._ . _ . - . tm _ -. fia i MI - ~- -
-214-The effects of restarting or not restarting the RCPs are perhaps the most difficult to estimate. In the TRAC calculation, the RCPs were restarted at
~525 a because the subcooling monitor was not modeled correc'tly. In reviewing the results, it appears that the RCPs should not have been resta'rted at that time, and probably would have not been restarted at all if modeled coriectly.
This is because the candy cane in loop B voided in the calculation and remained .-
voided for a considerable time. Also, the region in the vertical part of the hot leg where the fluid temperatures are measured was also voided. Therefore, the subcooling criteria for restarting the RCPs would not have been set. What most likely would have happened in the TRAC calculation if the subcooling monitor was modeled correctly is that the RCPs would not have been restarted at
~525 s, and the downconer temperatures would have continued to decrease at the same rate as prior to ~525 s until steam generator isolation at 600 s. Then the downconer fluid temperatures vould begin to increase and continue to increase for the remainder of the transient. Also, the system would repressurize to the PORV setpoint as shown in Fig. B-1.
Because there is only ~75 a worth of additional cooling if the RCPs had ,
not been restarted, there probably will not be auch difference in the minlaus downconer fit.id temperature calculated. Therefore, if all of these no
~**
so-l~ sw.w kin -
me h
we-- -
_jogo
./
I e,.
es.
/
. .o soo
, .o es. ,/ -
/
L / -see eg. .
f .
V/,/ ..
soo 30 ,
e eso sooo soon eooo sooo esos moo sooo TIWC (s)
Fig. B-1. -
Pressurizer pressure.
-215-see ,
see; 'l ; i 55 -
see. .
z ,,,.
I l.
ene- l .
I R TH 2 216 .
- 226 1 ese- .j 236 .
'; 246
- i) =256 -
f a266 aeo ,
e soo aseo sooo aoos sooo sooo moo sooo MM ,
Fig. B-2.
Downconer liquid temperatures at vessel axial level 6 (all azimuthal sectors).
uncertainties were removed, it is estimated that the ainlaus downcoser fluid temperature would be 405 K 2 25 K. ;
II. TBV TRANSIENTS The TRAC-PF1 calculated results of six TBV failure transients are presented in the body of the report (Sec. III.D and Sec. III.E). . Each of the calculated transients ended at a time equal to or less than 1500 s. In this section, these results are extrapolated to 7200 s. The parameters extrapolated are the system pressures, downconer liquid temperatures, and the heat-transfer coefficients in the downconer.
The extrapolated pressure histories are presented in Fige. B-4 and B-5.
Following an initial depressurization, the system repressurized to the PORV setpoint in four cases; 5A, 58, 6A and 65. The systes did not repressurize in cases 5C and 6C because the HPI was throttled upon attainment of sufficient primary-system subcooling.
l The extrapolated downconer liquid temperatures are presented in Figs. B-6 I through B-11. A discussion of factors expected to -influence the transient hi, stories through the extrapolation period are presented below.
l
-216-
. asoco ,
$2 g -- -
~
E
- ia000-B- ~
g 30000- .
g_
4000- F z
4000-z s sein sien mInn seine asIno sino Suise asas TIME (s)
Fig. B-3.
Heat-transfer coefficients at vessel axial level 6 (all azimuthal sectors).
The extrapolated heat-transfer coefficients in the downconer are presented in Figs. B-12 tarough B-17.
A. General -
For the TBV failure transients, ICS failure to run back MFW to the affected SG(s).was specified. The method chosen to simulate this failure was to fix the MW pump at its specified value and fix the SUFCV and MFCV in the affected loop (s) in the steady-state position. These valves were maintained in that position throughout the transient. The open position of the SUFCV(s) has proven to be significant. Although the MW pump is tripped and the EW pumps are operating, a significant flow from the hotwell and through the MW pump continues. This flow continues to the affected SG(s) even following tripping of the EFW pumps.
Case 5A The EW pumps begin operation at ~210 s. The two motor-driven and one turbine-driven EW pumps take suction from the surge tank, which empties at i -25A s. At this tine the suction of the turbine-driven pumps only is switched to the hotwell and continues to operate to the end of the transient. Before
~2500 s, the flow through the EFW header is ~255 kg/s of which ~117 kg/s is provided by the EFW pumps and ~138 kg/s cores through the MFW pump and SUFCV.
-217-100
-25 5 .
j~ '
ise- } -2am g
-- -= 3 wu) a 6 1 ,desh) PARAMETIt1C 1 g Lchaden) PARAMETRIC 2 -165 'm a MPL- E
~E I t a- L
.(T .
-m2 W ~
- SS- . * * % . ,,,, , ,7g 30 , , , , , , , 435 0 1000 E00 3000 4000 8000 8000 'P000 Fig. B-4.
Pressurizer pressure histories for Case 5 (Case SA-base; Case SB parametric 1; case 5C parametric 2).
350
-2535 386 -
- 2185 f
k IE' -18 5 3 S P urTRic 1 g .
hehadsh) PARAMrTRIC 3 m a 106- -1485 "
~
lf 80- .g33 i$
OS- ' r. 785 r\
- 30 , , , , , . . 435 0 3000 3000 3000 4000 8000 8000 'f000 Fig. B-5.
Pressuriacr pressure histories for Case 6 (Case 6A-base; Case 68 parametric 1; case 6C parsnetrie 2). ,
4 o
v , - , - , - ~-,n.e -
-218-MO- . THETA = 1 *
- THETA = 3
~ MOO
- g. -THETA = 3
- THETA = 4 .g i
- THETA = S
^gam-
- thera = e E '~
u -4283u
$400- "
U g -3683a:
- a. 450- $
56 5
-3083>
ago. -248.3 l
Seo , , , , , , , 3383 0 1000 2000 3000 4000 5000 0000 Tow Fig. E-6.
Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors) for Case SA.
980 ?B43
- THETA = 1 i
- THETA = 3 886 - + THETA = 3 - L'33
- THETA = 4
_
- THETA = 5 g aso-
- THrm = e -4s4g u u ac a hSo6- -449k g
oc
$ E f400- - 404 456- -3693 l
I 430 . . . . . . . 314 2 e 1000 3000 3000 4000 8000 0000 T000 Fig. B-7.
Downconer liquid temperatures at vessel axial level 6 (all
. azimuthal sectors) for Case 58. -
e
l
-219-gyg- aTHrfA = 1 *
- THETA = 2 .g
.THrfn = 3 645- =THr!1 = 4
- THER = 6 F
w 3g0-wrm = s ~4% '
M M E E
.- -446%
' a a.
M -396.3 3 l': 470 - e- l l
446- "M03 l
)
i i e i s s e e s i 0 1000 2000 3000 4000 5000 8000 7000 '
Fig. B-8.
. Downconer liquid temperatures at vessel axial level 6 (all azicuchal sectors) for Case SC.
See
~
i . THETA = 1
- THETA = 3 640- *wrfA = 3 ,
= THETA = 4
- THEIR = 6 E =-
u wem = . _, du E E E -- . ,.39 -
n A.
a a.
l f430- -2923 ago- -2223 I
SM . . . . . . . 352.3 0 1000 2 00 3000 4000 8000 8000 1000 Fig. B-9.
Downconer liquid temperatures at vessel axial level 6 (all azimuthal sectors) for Case 6A.
. s
-220-
= mm = 1 - 57tL3 ,
q e THETA = 2 gm- . THETA = 3 ,
= THETA = 4 -
- THETA = 5 g m. . wm = . _,o.3E w u
u E 830"
- 473.3 E
800-I 2
-43&3$
400- -4033 4e0 , , , , , , , 36&3 0 1000 2000 3000 4000 n000 8000 7000 Fig. B-10.
Downcomer liquid temperatures at vessel axial level 6 (all
( azimuthal sectors) for Case 6B.
~
i
. THETA = 1
- THETA = 3 S40- + THETA = 3 ,
a THETA = 4
- THETA = 5 g n00 . wm = . _43,3E u u u u 480-
-3823 m n E n.
2 3
$ M" -2923$
380-
-2223 i
340 , , , , , , , 15E3 0 1000 2000 3000 4000 6000 8000 1000 Fig. B-11.
Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors) for Case 6C.
e
-221-Ranco ,
locon- -
w
- nacco-g anoco- - -
0000-g_
<4 ago.
ha000-m 1
0 , , , , , , ,
. c anos sono sono sono sono sono sees sous TIME (s)
Fig. B-12.
Heat-transfer coefficients at vessel axial level 6 (all azimuthal sectors) for. Case 5A.
secco ,
- t 1400o <
5
- lanco- ,
g 1000o-8000- -
i l
.Y* 4 coo-l e-g 3000-m -
e , , , , , ,
e anno woo aseo anos seco esso sono sumo TIME (s)
Fig. B-13.
Heat-transfer coefficients at vessel axial level 6 (all azimuthal sectors) for Case 5B.
-222-I l
ps Hono- .
v 3s000-w E
g socco- -
seco-g _.
< dono-l "- .
M
. - - - - - - + -
TIME (s)
Fig. B-14.
Heat-transfer coefficients at vessel axial level 6 (all azimuthal sectors) for Case SC.
88000 i
su g14000-E
- isogo-g 10000-seco- ,
coco-
, 4 4 coo-I E
B ==00- -
e mins sine mise aims oise eies s anse 11ME (s)
Fig. B-15.
Heat-transfer coefficients at vessel axial level 6 (all I
azimuthal sectors) for Case 6A.
- _ _ . _ ________m__ _ _ _ _ _ ___.__ _ -- ___ _ _ _ __ - - _ _ _ _ _ _ _ __ -- _ _ _
, -223- -
8D8 i st
{84000-E w 33000-E U "" .
l--
g_
L N-5 ===- -
z e , , , , , , ,
e nos asso sono soon esos anno sono asas TlME (s)
Fig. B-16.
' Heat-transfer coefficients at vessel axial level 6 (all azimuthal sectors) for Case 6B.
M, 14000-5
- 33000- '
E g 30000-seco-8 .
8000-
< acco.
hsooo- ~
a 0
e mise mise mise en'ee esimo esino se'so esse M E (s) -
i Fig. B-17.
Heat-transfer coefficients at vessel axial level 6 - (all azimuthal sectors) for Case 6C.
l
-224-1 I
Following surge tank depletion only the turbine-driven EFW pump operates, ,
delivering ~59 kg/s. Total flow through the EFW header is estimated at
~1% kg/s. Af ter ~2500 s the cooldown rate is reduced by tw'o factors. First, the reduced flow to the SG reduces the secondary-side heat-transfer coefficient.
Second, reduced flow from the EW pumps results in a higher mixed-sean temperature for the flow through the EFW header. These two effects reduce the j cooldown rate by ~25% af ter 2500 s. The estimated extrapolated downconer liquid temperature at 7200 e is ~365 1 30 K.
Case 5B The EFW pumps begin operation at ~210 s, but EFW flow is terminated at
-400 s following closure of the loop-B EFW valve on high SG liquid level. Flow through the SUFCV continues at ~125 kg/s. The cooldown rate between ~800 and
~1100 e is -0.0130 K/s. This rate is used to estimate the downconer liquid temperature a; 7200 s. Ihe estimated temperature is -440 30 K.
t Case 5C l
! The EFW pumps bef;in operation at ~210 s, but EFW flow is terminated at 400 s following closure of the locp B EFW valve on high SG liquid level. Flow through the SUFCV continues at ~120 kg/s. The estimated cooldown rate between
~1200 and. ~1500 e is -0.0103 K/s. This rate is used to estimate the downconer liquid temperature at 7200 s. The estimated temperature is ~430 2 30 K.
Case 6A The approach used to estimate the downconer liquid temperature is similar to that discussed for Case 5A. The EFW pumps begin operation at ~150 s. The surge tank empties at ~1750 s. Before ~1750 s the flow through the EFW header l to one SG is -225 kg/s of which ~88 kg/s comes from the EFW pumps. Following surge tank depletion at ~1750 s, total flow through the EFW header is ~180 kg/s of which ~44 kg/s is from the turbine-driven EFW pump. After ~1750 s the cooldown rate is reduced by ~37%. The estimated extrapolated downconer liquid temperature is ~350 30 K.
Case 68 The EFW pump does not operate during this transient because the liquid level in both SGs is too high following MFW pump trip. The downconer liquid temperature cooldown has stopped by the end of the calculated transient. The minimum downconer liquid temperature of ~465 K occurs at ~950 s.
l l
t
-225-Case 6C -
The EFW pump does not operate during the transient. The estimated cooldown rate between ~750 and ~1500 e la -0.0203 K/s. This rate is used to estimate the downconer liquid temperature at 7200 s. The estimated temperature is 450 2 30 K.
B. Summary The extrapolated results for the TBV transients are summarized in Table B-1.
III. SBLOCA TRANSIENTS Extrapolations of the key parameters (primary system pressures, vessel downconer liquid temperatures, and downconer heat-transfer coefficients) to 7200 s are presented in Figs. B-18 through B-23 for the PORV and 4-in-dias.
SBLOCAs. For each transient, the extrapolaticn assumptions, modeling assumptions / uncertainties, and effect of the assumptions / uncertainties are described. Also, uncertainties on the extrapolated results are estimated.
A. PORV LOCA Extrapolation Extrapolation Assumptions The extrapolation of the PORV LOCA primary system pressure, vessel downcomer liquid temperatures, and vessel downconer heat-transfer coefficients presented in Figs. B-18 through B-20 assume the following:
- 1. HPI continues to operate
- 2. PORV remains open
- 3. Accumulators and LPI will not actuate TABLE B-1 EXTRAPOLATED RESULTS FOR TBV TRANSIENTS AT 7200 s Downcomer Liquid Pressure Heat-Transfer Coefficient Case Temperature (K) (bars) (W/m2 K) 5A ~365 2 30 ~170 2 5 ~1200 400 55 4 40 i 30 ~170 2 5 ~1200 2 400 5C 430 230 -40220 6000-7500 1 400 l 6A ~350 1 30 ~170 1 5 ~1200 i 400 63 4 658 2 30 ~170 2 5 ~1200 1 400 6C ~350 2 30 -40 20 ~6000-7500 i 400 0
- sinimum occurs at .950 s 1 . . . -
-226-we' . . . . . . .
g, .
. mes -
EXTRAPOLATED wd- ,
= = -
seen og- -
ase I =<-
ud-
- : : : 1 1 -
}
f -
wd. -
. mas end- -
ast-
) -
ase e d-. -
mes s.e # , . . , .
a noe asas asse esos anos esos mise sees MW .
Fig. B-18.
PORV LOCA extrapolated systern preseure.
se . . . . . . .
ses gas- -
see- k EXTRAPOLATED g ,se-. = = . -
g ase- -
-sas sa-- -
sas see- ,
sae- 1 -
see
.3.- T .
- me see- -
. me We .
m e use asas anos mees sene esse sees MW Fig. B-19.
.PORV LOCA extrapolated downcomer liquid temperature.
9 I
e
I
~
l
-227-3A000 .
~
.E.co.. '
g EXTRAPOLATED 1 i g gloo.-
l_
i g
< m. i N
3000- g x : : : 1 1 1 e , , , , , , ,
e mee sene seen mee sees esse 1ees mam TIME W Fig. B-20.
( PORV LOCA extrapolated downconer heat-transfer coefficient.
wt . , , , , ,
ene
- sa d- .
EXTRAPOLATED (NO LPI)
- ud- s..s .
wie 8-. .
ene EXTRAPOLATED (LPI) i
$ ud- -
see R
4 w d- .
~
od. -
. een a d- ' '. -
me 4 - . .. . . . _a ud- .
~
wd. -
"2 . . . . . . . e e see sees asse mes esse esse mee sees l ThE S Fig. B-21. l 4-in-dian. SBLOCA extrapolated systes pressure. j i
)
1 1
1
-228-ese , . . . . . .
spa- -
ese-
\ see sas-l ,
g see. \ . .
~* ~ s g
. - s - s . .a
. I lm..
e- .
aee
- m. q I
=
"- EXTRAPOMTED (NO LPI) ' g
. . _ . .ses ans. -
"~ '
. EXTRAPOLATED (LPI) .
see. -
m ,
me e use noso sees esse esse sees asse TIME fa)
Fig. B-22.
4-in-dias. SBLOCA extrapolated downcoser liquid temperature.
neec0 m EXTRAPOI.ATED (NO LPI) y isooo- *-a 2
h1400o-p EXTRAPOLATED (LPI) 3 manoo- 2 :
U C 10000-b .
8 .ooo.
~
h 40co-p
'f h I s -- 2 . . . . _ _
i hacco-0 , , , . , , ,
e anos asse anno ecco esso spoo inne sees TIME (e)
Fig. B-23.
4-in-dias. SBLOCA extrapolated downcoser heat-transfer coefficient.
~
-229-
- 3. Accumulators and LPI will not actuate
- 4. No operator action.s taken
- 5. ICS, trips, and system components function correctly.
Modeling Assumptions / Uncertainties The following modeling assumptions possibly affected the calculated -
results:
- 1. MFW pump speed increased to maximum rated speed
- 2. MFCVs fixed open (at steady-state flow area) until realignment trip.
No uncertainties in the TRAC modeling such as, failure of the TRAC ICS/ trips to function as the B&W ICS/ trips would function for this particular accident sequence were found.
Effect of Modeling Assumptions / Uncertainties The modeling assumptions listed probably would not significantly affect the calculated final vessel downconer liquid temperature results. The uncertainty of the modeling has essentially no effect on the calculated downconer liquid temperature. Based on the modeling assumptions described, a downconer liquid temperature uncertainty of 215 K has been estimated for the extrapolated results. ,
B. 4-In-Dias. SBLOCA Extrapolation Extrapolation Assumptions The extrapolation of the 4-in-dirm. SBLOCA primary system pressure, vessel downconer liquid temperatures, and vessel downconer heat transfer coefficients presented in Figs. B-21 througl. B-23 assume the following:
- 1. RPI, accumulators, and LPI continue to operate
- 2. Break is not isolated (closed) -
- 3. No operator actions taken
- 4. ICS, trips, and system comporents function correctly.
Modeling Assumptions / Uncertainties The following modeling assumption affected the calculated results: total LPI volumetric flow rate of 6000 spa (2 pumps) at 50 'F.
No uncertainties in the TRAC modeling such as, failure of the ICS/ trips to
, function as the B&W ICS/ trips would function for this particular accident
, sequence were found.
Effect of Modeling Assumptions / Uncertainties The LPI modeling assumption does affect the calculated final systes
, pressure, vessel downconer liquid temperr.ture, and heat-transfer ' coefficient results. The LPI volumetric flow of 6000 spa (2 pumps) at 50 'F reflects the maximum discharge rate end temperature for the '.PI system obtained from the
-230-FSAR. Variations in the volumetric flow (as a function of system pressure) would somewhat alter the slope of the system pressure curve, the downconer liquid temperature profiles, and calculated heat-transfer 'coef ficients. The extrapolation of the three key parameters following the LPl actuation (~1265 s) is very difficult and should be recognized as a rough approximation. Based upon the modeling assumptions made, the following uncertainties were estimated for -
the extrapolated results:
- 1. System pressure - 1 2.0 x 105 p.
- 2. Downcomer liquid temperature - 1 30 K
- 3. Downcomer heat-transfer coefficient - 12000 W/cm2 K.
C. SG Dryout Followed by EFW Overfeed (Rancho Seco-Type Transient)
Extrapolation Assumptions The extrapolation of the SG dryout followed by EFW overfeed transient primary syst m pressure, vessel downcomer liquid temperatures, and vessel downconer heat-transfer coefficient presented in Figs. B-27 thru B-29 assume the following:
- 1. SG level will not be restored to correct level by 7200 s.
- 2. HPI will not be throttled to maintain 50 F subcooling by 7200 s.
. 3. No further operator action taken.
- 4. ICS, trips, and system components function correctly.
The following uncertainties were estimated for the extrapolated results:
- 1. System pressure - 1 2.0 x 105 p.
- 2. Downcomer liquid temperature - 2 30 K
- 3. Downcomer heat-transfer coefficient - 1 500 w/m2 K.
Modeling Assumptions / Uncertainties .
No modeling assumptions or uncertainties in the TRAC model were found to affect the calculated results.
O t
-231-we' .
.mee ow- .
ad-mW- ,- .
. e ,,
I so W- EXTRAPOLATED .
{
1 _
.een &
sw l ag-
. eas u d- -
sg. . -
.see wd- .
s: . , , . . .- e o see sees sees ases anos sees mee eens MW Fig. B-24.
4-in-dias. SBLOCA extrapolated system pressure.
eee . . . . . . .
e,s- .
ses- .
gps. .
'888 g see- EXTRAPOLATED -
as -
- nee E
I me-.
as- .
.3.e ane- .
..e ase- .
me- . . .
.ao ass- -
ars , , , , , , ,
, e use sees asse mese sees sees mee sees .
l 1 MW Fig. B-25.
4-in-dian. SBLOCA extrapolated downconer liquid temperature.
-232- -
,acco ,
- EXI1 TAP 0 LATED k
- goon- _
E con o .
300o-8 800C- . .
< m. '
$noco- h/p[.f
=
f/
t h a , , , , .
sano Sees sees e sono sono seco esse esos Tike (e) l Fig. B-26.
4-in-dias. SBLOCA extrapolated downconer heat-transfer
)
coefficient.
mesesse . . . . . . . -asse seesses , ___.. .anee i
meseems 0
. sees J
t I
I"eness-useemos
.mee l necesse t
. .mee
' neeeeee . mas amaeses .
e see zees asse sees asas esse mee sees MW i Fig. B-27.
' ' Rancho-Seco type transient extrapolated system pressure.
l
I
-233-ese . . . . . . . .
ese- /-
een
.ao -
eas. .
g -
see i.- . -
.es ens- - ase de 1ase- , .
d
,o .ee ene- ,s .
,o .mo e
en- ,a ed . . . . . . .
e see asse asse esse esse esse mee esse MW ris. B-28.
Rancho-Seco type transient extrapolated downconer liquid temperature.
Moco gnoco. ,, **
B wooo. .
- _ 1a000-
' ***== 8 k i i , e a a
- e sees snee asse seen sees sees use esos 198E W Fig. B-29.
Rancho-Seco type transient extraplated downconer heat-tranefer coefficients.
I 13 m-- - .. . - _ - _ . _ __
-234--
APPENDIX C .
UNCERTAINTIES IN OCONEE FTS CALCULATIONS
' ~
- 1. INTRODUCTION -
Any realistic evaluations in uncertainties occurring in the Oconee PTS calculations should be obtained by sensitivity analyses. Such analyses require time, manpower, and money, none of which were available to assess the uncertainties in this study. In the absence of such resources, we used an algorithm that has a weak basis, but can be used to estimate the influence of uncertainties on the calculations.
Contributors to uncertainty in the calculations include (1) physical models (heat transfer, flow regime, choked flow, equation of state, condensation, frictional losses), (2) component models (fuel rod, staam generator, valves, pumps), (3) initial conditions (operating power, system pressure, primary flow rate, steam generator inventory, pressurizer inventory),
(4) plant model (noding, combined components, setpoints, control delays, shutdown margin), (5) operator actions, and (6) numerical methods. For the same accident initiator, changes in these contributors can cause a wide spread in results, particularly if one focuses on the results at a given instant in time.
One can, by definition, fix the transient' by declaring 'that the only uncertainties in which we have interest are those arising from the TRAC code, i.e., physical models, component models, numerical methods, and plant input deck (excluding setpoints etc.). Even with this restricted definition, the temperature and pressure uncertainties can cause a setpoint to be reached earlier or later such that the transient takes a different path and the subsequent uncertainty at a given instant in time can still be large.
- Such nonlinear behavior and the overall nonlinearity of the equations being solved make estinating uncertainties extremely difficult. One must als,o be careful about arbitrarily picking temperatures and pressures from within the uncertainty ranges; they are not necessarily independent because the uncertainties that may cause the temperture to be lower than the best estimate will probably cause the pressure to behave in a like manner.
e a, , ,,
9
~.,4 - - . - ., - - - - - - - - - - _ . - - - - _ _ " - v ~
l
-235-
~
. j l
II. UNCERTAINTY' ALGORITHM ,
If we concentrate on physical models, we know that heat-transfer correlations match the data to within 10-20%. We have also used TRAC to predict
- PORV flowrates to within 15-25%. Other uncertainties may be within similar ranges. On the other hand, we knew we were within 2-5 K on initial temperatures, and we set the pressure to be the normal operating pressure.
Thus, we assumed that the initial uncertainty was close to zero given that the initial conditions were defined.
Our algorithm ignored the nonlinear effects and accounted for the initially small uncertainty. Basically we assumed that the uncertainty was proportional to the deviation from the steady-state conditions. For the proportionality constant, we relied on the uncertainties seen in heat-transf tr and choked-flow correlations, 10-20%. We used 20% for this study. Please note that in testing TRAC against data from integral experiments, we have been able to predict results to better than 20% but usually only af ter adjustment of the input model to obtain better resolution in specific regions of the calculations.
R us, our algorithm for the temperature uncertainty was
. 6T = 0.2 lTg - T,1, and the pressure uncertainty was 6p = 0.2 lp g p,l, where T g and P g are the transient temperature and pressure, respectively, and T, and p, are the rteady-state temperature and pressure, respectively.
As the transient values began to approach the steady-state values, then the maximus uncertainty predicted so far was used. These algorithms were used directly to obtain the uncertainty in the primary-system pressure and the downconer liquid temperature. ne initial primary systes pressure was 15.03 MPa, and the initial downcomer temperature was 563 K. We 'used a 20%
uncertainty in heat-transfer coefficient at all times.
_ - - - . _ _ _ _ - ?- --
-236-III. UNCERTAINTY ZFFECTS ON PTS TRANSIENTS ,
We exasined each transient to determine how these uncertainties might affect certain system trips listed in Table C-I. Almost all of these systems were tripped by pressure. We did not account for overlapping uncertainties arising from uncertainties in both the setpoints and the pressure. Table C-I also includes the effective uncertainty range for this trips. For example, our best-estimate calculations used 10.44 MPs to trip on high pressure injection.
However, at 11.21 MPs, 20% uncertainty might also cause the HPI to trip on, or with 20% uncertainty, the trip might be delayed until the best-estimate pressure reached 9.29 HPa. In other words the effective uncertainty range is not the uncertainty in the setpoint, but how the pressure uncertainty can be translated into an effective setpoint uncertainty. In the following we examine the possible effect of these uncertainties on the transients.
A. Main Steam-Line Break The locp-B TBV was tripped (7.064 MPa) open at 5 s; the loop-B pressure increased so rapidly that its uncertainty should have little effect on the transient. At 21.2 s the HPI was tripped on (10.44 MPa); the cooling effect of the MSLB so overwhelmed the calculation that an advance or delay in HPI should
- t. ave little effect. Advance or delay of the RCP trip, which occurred 30 s later, may have some effect because the high flows associated with RCPs on -
enhance the heat transfer to the steam generator. However, the pressures dropped so rapidly that the advance or delay would be only a few seconds. At 526 s the subcooling margin 42 K was reached, and the HPI was shut off and the RCPs were restarted. A delsy would have given colder downconer temperatures at a time when the pressure was increasing. An advance may have had the opposite effect. The accumulators were predicted (4.17 MPa) to inject at 531 s. The TABLE C-I SYSTEMS AFFECTED BY UNCERTAINTIES Effective System Setpoint Uncertainty Range Turbine Bypass Valve (TBV) 7.064 MPs 6.94-7.26 High Pressure Injection (HPI) on 10.44 MPa 9.29-11.21 Reactor Coolant Pumps (RCPs) 30s after HPI on Accumulators 4.17 MPa 1.46-5.98 PORV 16.9 MPa 16.59-17.37 Low Pressure Injection (LPI) 1.0 MPs 0.0-3.34 High Pressure Injection Off 42 K + 12.5 K subcooling e
t
-237-uncertainty is such that they could have begun dumping as early as ~70 s. This -
could have caused more cooling, possibly mera subcooling with the miditional possibility for the RCPs to be restarted earlier. The PORV setpoint was hit at 4678 s; the results are insensitive to the uncertainty in when the PORV opened.
B. PORV LOCA The trips observed in this transient were TBV opens at 4.4 s, the HPI comes on at 70 s and the RCPs are shut of f at 100 s. The transient would be insensitive to the uncertainties that might change the timing of these trips.
No accumulator injection occurred and none would be expected, even with the.
pressure uncertainty.
C. TBV Failure-One Bank All three transients would be insensitive to the uncertainty in the opening of the loop-A TBV at 4.1 s. The HPI was initiated at 153 s with subsequent RCP shutoff at 183 s. The pressure uncertainty may advance or delay this trip, but it should have no effect on be transient. The pressure plateau from 180 s to 380 e could be shortened or lengthened. In the second parametric case, the RCPs were restarted when the subcooling margin was reached at 383 s.
Uncertainty in the subcooling margin could advance or delay this restart, which causes the pressure to drop and the downcomee temperature to increase. Although the accumulators were not predicted to actuste, the pressure uncertainty could cause accumulator actuation at approximately 900 e in the second parametric case.
D. TBV Failure-Two Banks Again all three transients would be insensitive to the uncertainty in the timing of the TBV at 4.1 s. The HPI is tripped on at 87.5 s, with the RCPs tripped off 30 s later. With the rapidly decreasing pressure at 87.5 s, these trips and their effects would be insensitive to the pressure uncertainty. In the second parametric case, RCP restart caused a rapid pressure decrease such that at 565 s the accumulators were actuated. Again the pressure decreases rapidly and the accumulator actuation would be irsensitive to the pressure uncertainty. Uncertainty in the subcooling monitor trip could advance or delay the RCP restart or the HPI throttling at 485 s. This would be expected to have little effect on the transient. The pressure uncertainty in the base case approaches the accumulator setpoint at approximately 300 s. The pressure increase that occurs immediately thereafter indicates that the accumuistors would probably shut off almost immediately. .
'%n
-238-E. Two-Inch SBLOCA The transient would be insensitive to the timing of 'the TBV opening at 4.1 s. BPI is initiated at 43 s and RCP trip at 73 s; again this timing is not very sensitive because the pressure is decreasing so rapidly. Accumulator
'~
actuation is predicted at 1800 s; uncertainty in pressure could lead to actuation as early as 1200 s.
F. Four-Inch SBLOCA Again the timings of the TBV opening, HPI initiation, and RCP trip would be insensitive to the pressure uncertainties. The pressure uncertainty could lead to accumulator actuation as early as 280 e instead of the predicted 540 s, or it could be delayed until approximately 1200 s. Low pressure injectica could have started as early as approximately 600 e instead of the predicted 1236 s.
C. Rancho Seco The Rancho Seco transient is fairly insensitive to the pressure uncertainties. 'lhe timing of HPI initiation, predicted at 738 s, would be changed by the uncertainties, but would have little effect on the overall transient.
IV. CONCLUSIONS Overall, it is probable that uncertainties in the timing of the actuation of the engineered safety features, arising from thermal-hydraulic uncertainties, would have little effect on pressurized thermal shock results for these transients. Only a more extensive sensitivity study could verify this conclusion.
b 9
_ _ _ _ _ _ _ . _ _ _ . _ ____.____ ._