ML20135D344
| ML20135D344 | |
| Person / Time | |
|---|---|
| Site: | Oconee  |
| Issue date: | 08/31/1985 |
| From: | Joann Ireland LOS ALAMOS NATIONAL LABORATORY |
| To: | NRC OFFICE OF NUCLEAR REGULATORY RESEARCH (RES) |
| References | |
| CON-FIN-A-7217 LA-10055-MS, NUREG-CR-3706, NUDOCS 8509160026 | |
| Download: ML20135D344 (261) | |
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LA-10055-MS TRAC Analyses of Severe Overcooling Transients for the Oconee-1 PWR
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An Affirmative Action / Equal Opportunity Employer I
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i DISCLAIMER This report was prepared as an account of work sponsored by an stency of the Uruted States Government. !
Neither the Uruted States Government nor any agency thereof, not any of ther employees, makes any warranty,eapress or imphed, or assumes any legal habshty or responsibsty for the accuracy, completeness, or usefulness of any informaton, apparatus, product,or process dationed,or reprewnta that its use would not infrmge prwately owned rtshts. Reference herein to any spectfk commercial product, process,or service by trade name, trademark, manufacturer, or otherwise, does not necessardy constitute or imply its endorwment, recommendation.or favoring by the United States Government or any agency thereof, lhe
~ views and opuuons of authors expressed hereta do not necessardy state or reflect those of the Uruled l States Government or any agency thereof, j r
NUREG/CR-3706 LA-10055-MS R4 TRAC Analyses of Severe Overcooling Transients for the Oconee-1 PWR Cc.npiled by John R. Ireland Contributors B. Bassett J. Ireland B. Boyack J. Koenig M. Burkett J. Lime R. Neiton Manuscript submitted: February 1984 Date published: May 1985 Prepared for Divison of Accident Evaluation offee of Nuclear Regulatory Research US Nuclear Regulatory Commission Washington. DC 20555 NRC FIN No. A7217 8 w 08 _G@2. _ _(/:h@
. UG) Los LosAlamosNationalLaboratory Alamos,New Mexico 87545
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CONTENTS rt ABSTRACT................................................................ 1 i
l 1. INTRODUCTION AND
SUMMARY
.......................................... 2 II. TRAC INPUT MODEL DESCRIPTION AND STEADY-STATE RESULTS............. 6 A. Primary Side.................................................. 6
- 1. Yessel.................................................... 6
- 2. Hot Legs.................................................. 9
- 3. Steam Generators.......................................... 9
[ 4. Cold Legs................................................. 10 4
- 5. Emergency Core-Cooling System (ECCS)...................... 10 '
) B. Secondary Side................................................ 11 i 1. Feedwater Train........................................... 11 L
.' 2. SG Control Va1ves......................................... 13 1 1
- 3. Emergency Feedwater....................................... 15
- 4. Steam Lines............................................... 15
- C. Control System................................................ 15 l 1. Trips..................................................... 16 3 2. 1C5....................................................... 20 D. S te ad y-S t a te Calcul at ion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 29 l
i j III. TRAC TRANSIENT CALCULATIONS....................................... 37 l A. 0conee-3 Turbine Trip......................................... 37 1 1. I n t r od uc t i o n . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 37 l 2. Model Description and Assumptions......................... 38 i 3. Transient Calculation..................................... 38 i 4. Summary................................................... 45
- 5. The NSLB...................................................... 49 l 1. In t r od uc t io n and S umm ar y. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 49 j 2. Model Description......................................... 50
- 3. Resulte................................................... 51
! a. Base Case (Case 1).................................... 51
! b. Parametric Case (Case 2).............................. 69 i c. Parametric Case (Case 3 ) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 86 i d. Par ame t ric Case ( Case 4 ) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 95 i 4. Co nc l u s i o n s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 104
- C. PORY L0CA.................................................... 109 j 1. In t r od uc t ion and S ummar y. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 109
- 2. Model Description and Assumptions........................ 110
- 3. Tr ansient Calculation. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 111 l 4. Summary.................................................. 123 J D. TBV Failures................................................. 123 l 1. One Bank of Two TBVs..................................... 123
{ a. Introduc tion and Summary. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 123
- b. Mod e l De sc r i p t i o n . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12 4 l
- c. Re s u l t s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12 5 i
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- i. Base Case...................................... 125
- 11. Parametric Case 1.............................. 138 111. Parametric Case 2.............................. 141
- d. Conclusions........................................... 142
- 2. Two B a nks o f Two T B V s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 144
- a. Introduction and Summary............................. 144
- b. Results.............................................. 145
- 1. Base Case....................................... 145
- 11. Parametric Case 1.............................. 147 111. Parametric Case 2.............................. 148
- c. Conclusions.......................................... 149 E. Hot-Leg Break L0CAs........................................... 164
- 1. Two-Inch Break........................................... 164
- a. Introduction and Summary............................. 164
- b. Model Description and Assumptions..................... 165
- c. Transient Calculations............................... 166
- d. Analysis of the Loop-Flow gsci11ations. a . . . . . . . . . . . . . . 176
- e. Summary.............................................. 179
- 2. Four-Inch Break.......................................... 180
- a. Introduction and Summary............................. 180
- b. Model Description and Assumptions.................... 180 C. Transient Calculation................................. 180 ,
- d. Summary.............................................. 193 g F. Rancho Seco-Type Transient (SG Dryout Followed by 1 E FW 0v e r f e e d ) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19 3
- 1. Introduction and Summary................................. 193
- 2. Model Description and Assumptions......................... 194
- 3. Results.................................................. 194
- 4. Conclusions.............................................. 206 IV. CONCLUSIONS AND RECOMMENDATIONS.................................. 206 AC KNOW LE D GME NT S . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 207 REFERENCES............................................................. 208 APPENDIX A OCONEE ICS CONTROLLER FOR LOOP A. . . . . . . . . . . . . . . . . . . . . . . . . . . 209 APPENDIX B EXTRAPOLATIONS ............................................ 219 APPENDIX C UNCERTAINTIES IN OCONEE PTS CALCULATIONS................... 241 vi
FIGURES
- 1. Pr imar y-s id e mod el f o r Oconee-1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7
- 2. Ve n t- v a l v e m od e 1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8
- 3. Secondary-side model for Oconee-1................................. 12
- 4. Cross section of feedwater-heater mode 1........................... 14
- 5. SG control-valve schematic........................................ 14
- 6. Reactor trip system............................................... 17
- 8. Hotwell, condensate booster (CB), and MFW pumps trip system....................................................... 20
- 9. SG isolation trip system.......................................... 21
- 10. EFW trip system................................................... 22
- 11. ICS organization.................................................. 23
- 12. Feedwater-control section o f the 1CS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 24
- 13. BTU-limiter control b1ocks........................................ 25
- 14. Neutron power c ross-limiter control blocks. . . . . . . . . . . . . . . . . . . . . . . . 26
- 15. Feedwate r-flow control b1ocks . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 28
- 16. Level-limiter control b1ocks...................................... 31
- 17. Feedwate r-valve adjustment cont rol b1ocks . . . . . . . . . . . . . . . . . . . . . . . . . 31
- 18. MFW p um p-s pe ed co nt rol b1ocks . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 33
- 19. Main steam safety valve modeling for Oconee-3 t ransient. . . . . . . . . . . 39
- 20. Measured MFW flow rates........................................... 40
- 21. TRAC-calculated decay power and Oconee-3 neasured t h e rm al po we r . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 40
- 22. Calculated and measured primary pressures......................... 43
- 23. Calculated and measured pressurizer water levels.................. 43
- 24. Calculated and measured hot- and cold-leg temperatures for loop A........................................................ 44
- 25. Calculated and measured hot- and cold-leg temperatures for loop B........................................................ 44
- 26. Calculated and measured SG A secondary pressures.................. 46
- 27. Calculated and measured SG B secondary pressures.................. 46
- 28. Calculated and measured water levels in SG A...................... 47
- 29. Calculated and measured water levels in SG B...................... 47
- 30. Pressurizer pressure (0-900 s)--base case......................... 54
- 31. Pressurizer pressure (0-7200 s)--base case........................ 54
- 32. Pressurizer water level (0-900 s)--base case...................... 55
- 33. Pressurizer water level (0-7200 s)--base case..................... 55
- 34. Downcomer liquid temperatures (0-900 s) at vessel axial level 6 (all azimuthal sectors)--base case.............................. 56
- 35. Downcomer liquid temperatures (0-7200 s) at vessel axial level 6 (all azimuthat sectors)--base case.............................. 56
- 36. To tal vent-valve flow into downcomer--base case . . . . . . . . . . . . . . . . . . . 58
- 37. Hot-leg liquid subcooling (0-900 s)--base case.................... 58
- 38. Ho t-leg liquid subcooling (0-7 200 s)--base c ase . . . . . . . . . . . . . . . . . . . 59
- 39. Hot-leg mass flows (0-900 s)--base case........................... 59
- 40. Hot-leg mass flows (0-7200 s)--base case.......................... 60
- 41. Candy-cane void fractions--base case.............................. 60
- 42. Upper plenum liquid volume fraction--base case.................... 61
- 43. Loop-A cold-leg mas s flows--base case . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 61
- 44. Loo p-A cold-leg liquid tempe ratures--base c ase. . . . . . . . . . . . . . . . . . . . 62
- 45. Loo p-B cold-leg mass flows--base case . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 62 vii
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- 46. Loop-B cold-leg liquid temperatures--base case.................... 64
- 47. SG A secondary-side water inventory--base case. . . . . . . . . . . . . . . . . . . . 64
- 48. SG A secondary-s ide p res s ure--ba se ca s e . . . . . . . . . . . . . . . . . . . . . . . . . . . 65
- 49. SG A steam-line flow--base case................................... 65
- 50. Counter current flow limiting phenomena in affected steam generator (SC A)--base case....................................... 66
- 51. SG B secondary-side water inventory (0-900 s)--base case.......... 66
- 52. SG B secondary-side water inventory (0-7200 s)--base case......... 68
- 53. SC B secondary-side pressure (0-900 s)--base case................. 68
- 54. SG B secondary-side pressure (0-7200 s)--base case................ 70
- 55. MFW pump speed--base case......................................... 70
- 56. MFW liquid temperature--base case................................. 71
- 57. MFW mass flows--base case......................................... 71
- 58. EFW mass flows (0-900 s)--base case............................... 72
- 59. EFW mass flows (0-7200 s)--base case.............................. 72
- 60. EFW liquid temperature at pump discharge--base case............... 73
- 61. EFW liquid temperatures at injection locations--base case......... 73
- 62. Loop-A HPI flows--base case....................................... 74
- 63. Loop-B HPI flows--base case....................................... 74
- 64. Ac cumul a t o r wa t e r le ve ls --ba s e ca s e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 75
- 65. Accumulator liquid volume discharged--base case . . . . . . . . . . . . . . . . . . . 75
- 66. PORV mass flow--base case......................................... 77
- 67. P re s s u r i z e r p r e s s u re --Ca s e 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 77
- 68. P re s s u r i z e r wa t e r l e ve l-- Ca s e 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 78
- 69. Downcomer liquid temperatures at vessel axial level 6 (all a z imu t h a l s e c t o r s ) --Ca s e 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 78
- 70. Ho t -l e g ma s s f l ow s -- Ca s e 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 80
- 71. Ca nd y-c a ne void f ra c t ion s--Ca s e 2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 80
- 72. Loo p- A co ld-l e g ma s s f l ows -- Ca s e 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 81
- 73. Loo p-B cold-le g ma s s f lows--Ca s e 2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 81
- 74. Loop-A cold-leg liquid tempera ture s--Case 2. . . . . . . . . . . . . . . . . . . . . . . 82
- 75. Loop-B cold-leg liquid tempe ra tures--Case 2. . . . . . . . . . . . . . . . . . . . . . . 82
- 76. SG A s econda ry-s ide wa te r i nvent o ry--Ca s e 2. . . . . . . . . . . . . . . . . . . . . . . 83
- 77. SG A s econda ry-s ide p re s s ure--Cas e 2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 83
- 78. S G A s t e am-l i n e f low-- Ca s e 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 84 1
- 79. SG B s econda ry-s ide wa t e r invent o ry--Ca s e 2. . . . . . . . . . . . . . . . . . . . . . . 84 I
- 80. SG B s econda ry-side p re s su re--Case 2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 85
- 81. M FW p um p s p e e d -- Ca s e 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 85
- 82. H P I f l o w s -- Ca s e 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 87
- 83. P r e s s u r i z e r p re s s u r e -- Ca s e 3 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 87
- 84. P r e s s u r i z e r wa t e r l e ve l-- Ca s e 3 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 88
- 85. Downcomer liquid temperatures at vessel axial icvel 6 (azimuthal s e c t o r s ) -- Ca s e 3 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 88
- 86. Ho t-l e g ma s s f l ow s -- Ca s e 3 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 90
- 87. Ca nd y-ca ne vo id f ra c t i on s--Ca s e 3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 90
- 88. Loop-A cold-le g na s s f lows --Ca s e 3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 91
- 89. Loop-B cold-leg mass flows--Case 3................................ 91
- 90. Loop- A cold-leg liquid tempe ra tures--Case 3. . . . . . . . . . . . . . . . . . . . . . . 92
- 91. Loop-B cold-Icg liquid tempe rature s--Case 3. . . . . . . . . . . . . . . . . . . . . . . 92
- 92. SG A se conda ry-s ide wa t e r i nve n t o ry--Ca s e 3. . . . . . . . . . . . . . . . . . . . . . . 93
- 93. SG A seconda ry-s ide p res su re--Cas e 3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 93
- 94. SC A s t eam-l i ne f l ow--Ca s e 3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 94
- 95. SG B seconda ry-side wa te r invento ry--Case 3. . . . . . . . . . . . . . . . . . . . . . . 94
- 96. SG B seconda ry-s ide p re s su re--Ca s e 3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 98 viii
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- 97. MW pump speed--Case 3............................................ 98
! 98. System pressure for Case 4 and base case.......................... 99
- 99. Downconer liquid temperatures at vessel axial level 6 for j Case 4 and base case.............................................. 99 j 100. Break flow for Case 4............................................ 100 101. SG A secondary pressure for Case 4. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 100 1 102. SG B secondary pressure for Case 4. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 101 i 103. EN flow (SG A and SG B) for Case 4. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 101 j 104. SG A total feedwater flow for Case 4............................. 102
- 105. SG B total feedwater flow for Case 4. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 102 106. SG A tube-bundle-region mass inventory for Case 4................ 103 l 107. SG B tube-bundle-region mass inventory for Case 4................ 103
] 108. Hot-leg A and B mass flows for Case 4. . . . . . . . . . . . . . . . . . . . . . . . . . . . 105 J
109. Hot-leg A and B liquid temperatures for Case 4. . . . . . . . . . . . . . . . . . . 105 t j 110. Loop-A cold-leg mass flows for Case 4............................ 106 i 111. Loop-A cold-leg liquid temperatures for Case 4. . . . . . . . . . . . . . . . . . . 106 !
- 112. Loop-B cold-leg mass flows for Case 4. . . . . . . . . . . . . . . . . . . . . . . . . . . . 107 l 113. Loop-B cold-les liquid temperatures for Case 4. . . . . . . . . . . . . . . . . . . 107 ,
i i 114. To t al HP I f low f o r Case 4 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 108 115. Total mass flow through vent valves for Case 4. . . . . . . . . . . . . . . . . . . 108 :
116. Upper plenum liquid temperature for Case 4. . . . . . . . . . . . . . . . . . . . . . . 109
- 117. SG A secondary pressure.......................................... 112 1 118. SG B sec ond ar y pr e s sure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 112 j 119. NN f l ow--loo ps A and B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 114
{ 120. NW liquid temperature-loops A and B. . . . . . . . . . . . . . . . . . . . . . . . . . . . 114
, 121. Realignment mass flow--loop A (negative flow is into i
- steam generator)................................................. 115 r i 122. Realignment mass flow--loop B. (negative flow is into '
s t e am g en e r a t o r ) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 115 !
l 3 123. Stean generator secondary inventory--loop A...................... 116 124. Stems generator secondary inventory--loop B...................... 116
, 125. Pressuri se r pres sure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 117 i 126. P r e ssur i se r wa t e r leve 1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 117 l 127. PORV mas s f l ow . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 118 l 128. PORV vapo r f rac t ion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 118
{ 129. Ho t-leg mas s flows--loo ps A and B. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 119 l 130. Hot-les liquid temperatures--loops A and B. . . . . . . . . . . . . . . . . . . . . . . 119 [
l 131. Cold-les mass flows--loops Al and A2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 120 t l 132. Cold-les mass flows--loops B1 and B2............................. 120 ['
! 133. Cold-leg liquid temperatures--loops Al and A2. . . . . . . . . . . . . . . . . . . . 121
- 134. Cold-les liquid temperatures--loops 81 and 82. . . . . . . . . . . . . . . . . . . . 121
! 135. Candy-cane void fractions--loops A and B......................... 122 j 136. Vessel upper plenum void fractions--all animuthal ce11s... ..... . . 122 i 137. Downconer liquid temperatures at vessel asial level 6-all j l as inut ha t sec t o r s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12 3 :
) 138. P re ss u r i se r pr e s su re . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12 6 ,
! 139. SG second ar y pressure--loop A. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12 6 j 140. SG secondary pressure--loop B.................................... 127 1 1
141. NN f l ow-- l oo p A . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12 7 142. MW flow--loop B................................................. 129 ,
143. N N liquid temperatures--loop A.................................. 129 144. NN l iqu id t em pe r a t ur e s--loo p B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 130 145. Flow through E N header--loop A.................................. 130 146. Flow through EN header--loop B. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 131
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147. Liquid temperatures in the emergency-feedwater header
--loop A......................................................... 131 148. Liquid temperatures in the emergency-feedwater header
--loop B......................................................... 132 149. SG s econdary invento ry-loop A. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 132 150. SG secondary inventory--loop B................................... 133 151. Ho t- le g f l o w--loo p A . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13 3 152. H o t- l e g f l o w-- l o o p B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13 4 153. Co ld- l eg f l o w--loo p A1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13 4 154. Cold-leg flow--loop A2........................................... 135 155. Co l d-l eg f l ow--l oo p B 1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13 5 156. Co ld-leg f l ow-- loo p B2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13 6 157. Cold-leg liquid temperatures--loop B1............................ 136 158. Cold-leg liquid temperatures--loop B2............................ 137 159. Cold-leg liquid temperatures--loop A1............................ 137 160. Cold- leg liquid temperatures--loop A2. . . . . . . . . . . . . . . . . . . . . . . . . . . . 139 161. Canriy-cane vapor fraction--loop A................................ 139 162. Candy-cane vapor fraction--loop B................................ 140 163. Pressurizer water leve1.......................................... 140 164. Downcomer liquid temperatures (base case) at vessel axial level 6 ( all azimuthal sec to rs ) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 143 l 165. Downcomer liquid temperatures (parametric case 1) at vessel axial level 6 (all azimuthat sec tors) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 143 166. Downcomer liquid temperatures (parametric case 2) at vessel axial level 6 (all azimuthal sectors).................................. 144 167. Pressurizer pressure............................................. 150 168. SG secondary pressure--loop A.................................... 150 169. SG s ec o nd ar y p r e s s u re--loo p B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 151 170. M W f l ow-- l oo p A . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 151 171. MW f l o w-- l o o p B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 2 172. MW liquid tempe ra t ures--loo p A. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 152 173. MW liquid t empe rature s--loop B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 153 174. E W f l o w-- l o o p A . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 3 175. E FW f l o w-- l o o p B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 4 176. E W liquid temperatures--loop A.................................. 154 177. EW liquid t em pe r a t u r e s--loo p B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 155 178. SG s econd ary invento ry-loo p A. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 5 179. SG s econd ary invento ry-loo p B. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 156 180. Ho t- leg f l ow--loo p A . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 6 181. Ho t-l eg f l o w--lo o p B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 7 182. Co ld- l eg f l ow--l oo p A1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 7 183. Co ld-l e g f l o w--loo p A2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 8 184. Co ld- l e g f l ow-- l oo p B 1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 8 185. Co ld-l e g f l ow--loo p B 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 9 186. Cold-leg liquid temperature--loop A1............................. 159 187. Cold-leg liquid temperature--loop A2............................. 160 188. Cold-leg liquid temperature--loop B1............................. 160 189. Cold-leg liquid temperature--loop B2............................. 161 190. Cand y-cane va po r f rac t ion--loo p A. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 161 191. Candy-c ane vapo r f r ac t ion--loo p B. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 162 192. P r e s s u r i z e r wa te r l ev e 1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 162 193. Downcomer liquid temperatures (base case) at vessel axial level 6 (all azimuthal sectors)........................................ 163 X
194. Downcomer liquid temperatures (parametric case 1) at vessel axial level 6 (all azimuthal sectors)............................ 163 193. Downcomer liquid temperatures (parametric case 2) at vessel axial level 6 (all azimuthat sectors)............................ 164 196. SG se conda ry-side p re s sure--loop A. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 167 197. SG s econda ry-side p res su re--loop B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 167 198. SG secondary-side inventory--loop A.............................., 168 199. SG secondary-side inventory--loop B............................... 168 200. MFW ma s s f lows--loo p s A a nd B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16 9 201. MFW liquid temperatures--loops A and B............................ 169 202. Emergency / realigned mass flows--loops A and B..................... 170 203. Emergency / realigned liquid temperatures--loops A and B. . . . . . . . . . . . 1/C 204. Pressurizer pressure.............................................. 171 205. Pressurizer water leve1.......................................... 171 206. Break mass f1ow................................................... 172 207. Break void fraction.............................................. 172 208. Candy-cane void f rac t ions--loops A and B. . . . . . . . . . . . . . . . . . . . . . . . . . 173 209. Cold-leg mass f lows--loops Al and A2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 17 3 210. Cold-leg ma s s flows--loops B 1 and B2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 174 211. Cold-leg liquid tempe ra tures--loops Al and A2. . . . . . . . . . . . . . . . . . . . 174 212. Cold-leg liquid temperatures--loops B1 and B2.................... 175 213. Ho t-le g ma s s flows--loo p s A a nd B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 175 214. Hot-leg liquid temperatures--loops A and B....................... 177 215. Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors)........................................... 177 216. Downcomer liquid temperature comparison for 2-in. break case (vent valves vs no vent va1ves)............................. 178 217. To t a l ve n t -va l ve ma s s f 1 ow . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 178 218. SG s econda ry-s ide p re s su re--loo p A. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 181 219. SG secondary-side pressure--loop B................................ 181 220. SG secondary-side inventory--loop A............................... 183 221. SG se conda ry-s ide inve n to ry--loop B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 183 222. M FW f l ow s --lo o p s A a nd B . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 184 223. MFW liquid tempe ra tures--loops A and B. . . . . . . . . . . . . . . . . . . . . . . . . . . . 184 224. Realigned mas s flows--loops A and B. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 185 225. Realigned liquid tempe ra ture s--loops A and B. . . . . . . . . . . . . . . . . . . . . 185 226. Pressurizer pressure.............................................. 187 227. Pre s s u r i z e r wa t e r le ve 1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18 7 228. Break mass f10w.................................................. 188 229. Break void fraction............................................... 188 230. Candy-cane void f rac t ions--loops A and B. . . . . . . . . . . . . . . . . . . . . . . . . 189 231. Ho t-leg ma s s f lows--loops A a nd B. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 189 232. Hot-leg liquid temperatures--loops A and B....................... 190 233. Cold-leg ma s s f lows--loops A l and A2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 190 234. Cold-leg mass flows--loops B1 and B2............................. 191 235. Cold-leg liquid t empe ra tures--loops Al and A2. . . . . . . . . . . . . . . . . . . . 191 236. Cold-leg liquid t empe ra tu res--loops B l and B2. . . . . . . . . . . . . . . . . . . . 192 237. Total positive vent-valve vapor mass f1ow......................... 192 238. Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors).......................................... 194 239. Primary system pressure........................................... 196 240. P re s s u r i z e r wa t e r le ve 1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19 6 241. PORV mass f1ow.................................................... 197 242. SG A s e conda ry-s ide pre s sure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19 7 xi
I, 243. SG B secondary-side pressure..................................... 198 244. EFW mass f1ows................................................... 198 245. EFW liquid temperatures at injection point........................ 199 246. SG A secondary-side inventory.................................... 199 247. SG B secondary-side inventory.................................... 200 248. Loop-A cold-leg liquid temperatures.............................. 200 249. Loo p-B cold-leg liquid t em pe rature s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 201 250. Hot-leg liquid temperatures...................................... 201 251. Lo o p- A H P I m a s s f 1 ows . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 02 252. Lo o p- B H P I m a s s f 10 w s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 0 2 253. Loo p- A c old-leg m as s f 1ows . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 203 254. Lo o p-B cold-leg m a s s f 1ows . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 203 255. Hot-leg mass f1ows............................................... 205 256. Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors).......................................... 205 A-1. TRAC-PF1 UCS model tor occase-1.................................. 217 A-2. Trip System Legend (for Section 11.C)............................ 218 B-1. Pressurizer pressure............................................. 220 B-2. Downcomer liquid temperatures at vessel axial level 6
( all a z imut hal s ec to r s ) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 21 B-3. Heat-transfer coefficients at vessel axial level 6 (all azimuthal sectors).......................................... 222 B-4. Pressurizer pressure histories for Case 5 (Case SA-base; case 5B parametric 1; case 5C parametric 2 ) . . . . . . . . . . . . . . . . . . . . . . 225 B-5. Pressurizer precoure histories for Case 6 (Case 6A-base; Case 6B parametric 1; Case 6C parametric 2 ) . . . . . . . . . . . . . . . . . . . . . . 226 B-6. Downcomer liquid temperatures at vessel axial level 6 (all azimuths f sectors) for Case 5 A. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 226 B-7. Downcomer liquid temperatures at vessel axial level 6
( all azimuthat sectors) fo r Case 58. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22 7 ,
B-8. Downcomer liquid temperatures at vessel axial level 6 I
( all azimuthat sectors) for Case 5C. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 227 B-9. Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectore) for Case 6A.............................. 228 B-10. Downcomer liquid temperatures at vessel axial level 6 (all azimuthat sec tors) for Case 6B. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 228 B-11. Downcomer liquid temperatures at vessel axial level 6 (all azimuthat sectors) for Case 6C.............................. 229 B-12. Heat-transfer coefficients at vessel axial level 6 (all azimuthal sectors) for Case 5A.............................. 229 B-13. Heat-transfer coef ficients at vessel axial level 6 (all azimuthal sectors) for Case 58.............................. 230 B-14. Heat-transfer coef ficients at vessel axial level 6 (all azimuthat sectors) for Case 5C.............................. 230 B-15. Heat-transfer coefficients at vessel axial level 6 (all azimuthat secto rs) f or Case 6 A. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 231 B-16. Heat-transfer coefficients at vessel axial level 6
( all azimu that sec to rs) f o r Ca se 68. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 231 B-17. Heat-transfer coef ficients at vessel axial level 6 (all azimuthal sec to rs) fo r Case 6C. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 232 B-18. PORY LOCA ext rapolated sys t em pres sure . . . . . . . . . . . . . . . . . . . . . . . . . . . 234 B-19. PORV LOCA extrapolated downcomer liquid temperature. . . . . . . . . . . . . . 235 B-20. PORV LOCA extrapolated downcomer heat-transfer coef ficient.. . . . . . 235 B-21. Four-inch-diameter SBLOCA extrapolated system pressure. . . . . . . . . . . 236 x11
B-22. Four-inch-diameter SBLOCA extrapolated downconer l iqu id t em pe r a t u r e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 236 B-23. Four-inch-diameter SBLOCA extrapolated downconer heat-transfer coefficient...................................................... 237 B-24. Four-inch-diameter SBLOCA extrapolated system pressure........... 237 B-25. Four-inch-diameter SBLOCA extrapolated downconer liquid temperature...................................................... 238 B-26. Four-inch-diameter SBLOCA extrapolated downconer heat-transfer coefficient...................................................... 238 B-27. Rancho Soco-type transient extrapolated system pressure.......... 239 B-28. Rancho Seco-type transient extrapolated downconer l i qu id t em pe r a t u r e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 3 9 B-29. Rancho Seco-type transient extraplated downcoeer heat-transfer coefficients............................................ 240 t
xiii
I TABLES
- 1. LOS ALAMOS OCONEE-1 PTS OVE14C00 LING TRANSIENT CALCULATIONS...... 3 II. TRAC OCONEE-1 TRANSIENT RESULTS................................. 5 III. SIMPLE TRIPS.................................................... 16 IV . BTU-LIMITER CONTROL-BLOCK EQUATIONS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25 V. NEUTRON-POWER CROSS-LIMITER CONTROL-BLOCK EQUATIONS............. 27 VI . FEEDWATER-FLOW CONTROL-BLOCK EQUATIONS . . . . . . . . . . . . . . . . . . . . . . . . . . 28 VII. LEVEL-LIMITER CONTROL-BLOCK EQUATIONS........................... 30 VIII. FEEDWATER-VALVE ADJUSTMENT CONTROL-BLOCK EQUATIONS. . . . . . . . . . . . . . 32 IX. MFW PUMP-SPEED CONTROL-BLOCK EQUATIONS. . . . . . . . . . . . . . . . . . . . . . . . . . 34 X. PRIMARY-SIDE STEADY-STATE CONDITIONS............................ 35 XI. SECONDARY-SIDE STEADY-STATE CONDITIONS.......................... 36 XII . INITI AL STEA DY-STATE CONDITIONS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 41 XIII. MSSV AND TBV SETP0!NTS.......................................... 42 XIV. SEQUENCE OF EVENTS.............................................. 42 XV. COMPARISON OF TRAC AND OCONEE-3 RESULTS......................... 48 XVI. SEQUENCE OF EVENTS.............................................. 52 XVII. MSLB (CASE 2) SEQUENCE OF EVENTS................................ 76 XVIII. MSLB (CASE 4) SEQUENCE OF EVENTS................................ 96 XIX. MSLB ( CASE 3 ) SEQUENCE OF EVENTS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 110 XX. PORV LOCA EVENT SEQUENCE....................................... 111 XXI. TBV EVENT SEQUENCE, BASE CASE.................................. 125 XXII . TBV EVENT SEQUENCE , PARAMETRIC CASE 1. . . . . . . . . . . . . . . . . . . . . . . . . . 138 XXIII. TBV EVENT SEQUENCE, PARAMETRIC CASE 2. . . . . . . . . . . . . . . . . . . . . . . . . . 141 XXIV. TBV EVENT SEQUENCE, BASE CASE.................................. 146 XXV. TBV EVENT $EQUENCE, PARAMETRIC CASE 1.......................... 147 XXVI . TBV EVENT SEQUENCE , PARAMETRIC CASE 2. . . . . . . . . . . . . . . . . . . . . . . . . . 149 I XXVII . HOT-LEG BREAK LOCA--2-IN. BREAK SEQUENCE OF EVENTS. . . . . . . . . . . . . 165 XXVIII. HOT-LEG BREAK LOCA--4-IN. BREAK EVENT SEQUENCE BASE CASE....... 182 7XIX. RANCHO SECO-TYPE TRANSIENT INITIAL CONDITIONS AND POSTULATED EVENT SEQUENCE................................................. 195 XXX. RANCHO SECO-TYPE TRANSIENT SEQUENCE OF EVE NTS. . . . . . . . . . . . . . . . . . 204 B-1. EXTRAPOLATED RESULTS FOR TBV TRANSIENTS AT 7200 s . . . . . . . . . . . . . . 224 C-I. SYSTEMS AFFECTED BY UNCERTAINTIES.............................. 243 i xiv
ACRONYMS ANS American Nuclear Society BTU British Thermal Unit B&W Babcock & Wilcox CB Condensate Booster ECCS Emergency Core-Cooling System EFW Emergency Feedwater EFWV Emergency Feedwater Valve FSAR Final Safety Analysis Report HPI High-Pressure Injection ICS Integrated Control System LOCA Loss-of-Coolant Accident LPI Low-Pressure Inject ion LPIS Low-Pressure Inject ion System MFCV Main-Flow-Control Valve MFW Main Feedwater MSLB Main Steam-Line Break MSSV Main Steam Safety Valve NDT Nil-Duct ility Temperature NRC Nuclear Regulatory Commission ORNL Oak Ridge Nat ional Laboratory PORV Power-Operated Relief Valve PTS Pressurized Thermal Shock PWR Pressurized Water Reactor RCP Reactor Coolant Pump SBLOCA Small Break LOCA SC Steam Generator SC A Steam Generator A SC B Steam Generator B SUFCV Startup Flow-Control Valve TBV Turbine-Bypass Valve TRAC Transient Reactor Analysis Code TSV Turbine-Stop Valve USI Unresolved Safety Issue xv
TRAC ANALYSES OF SEVERE OVERC00 LING TRANSIENTS FOR THE OCONEE-1 PWR Compiled by J. Ireland Contributors B. Bassett J. Ireland B. Boyack J. Koenig M. Burkett J. Lime R. Nelton ABSTRACT This report describes the results of several Transient Reactor Analysis Code (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 input on thermal-hydraulic data to Oak Ridge Nat ional Laboratory for pressurized thermal-shock analyses. The t rans ient calculat ions performed were plant spec ific in that details of the primary system, 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 downcomer region of the vessel. The power-operated relief valve loss-of-coolant acc ident transient was the least severe in terms of downcomer liquid temperatures because of vent-valve fluid mixing and near-saturated condit ions in the primary system. It is recommended that future calculations consider a wider range of operator actions to cover the spectra of overcooling transient sequences more completely.
1
I. INTRODUCTION AND
SUMMARY
Pressurized thermal shock (PTS) in pressurized water reactors (PWRs) has been ident if ied by the Nuclear Regulatory Commis s ion (NRC) as an unresolved safety issue (USI A-49). Because of this, the NRC has a major program distributed among several organizat tons to help resolve the PTS issue. The goal of this project is to determine the potential risk of older reactor vessels to severe overcooling t ransients that rapidly cool the primary syutem.
The Los Alamos contribution to this project is to use the mult1-dimensional, two-fluid, nonequilibrium numerical simulat ion code, the Transient Reactor Analysis Code (TRAC)-PF1,1 to provide accurate thermal-hydraulic cond it ions during postulated PTS acc ident s in selected PWRs. This report presenth the results of several TRAC-PF1 thermal-hydraulic calculations performed for the Oconce-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 calculat ions was to determine which of the overcooling t ransient s specified by Oak Ridge Nat ional Laboratory (ORNL) was the most severe in terms of cold liquid temperatures in the downcomer EE region of the reactor vessel. These ORNL-specified transients are listed in Table I.
The concern over PTS arises because the material propert les of the vessel wall change after several years of irradiatfon.2 The vessel wall becomes embrit t led and its nil-duc t il it y temperature (NDT) increases. If, during an accident, overcooling of the primary-system liquid cools the vessel wall helow the NDT (the NDT for Oconce-1 is ~365 K) and the system subsequently repressurizes, defects could be initinted or propagated in the vessel wall.
Such overcooling of the primary-system liquid may result from the high pressure inject ton (llP1) system or rapid cooling by the secondary system.
Because the risk of init tat ing or propagating flaws in the vessel wall depends on the coupling of the thermal stresses produced by overcooling with the mechanical stresses f rom repressurizat ion, det ailed system models are required.
Modeling both the primary and secondary systems of the reactor plant is necessary to properly analyze the PTS phenomena. The nteam generator (SG) necondary-side inlet condit ions directly af fect primary temperature, pressure, and the emergency core-coolant injectfon. Secondary-side inlet cond it ions are highly dependent on main feed pump and SG cont rol-valve operations as well as the termination of the extracted steam supply to the feedwater heaters. Other 2
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 (MSLB) 34-in. steam-line break; all systems operate as designed; steam generators isolated at 10 min, unaf f ected steam generator refilled at 15 min.
- 3. Small-break loss-of-coolant Pressuriser relief valve sticks open; accident (SBLOCA) ICS f ails to run back main feedwater; power-operated relief primary coolant pump trip.
valve (PORV) stuck open
- 4. Turbine-bypass valve (TBV) One bank of TBVs fails to failure (one bank of two valves) resent af ter opening.
- a. SG 1evel control fails
- b. SG 1evel control does not fail
- c. Reactor coolant pusp (RCP) restart; HP1 throttled
- a. SG 1evel control fails
- b. SG 1evel control does not fail
- 6. SBLOCA 2-in.-dias. hole in (2-in. hot-leg break) pressuriser surge line; RCP trip; all systees operate as designed.
- 7. SBLOCA 4-in.-dias. hole in (4-in. hot-leg break) pressuriser surge line; RCP trip; all systems operate as designed.
- 8. Rancho Seco-type Initial loss of feedwater transient followed by runaway emergency feedwater to both steam generators.
3
l important systems modeled in the TRAC input deck include a model of the B&W Integrated Control System (ICS) used at the Oconee-1 plant. The ICS monitors the primary flows and temperatures to determine the feedwater demand. It also regulates the main and startup flow-control valves (SUFCV), the main feedwater (MFW) pumps, and the turbine-bypass valves (TBVs). Details of the ICS are pre-sented in Appendix A.
Several overcooling transients have been identified by ORNL,3 and additional transients may be specified after these initial results are evaluated. The initial transients include a main steam-line break (MSLB) with a delay in isolating the affected steam generator, a small-break loss-of-coolant accident (SBLOCA) [ full-open failure of the power-operated relief valve (PORV)]
with failure of the ICS to throttle the MFW flow and trip the reactor coolant pumps (RCPs), and TBV transients with SG overfeed. An actual plant transient (Oconee-3 turbine trip) was also simulated by TRAC to compare with actual plant data to verify the code 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 downcomer fluid mixing.
Except for the small hot-leg-break cases, all calculations showed significant primary-system depressurization followed by repressurization if the HPI system was not throttled. System repressurization did not occur for the SBLOCA cases because the break sizes were too large. Some overcooling was obtained in all calculations, as evidenced by highly subcooled liquid temperatures in the downcomer. The most severe transient in te rms of overcooling was the TBV transient in which both banks of TBVs were assumed to fail open, and the least severe was the PORV-LOCA transient. Table II summarizes the key results calculated for these overcooling transients. Not all the transients were run to 7200 a because of computer-tima limitations. Once 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 influence 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 actions should be considered to fully cover all possible overcooling scenarios. Also, in the case of the SBLOCAs, other break sizes and locations should be investigated.
4 l
l l
l
TABLE II TRAC OCONEE-1 TRANSIENT RESULTS Minimus Minimum Downcomer Cold-Leg System Temp. Pressurea Temp. Repressuri-Transient (K) (MPa) (K) izat ion?
- 1. MSLB
- a. RCP restart; HP1 475 6.5 Loop A 475 Yes throttledb 1. cop B 400
- 3. TBV failurec (one bank)
- a. SG 1evel control 365 17.0 Loop A 448 Yes fails Loop B 375
- b. SG 1evel control 440 17.0 Loop A 477 Yes does not fail Loop B 430
- 4. TBV failure" (two banks)
- a. SG 1evel control 350 17.0 Loop A 441 Yes fails Loop B 446
- b. SG 1evel control 465 17.0 Loop A 463 Yes does not fall Loop B 465
- c. RCP restart; HP1 350 4.0 Loop A 465 No throttled Loop B 465
- 5. SBLOCAc (2-in. hot leg) 425 1.0 Loop A 370 No Loop B 410
- 6. SBLOCAc (4.in. hot leg) 320d 0.5 Loop A 430 No Loop B 425
- 7. Rancho Seco-type 452 14.0 Loop A 450 Yes transiente Loop B 450 aThis is the system pressure at the time of minimum downcomer temperature.
bror these MSLB calculations, the EFW system did not actuate because of input errors in the ICS and trip logic.
eThese calculations were extrapolated to 7200 e and the temperatures and pressures shown represent estimated values.
d For this calculation, the minimum temperatures corresponded to temperature
" spiking," as a result of accumulator injection. 5
- 11. TRAC INPUT MODEL DESCRIPTION AND STEADY-STATE RESULTS At the time of these calculations, the Oconee-1 model developed for the PTS study represented the most comprehensive podeling 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 uver 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 f rom ~300-590 K
(~80-600oy),
A. Primary Side The primary side of the Oconee-1 model is similar to other TRAC models that have been used (Fig. 1). It censists of the three-dimensional vessel, two hot legs, two once-through steam generators, four cold legs, and some parts of the emergency core-cooling system (ECCS) [the low pressure injection system (LPIS) was not modeled) .
The volumes, elevations, and pipe lengths from the Oconee-1 plant are 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.
Ileat transfer from the pipe walls to the environment is also allowed.
- 1. Vessel. The three-dimensional vessel is made up of 96 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 of four levels; the upper plenum of two lovels; Und the upper head region contains one level. he two radial rings are dividel so th,t the core region is bounded by the inner ring, and the downcomer annulus is modeled in the outer ring.
De vessel metal structures and wall masses are modeled by assigning 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. he 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 side 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.
6
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Primary-side inodel for Oconee-1.
1
The six azimuthal divis ions were chosen to allow for the angle and separation of the penetration of the hot and cold legs. Also, the accumulat o r injection ports were modeled close to their exact positions (both axially and azimuthally) in the vessel downcomer region.
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. Wit hout finer noding of the vessel, j the upper head is effectively isolated from the rest of the vessel in many c ircums t ances. To alleviate this problem, a connec t ion 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.
The vessel includes vent valves that are modeled to allow flow from the F h
upper plenum directly to the downcomer. This flow path is only available when the upper plenum pressure is higher than the downcomer 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- ,
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at 0 08 4 82 AP(INSIDC-OUTSIDE) (PSI)
Fixed closed value allows leakage for AP<0.08 Fixed open value for AP>0.12 Cloacd value based on design Icakage (W. Jensen, NRC)
Open valuu based on 1%W - 1628 table.
Fig. 2.
Vent-valve model.
d
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).
- 3. Steam Generators. A cnsplete mode" . for B&W once-through steam get.arators 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 SG tubes to the secondary coolant.
Cells 2-11 (Fig. 1) represent a composite of the volume inside and wall area of all tubes, and it is in these c sils that heat transfer from the coolant to the walls of the tubes occurs. Cells 1 and 12 model the upper and lower plenums.
The wall area and t hickness of the plenums 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 into four components (Fig. 3): two model the tube-bundle region, one models the downcomer, and the final one models the steam-outlet annulus. The tube-bundle region is made up of 10 cells (SG components 12-1 and 12-2 for loop A, and 2-1 and 2-2 for loop B in Fig. 3). These cells model the tot al volume and the tube wall area for the region between all the SG tubes, but model the heat-transfer characteristics of a unit cell of tubes. The top 4 of these 10 cells compone the superheated-steam region of the model during normal, full power operation. The connect ion at the top cell of the tube-bundle region is for the emergency feedwater (EFW) flow and the realigned MFW flow. This connection closely models the correct plant geometry so that any flow from the " upper header" is correctly injected into the tube bundle.
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The SG downcomer consists of seven cells (SG components 12-3 and 2-3 of Fig. 3), plus one cell for the MFW injection and one cell for the aspirator port. During norma 1' operation, the downcomer condenses enough superheated steam drawn through the aspirator port to heat the main feedwater to saturation temperature before steam enters the tube-bundle region. The final six SC 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 insuf ficient informat ion regard ing pressure tap locations on the SG secondary side, the SG water levels were modeled using a collapsed 11guld level calculat ion. This method is adequat e for transients in which relat ively 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 may not be accurate for a violent secondary-side transient.
- 4. Cold Legs. All four cold legs of the plant are modeled. These legs each consist of a loop seal (lowest primary-loop po int ) , a RCP, and the HPI connec t ion. The HPI nozzles are positioned close 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 maintained fixed while they are running during a transient. During normal operation, the four cold legs of this model have symmetric flows.
- 5. Emergency Core-Cooling System. 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 causA the primary pressure to fall below the 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 (500F) water into the primary through the four side nozzles of the cold legs. These nozzles are located so that they enter the main cold legs from the side at an elevation somewhat higher than the centerifne of the cold / hot legr.
The pressure-dependent flow rate of the two loop-B ports is ident ic al, 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 11
open. The initial accumulator pressure was ~4.2 MPa (~610 psia), with a coolant temperature of 305.4 K (90cy).
B. Secondary Side All major components of the secondary side are modeled except for the turbine generator equipment and various valves that are 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 term feedwater train indicates 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 it is delivered to the steam generators.
The condenser is modeled as a large tank with a very large wall area. The thin walls have a high thermal conductivity and a constant-temperature heat sink on the outside surface. This model f ully 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 providing an estimate of the available hotwell inventory. The coolan t supplied to the EFW system is taken f rom the hotwell/ upper surge tank combination.
The hotwell pump (component 51) includes the demineralizer/ aerator section l of the feedwater train. The model accounts for the ef fects of this section by including additional frictional losses and the heat addition to the coolant.
Each feedwater 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 steam / water mixture. The energy input from the extracted turbine steam is modeled by adding a volumetric heat source to the middle node of the steam / water mixture. This 6
12
1 l
heat source is controlled by a trip so that it can be ramped off following a turbine trip.
Se two parallel MFW pumps of the actual plant are combined into a single j ptmp 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 one of a variety of setpoints is reached, as described in Sec. II.C.
The coolant flow from the feedwater train splits to provide flow to the SG control-valve section of each major loop. Because the loop-A and loop-B flow-control valves are identical, a discussion of the va3ves for only one loop is necessary.
- 2. SG Control Valves. A diagram of the control-valve arrangement is shown in Fig. 5. Bis schematic can be correlated to the two control-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 (EFWV) is closed during steady-state operation, but opens and closes as necessary if emergency feedwater is demanded. (See Sec. II.C.l.)
The remaining four valves in Fig. 5 provide the major control for the MW flow. he valve area of the two flow-control valves is controlled by the ICS.
During steady-state operation, the SUFCV is 100% open and the main-flow-control valve (MFCV) is approximately 50% open. he 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 Sec. II.C.1).
When one of the two realignment valves, valves 4 or 5, is open, the other must be closed. During steady-state operation, or any time before the f et.dwater-realignment trip is hit, valve 4 is closed and valve 5 is open.
Following the realignment, valve 4 opens and the MFCV and valve 5 close.
There are several different flow combinations that can occur. During 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 through the SUFCV, valve 5, and check valve 3. If feedwater realignment occurs, the flow is from the feedwater train 13
Energy -Input From Turbine l Extraction Steam q" I
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7 SUFCV-LA w2 CENE TOR S' F' Hea er NI ;F WFCV 1: EFW Check Valve 2: Realignment Check Valve 3: MFW Check Valve 4,5: Feedwater Realignment Valves EFWV: EFW Valve SUFCV: Startup Flow-Control Valve MFCV: ' Main Flow-Control Valve Fig. 5.
SG control-valve schematic.
14
t:
.through the SUFCV, valve 4, check valve 2, and into the EFW header. The I
emergency feedwater may also run so that flows can enter both SG headers, or the
! emergency feedwater may mix with the MFW flows 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. he EFW system is modeled so that it takes water f from the hotwell and delivers it to the EFWVs of both steam generators. The EFW l pump is modeled as a composite of the turbine- and motor-driven pu.nps in the actual plant, and therefore has a large capacity. he EFW is delivered to a tee (component 152, Fig. 3) that splits the flow between the two steam generators.
If both EFWVs are open, the EFW flow is symmetrically split as long as the
] secondary pressure of the two steam generators is equal. If only one EFWV is i o pen, the full available flow is delivered to that steam generator. The EFW i- flow stops if the-hotwell inventory has been depleted.
) 4. Steam Lines. he steam line for steam generator B (SG B) is longer ,.
1 than the steam generator A (SG A) steam line. Other than this difference, the i
f steam lines and their valves are identical. A pressure boundary condition i j models the steam flow exiting the secondary into the turbine inlet. During ;
l steady-state operation, the steam-mass flow is modeled to re-enter the secondary
} as boundary condition inlets to the condenser and the two heater drains.
l Following a turbine trip, the turbine-stop' valves (TSVs) close the steam line s.
} If any pressure relief from the closed steam lines is necessary, steam is [
l released through the TBVs into the condenser. For a normal reactor / turbine-trip
{ transient, the steam relief from the TBVs is adequate so that modeling the main steam safety valves is not necessary. Following a turbine trip, the secondary
{
! model is a closed loop unless a transient similar to a MSLB is modeled. For the t
j MSLB, one .of the turbine inlet boundary conditions is set to constant l containment pressure and the corresponding TSV is fixed open.
- C. Control System The control system for the various components of the TRAC input model is l provided in two ways: with trips and with control blocks. Trips basically turn
, something on or off, depending on certain conditions being met. In this model, control blocks use mathematical relations between system variables to adjust valve areas or pump speeds. h e control blocks are used to model the B&W ICS.
i l
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'l. Trips. Of the more than 40 trips used in this model, only 5 are s imple trips. Simple trips have only a system variable as input, and their
- output is only used to control some component act ion. A summary of the simple trips is presented in Table III. The TBVs are actually controlled 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 i
output from the reactor trip is input to the TSVs, the condenser feed, the feedwater heaters, and the heater drains. Most of 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 occurred immediately. If additional transients were to require the reactor trip
) to occur after other criteria had been met, this model would have to be I mod if ied. The delays and rates are not included in the information presented in Fig. 6, so these parameters will be discussed in the following paragraph.
l TABLE III 1
SIMPLE TRIPS Description System Variable Setpoint [(MPa)/ Action]
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 16
~
The reactor trip is 4.5 s from the. beginning of the transient. To model the insertion of control rods into the core, a negative reactivity insertion of
-0.0536 Ak' 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 TSVs start closing simultaneously with the reactor trip and take
~1.0 s to fully close (except in the MSLB t ransient, where the loop-A TSV is 3
fixed open). The condenser feed trip occurs ~1.0 s after 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 through the heater drains are tripped 4.5 s af ter the reactor trip and also take 5.0 s to decay to
' Zero.
1 111 2511 1 Reactor Trip TSV .A S.V. Time 2559. E >05 0*1 04 1 4
2522 TSV - B 2559. E >05 0+1 2911 Condenser Teed 2559. E >05
, 0-1 2611 Heaters, Drains 2559. E >05 0*1 Fig. 6.
Reactor trip system.
I I' 17 r,
r t
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 input. The component numbers are enclosed in small circles and correspond to 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 to understand the interconnection between them. For further explanation of these figures, please refer to the trip system legend in Appendix A. !
Figure 7 presents the trip system that controls the HPI, the four RCPs, and the MFW realignment valves. HPI initiation occurs if the system pressure drops below 10.44 MPa, but is also cycled on and of f to keep the least cooled hot leg subcooled at 42 7 K (75 12.5 F). For some transients reported, the HPI subcooling monitor was not used. For those cases in which a particular trip is not wanted, it is left in the system but given a setpoint that cannot be reached. All four RCPs also use the HPI low pressure setpoint, but have a 30.0-s delay from the time the setpoint is reached until the pumps trip. The Al and B1 pumps are under separate trip control from the A2 and B2 pumps, so that pumps Al and B1 can be turned back on if the least cooled hot leg reaches 42 K while subcooling. For some of these transients, the pump subcooling monitor is also disabled.
Following any RCP trip, the main feedwater is realigned to flow into the steam generators through the EFW header. This realignment occurs by taking the control of the MFCVs away from the ICS, ramping the the MFCVs closed at a flow area fraction closing rate of 25% per second, and opening or closing the appropriate realignment valves over a 5.0-s period. Further information on this valving arrangement is in Sec. II.B.2.
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.
If this pump trips, then the other two must 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 upstream pump trips, either SG level greater than 9.27 m, a suction pressure less than 1.72 MPa, or a discharge pressure greater than 8.89 MPa.
18
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HPI, RCP, and M W realignment trip system.
i 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 EW valve close if SG 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 SC isolation trips is turned back of f, then EW can be delivered to that steam generator if it is required, such as in the MSLB transient.
The _ controlling trips for the EW system are presented in Fig.10. The trip conditions that demand emergency feedwater ' are at the top of the figure, and the conditions to throttle emergency feedwater are at the bottom. Emergency feedwater is demanded if the MW discharge pressure drops below 5.271 MPa. The EW demand opens the appropriate EW valve at a fractional rate of 33% per 19
=. . . , - - . - . .
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rii.eDr*!NN\
23:1 sy on, Tr _0*1 s 4 233e E >o.s i-
'l 0*1 Fig. 8.
Hotwell, condensate booster (CB), and MFW pumps trip system.
second. If either EFW valve starts to open, the EFW pump comes up to full speed in 4 s. he EFW valves close if the SG water level goes above 6.2 m, but reopen
. when the level drops below 6.0 m. he EFW valves also close if the hotwell level is less than 0.0254 m.- If both EFW valves close, the EFW pump trips of f.
I
- 2. ICS. h e B&W ICS matches the feedwater flow with the power demand and maintains a constant steam-line pressure and adequate superheat. he ICS can
, quickly respond to plant load demands while maintaining smooth plant operating parameters. he block diagram of Fig. 11 depicts how this control is i accomplished. he 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 t 20 1
I lon
( sv => / ["flhr'Lis')
\1
- O 4009. E >05 /
\ ~_.V 2711 3411 3311 EFW Valve A SUFCV A WFCV A Control Dwerride Closed Dverride Closed 3 2778. M >0.5 =- -+ 3419, E <0.5 --* 3319. E >05~C
- 1.-+ 1 o * ~1 0 * -1 e
I lation k O'Y "I Shr a
\1
- O ( 4019. E >05 )
rra 342e sace EFW Valve B SUFCV B MFCV B Control Dwerride Closed Dwerride Closed 3 2719,M >05 *- -* 3429 E <05 --* 3329. E >0,5 7
-1
- 1 0 * -1 0 * -1 Fig. 9.
SG isolation trip system.
the Integrated Master so that, if some section is not performing adequately, the other parts of the ICS balance the response to prevent any power, pressure, or feedwater flow mismatches.
The performance - of the ICS must 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 Sec. II.C.1) and only the feedwater-control section needs to have detailed 21
(
l
{
ischarge Pnen I tant k EY' " l I m
( CA -1994 ) [ ( CA -tied )
G A f IFW Valve A ) f EFW Valve 3 \
Demand l Demand )
--=( FF19. E >05 }- -( 2738, E >05 h N'Y N~Y 271I 2722 EFW Valve A EFW Valve-S Control Control
- 9 2778. M >05 == . _ _
-* 2779. M >OS N j -la 1 3011 -1 I l "W1r Pump Centrol 3000. E >-OS :
0 ++ 1 i
f n-AInvel)
Op. Range .
y f gm3 ,
[St n-B Lavolk Op Range;
( EV. O!! ) [1mvHotwell)
Inventory l
( EV. 010 /
( EV, 016 )
\~o/
v Fig. 10.
EFW trip system.
control-block modeling. If a transient were run with the reactor not tripped, 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.
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-l power cross limiter.
l 22
Auto Dispatch System Unit Load Demand i
l Integrated Master I I l l Pressure Reactor Feedwater Control Control Control I I I I Turbine Turbine- Control- SG A WFW SG B B
Valves Mes Valves Ruggpe Valves Dri es Fig. 11.
ICS organization.
Figure 12 presents the feedwater-control sect ion 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 with the reactor power, so that the feedwater flow can be reduced if the power is dropping. The requested feedwater flow is then compared with the actual feedwater flow to get a feedwater-flow error.
This error value will then open or close the MFCV unless the SG 1evel has 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 aP.,ng with the MFCV pressure drop to determine the MFW pump speed.
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 model. The loop-B model is similar except that it uses several of the signals generated for the loop-A side to avoid duplicat ion of coding. A complete listing of the flow diagram and mathematical . rapresentation of the TRAC-PF1 ICS j model is included in Appendix A.
23 J
w - , - , . - - . v
I l
A diagram of the BTU-limiter control blocks is given in Fig.13. Each of '
the blocks represents one signal manipulat ion. 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 identify 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 the International System of Units. The steady-state output voltage for each of the blocks is given in the table. A BTU-limiter output value of 8 V indicates the correct feedwater flow. For a value less than 8 V, less flow is needed, and for a value greater than 8 V, more flow is requested.
The neutron power cross-limiter sect ion 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 after 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 BTU Limiter Neutron Power Cross Limiter 3t AE Feedwater MFCV Pressure Flow Drop kvel MFW Pump Limiters Speed Control Feedwater Valves Control Fig. 12.
Feedwater-control sect ion of the ICS.
24
-- . . . _=. . - . - . . -
l Hot-Leg A flot-Leg A SG A Feedwater Mass Flow Temperature Pressure Temperature A _B C H
T Fig. 13.
BTU-limiter control blocks.
TABLE IV BTU-LIMITER CONTROL-BLOCK EQUATIONS Steady-State Equation Output Voltage 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 i F = 0.55555 + 0.055555
- E 1.0 H = -10.0 + A
- F 8.0 25
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 ,
for time-step size. The final output from this 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 frc,m block H and the neutron cross-limiter input from block SG. The blocks depict the selection of one limiter signal and the comparison of that signal with the current feedwater flow. The equations that make up this section of the ICS are presented in Table VI. All blocks in this table are output in volts except for block SL, which is in kilograms per second. 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.
Reactor Power Resetor Trip Feedwater Time Temperature 1
SP J'
JL A1 Jl B1 i
g C1 J
i D1 i
E1
.)I Y1
. 1, K1 SG Fig. 14.
Neutron power cross-limiter control blocks.
26 l
l t
.. .___ _ _ . ~ .._. - __.._ __ . _ . - . _ . - _ . _ . . _ . _ .. _ . .m._ _ . . ._ . - __ _ m___.._.. .._-
i I
TABLE V NEUTRON-POWER CROSS-LIMITER CONTROL-BLOCK EQUATIONS l
Equat ion Steady State A1 = TIME--(TIME OF REACTOR TRIP) 0.0 i
- n1 = 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 r
a 5.4 356.0 l
9.36 402.0 18.0 460.0 l' 21.42 483.0 1
i El = -460.0--D1 + 1.8
- FWTEMP 0.2 i F1 = 1.0 + 0.0013
- El 1.0 l
)i G1 = F1
- C1 18.0 SP = POWER--2568.0e6 0.0 4
1 4 First-order lag of power with 4.5-s time constant JL = JL + ((SP--JL)/4.5)
- DELT. 0.0 i
l J1 = 2568.0e6--JL 2568.0 x 106 l
} H1 = 1.6 + 14.4
- B1 16.0 I Il = -1.0 * (H1--6.23053e-9
- J1) , -10.0 < Il < 10.0 0.0 i
K1 = f(II) : Il K1 0.0
, -1E5 -1E0
- 0.5 0.0 r 0.5 0.0
- 10.5 10.0
4 SG = -10.0 + K1 + C1 8.0 i
I i
4 27 i
i
,. , . , - = .- - . . - , , . - . . . ,-...------..-.,.n..-~, . , - - , - - , . .n.,~n , , , , -
- . . ~. .
- - - __- _ _ . - . . _ . - ~ _ . _ - . .
4 t
t Loop-A BTU Neutron-Power Feedwater Flow Limiter, H Cross Limiter, SG l
FWB -
4
- R :
A --
FSL SL l
i _
- S1 :
. T i
Fig. 15.
Feedwater-flow control blocks.
E TABLE VI i
FEEDWATER-FLOW CONTROL-BLOCK EQUATIONS t
i Equations Steady-State Value R = min (SG,H) 8.0 FWB = FWFLOWA--680.4 0.0 ;
First-order lag of feedwater flow with 1.0-s time constant FSL = FSL + ((FW8--FSL)/1.0)
- DELT 0.0 SL = FSL + 680.4 680.4
$ St = 10.0 + R--0.026455
- SL 0.0 i
4 28
, , . ...,-.m -, - - - . .~.-,.m.._.- ., + . - , . , - - m_,,, ,.-, , - - -
The SG 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 high-level limit block (PL) 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 RCP trip has occurred, the higher value (6.096 m or 240 in.) is used to enhance natural circulation by maintaining a higher SG level. When the low-level limit is reached, P2 changes 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. Table VII gives the equations and steady-state values that correspond to these control blocks.
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 of 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 the ICS is low-level limited or not. The equations describing this action are given in Table VIII. If the ICS is 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 f rom both loops. The resulting voltage obtained from the comparison between the BTU-limiter output and the neutron cross-limiter output 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 are in Table IX. There is a constraint on the rate of change of pump speed built 2
into the TRAC-PF1 MFW pump model (27 rad /s ), so that an additional constraint was not needed in the ICS model.
D. Steady-State Calculation The primary-side steady-state operating conditions for the TRAC model are presented in Table X with operating specifications f rom the Oconee-1 plant for comparison. The two primary-system loops and all four cold legs had symmetric flows. The mass flow rates through the reactor pumps were controlled during the 29
i l-TABLE VII r
LEVEL-LIMITER CONTROL-BLOCK EQUATIONS Equations Steady-State Value
- operating level scale, 96 to 388 in. (level in meters)
HL1 = f( ALEV) : ALEV HL1 -10.0 2.438 -10.0 9.855 10.0 1
P1 = -2.0 * (HL1--7.0) 34.0 Q1 = min (S1,P1) 0.0
- startup level scale, 0.0 to 250 in.
LL1 = f(ALEV) : ALEV LL1 -3.3 TCTI -10.0 6.350 10.0
- pumps tripped: 240 in. = 6.096 m = 9.2 V IF(PTRIP .EQ. 1) STP = 9.2 pumps running: 24 in. = 0.61 m = -8.08 V IF(PTRIP .NE. 1) STP = -8.08 -8.08 P2 = -2.0 * (LL1--STP) -9.6 T1 = max (P2,Q1) 0.0 steady state by adjusting the pusp 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 pressure d rops. In Table X, 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, s teady-state operating conditions for this model. In general, the comparison -is very good with available Oconee-1 plant data. The feedwater-flow rates for the two loops were independently controlled by the ICS through control of the MFW pump speed and MFCV areas. The steam-outlet flow rate ' was approximately 2% higher than the 30
i l
Feedwater-Flow SG A RC Pump Comparison, S1 Water Level Trip Signal o
llL1 LL1 STP
_L n F_-
-8 G=
j Fig. 16.
l Level-limiter control blocks.
i Feedwater-Flow Low-Level Error, T1 Limit Signal. P2 h jCNST1 .----
t X 11l= CNST2 ----
r-I MFA i
[SUFVA (MFCVAl 1 1 to Comp to Comp 30 36 j Fig. 17..
Feedwater-valve adjustment control block. 3t
r TABLE VIII I
FEEDWATER-VALVE ADJUSTMENT CONTROL-BLOCK EQUATIONS i
. i
, Equation Steady-State Value i
- if P2 < 0, lov limit has not been hit IF(P2 .GE. 0.0) CNSTI = 0.12 0.1125 IF(P2 .LT. 0.0) CNSTI = 0.1125 IF(P2 .GE. 0.0) CNST2 = 2.4 0.9 IF(P2 .LT. 0.0) CNST2 = 0.9
- integrate, Ti ois the last time-step value of signal T1
]
U1 = U1 + CNSTI * (T1 + Ti o )/2.0
- DELT , -18.0 < U < 2.0 0.0 l X11 = U1 + CNST2
- T1 0.0 X1 = X11 + 8.0 , -10.0 < X1 < 10.0 8.0 s
- SUA = 64.1164 + 7.44164
- X1 , -10.0 < SUA < 10.0 10.0 i
i SUFVA = 0.1
- SUA , 0.0 < SUFVA < 1.0 1.0 l
j MFA = 0.5555
- X1--4.4444 , -10.0 < MFA < 10.0 0.0 I
T MFCVA = 0.5 + 0.5
- MFA , 0.0 < MFCVA < 1.0 0.5 4
i i feedwater flow, so that the SG inventory is slowly depleted with time. The
- approximate t ime it takes coolant to travel from the hotwell to a . steam
, generator is estimated from the average velocity and piM length to be over
't 11 min.
b 32 d
. - , -v - , - - , . - , . , , - -
l Loop A Limiter Loop B Lirniter MFtV-A MFCV-B Comparison, R Comparison, BR AP AP ,
1 1 IDPABl lDPSBl I A 4 DPAL DPBL A A DPA DPB
FC
I FD I
g FE FI =
iFC T to Comp 49 Fig. 18.
MFW pump-speed control blocks.
i 33
TABLE IX MFW PUMP-SPEED CONTROL-BLOCK EQUATIONS Equat ions Steady-State Value DPAB = DELPA--3.55e5 0.0 First-order lag with a 1.0-s time constant
- 40 psi limit on both sides DPAL = DPAL + ((DPAB--DPAL)/1.0)
- DELT , -2.4E5 < DPAL 0.0
< 2.4E5 DPA = DPAL + 3.55E5 3.55E5 FA = 2.90074E-5
- DPA--10.0 0.2975
- the loop-B pressure-drop equations are the same FC = min (FA FB) 0.2975 FD = FC--0.2975 0.0 FE = 0.2
- 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 $23.6
-270 370.4 0.0 392.8 6.0 460.0 10.0 586.43 34
. _ .. . ~ . . . - - . . _ . _ . .- .. _- _. - __... ..-. _ _ - - - - . . --.
IABLE X l
PRIMARY-SIDE STEADY-STATE CONDITIONS l
Oconee-1 Parameter TRAC Model Nuclear Station Power (MW) 2568.0 2568.0 1
Coolant flow rate, total (kg/s) 17640.0 17640.0 i Hot-leg temperature (K)a 589.56/589.33 589.3
- Cold-leg temperature (K) 563.81/563.81/ 563.5
- 563.61/563.61 l Primary pressure (MPa) i (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 temperature (K) 305.4 305.4 f I
l Hot-leg coolant velocity (m/s) ~19.5 -
i l
i Coolant flow path length (m) 1 (external to vessel) 63.5 -
1 1
aloop A/ loop B or cold-leg A1/A2/B1/82.
1 i ;
I l
r t
r i
35
TABLE XI SECONDARY-SIDE STEADY-STATE CONDITIONS Oconee-1 TRAC Model Nuclear Station l i
Feedwater flow, loop A/ loop B (kg/s) 679.04/678.38 680.4
< Feedwater 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 l Steam pressure at turbine inlet (MPa) 6.2/6.2 6.2 l
! SG secondary inventory (kg) 1.711E4/1.715E4 ~1.77E4 i
l Aspirator steam flow (kg/s) 95.4/95.3 -
1 l MFW pump inlet pressure (MPa) 2.77 2.68 -
] temperature (K) 461.0 462.0 l
Condensate booster pump inlet pressure (MPa) 8.0 6.9 i temperature (K) 308.3 309.0 Hotwell pressure (MPa) 0.015 0.01 temperature (K) 305.6 305.9 inventory (kg) 5.31E5 5.31ES Upper surge-tank inventory (kg) 2.69ES 2.98E5 '
- Feedwater train average coolant ~1.5 -
! velocity (m/s) i Hotwell to SG flow length (a) 1004.3 -
MFCV area fraction (t) 48.60/48.15 50 Loop-flow fraction (%) 85/85 85 l
t t
1 e
36 t
l I
- . . . - . - - - . , . w . , .. . . ,,,,e , ,
III. TRAC TRANSIENT CALCULATIONS i A. Oconee-3 Turbine Trip 1
l 1. Introduction. As a benchmark case for the Oconee-1 PTS s tudy, the Oconee-3 turbine-trip and SG overfeed transient of March 14, 1980, was simulated. The actual transient is documented in Ref. 4, and measured data are available for the first 3 min of the transient.* The data available include l primary- and secondary-system pressures, hot- and cold-leg temperatures, pressurizer and SG water levels, and MFW flow rates and supply temperatures.
In this transient, the plant was operating at 100% power when the electro- I hyd raulic 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. He SG water levels increased above the expected levels until the MFW pumps tripped on an SG high-water-level signal.
De 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) . He TBVs were also modeled. ne condensate-heater /feedwater-train modeling and the ICS modeling was not included for this transient because these systems were not necessary. Instead, the measured MFW 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. We actual transient had slightly more overcooling than the TRAC calculation. The calculated primary pressure, hot- and cold-leg temperatures, and pressurizer water level were slightly higher than the measured data. He major differences between calculated and measured results were in the SG secondary pressures and in the SG B water level. He 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.
" Data not shown in this report because they are proprietary.
37 m
The seasured secondary-side water level of SG B was found to be inconsistent with the measured MFW flow rate. H e steam generator refilled at a
( much faster rate than could be accounted for by just the MFW flow, which i
l indicates that auxiliary feedwater might have been inadvertently delivered to SG B during the transient. His inconsistency was not mentioned in Ref. 4, nor were 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. He basic TRAC model of the l
Oconee-1 PWR, Fig.1, was used to model the Oconee-3 transient. He steam lines l 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 4
I specified as boundary input in the TRAC calculation. He standard ANS decay power, Fig. 21, was used and is automatically calculated by TRAC. The Oconee-3 4
} measured reactor power is also shown in Fig. 21. He Oconee-3 power curve was used in a second TRAC calculation to determine the effect of reduced decay j power, even though it did not include gamma-ray heating.
I 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. Rese 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 1
transient. He primary-coolant pumps did not trip in this transient. No l
auxiliary-feedwater flow was assumed in either steam . generator. The HPI flow was assumed to be divided evenly among all four cold legs.
! well with the actual plant transient. De actual transient had slightly more overcooling than the TRAC calculation. He calculated primary pressure,
- pressuriser water level, and hot- and cold-leg temperatures were slightly higher 4
than the measured values. In general, differences between calculated and
- measured values were consistent throughout the transient.
l i
i 38 i
i l
l l
Ow(ii) @-
@ 7ao2cDoso @
31E 60 @
sti ti 2 ' 3 ' 4 i 5 i ro iB2 C Mssvs
@ @ _]904[ 79:scDo70 #F) '
s e.am ,
sc 4 "*
- g "
__391609170 @)
J91089120 @
J90eC ig4 {
l' Hest 3D @
4111121 3 i 4 i 5 I 6 1 7 from Il @ b 716140 @ '
SG B S
L[gMg 2 M
,- @ m es:52 o @
3 Ma D
~
8@n*j 4l
. ! TGV -
l @
5' I 6 188I 1 x 2 s7 tO @
5 1 6 1871 1 X 2 1860[ h 8TBV Fig. 19.
Main steam safety valve modeling for Oconee-3 transient.
%D i
soo , , ,
nso Loop A soo- .
.Leoc 8 -. moo uso l ? ?
, s neo- . s I i moc 2
h 400- .
no I h .
h S a
. see 20o- .
2sc
'.. . . , . . . . . . . . , . . ^ ..
o . 3 0 20 40 SC 53 100 20 HQ SC SC 200 Time (s)
Fig. 20.
Measured MFW flow rates.
c:
coicuto+.d aus decoy pe-er
-- weasse d secoy po.er (Do not include gefrwn0* rey Mti%) ,
Q 40- -
w I
E p.
w 40- -
3 4
go. .
O 20 40 SC 80 90 90 WO SC 53 20o Time (s)
Fig. 21.
TRAC-calculated decay power and Oconce-3 measured thermal power.
40
TABLE XII f
INITIAL STEADY-STATE CONDITIONS Parameter TRAC Oconee-3a Reactor power (100% power) 2568 MW Cold-leg ~ temperature 563.0 K (554 0F)
Hot-leg temperature 588.9 K (600 0F)
Primary-mass 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)
MFW flow 680.0 kg/s (1500 lb/s)
Feedwater temperature 510.9 K (460 0F)
" Reference 4 (proprietary data).
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 MPa (1881 psia), whereas the actual transient decreased to a slightly lower minimum. The calculated pressurizer water level dropped about 3.04 m (10 ft).
Figures 24 and 25 show the hot- and cold-leg temperatures of loop A and loop B, respectively. Temper'tures in loop A were slightly lower than in loop B because of a higher MFW flow rate in the loop-A steam generator. The init ial increase in cold-leg temperatures in the first 10 s of the transient resulted from 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.
41
TABLE XIII MSSV AND TBV SETPOINTS The following setpoints are FSAR values and were used in the TRAC l calculation. Actual Oconee-3 setpoints were not available.
! 1. ' MSSV Setpoints l
l Pressure (MPa) Pressure (psig)
! Open Close O_ojg1 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. TBV Setpoints l Pressure (MPa) Pressure (psig) l Open Close Open Close 7.067 6.998 1010 1000 TABLE XIV SEQUENCE OF EVENTS l
The following times are actual transient times and are the event times used in the TRAC calculation.
Event Time (s)
- 1. Turbine trip and reactor trip occurred. 0
- 3. HPI pump A started to assist pump B to maintain pressurizer water level. 30
- 4. MFW pumps tripped. 103
- 5. End of plant data and' calculation. 180 42
sae' ,
2xo caisuisted
...***" hieseur ed 2203 s
- u. _
2ec
- n m t 2
E un'- i:
- 2000 S
. i
.R
- '- uno 3
j
! ue'.' i s.
- a. a,
.'.., ...~,
.** o 00 j
_ _ . l ue'- .
1700
. ggg iMC' -
0 20 40 60 80 EDC 20 WO oC 10J 200 Time (s)
Fig. 22.
Calculated and measured primary pressures.
~
Colculated 7- ......**~ esessured -
. 20 4-a
^
o w
E g.
O %
e D - ,
$ e- I.
t i.
- . 10 o 2- .,
- 8 i
'.' =
2- =
t-0 0 . .
SC SO 200 0 20 40 4G 80 10 0 90 WO Time (s)
Fig. 23.
Calculated and measured pressurizer water levels.
43
1 800 , , gro <
l Celculeted hot-leg temperature 1
- .- Celsulated cold-leg temperature
...-.- essener ed hot-leg temper e f ure too. - - teessured 8018-186 temperature -
i a l 5 >
i e '
6 t
t 2 se0- . ~
~
3
.sec
- st0- ~* 'i - b 2 .h u0
}
a } %%'-
^' "
u0- %N?!*- . .=. .E_ ~3 -~2 7l"" *
- , - - ~
i 540 SSO 4 0 29 40 60 80 10C 90 WO SO ISD 200 3 Time (s)
Fig. 24.
Calculated and measured hot- and cold-leg temperatures for loop A.
800 s20 Calcuteted e e t-les tempereture
--- Celcuteted cold-leg tempereture
--. --- weesor.d ho t-ses t emper s f ure no- - --- . woesured colo-seg temperature -
s00 a m
V u -
y w
- i. !
2 m- .
2 2 i we
- i st0- i -
-=
\,,%.w- . u0
-& J , r:- - .&.
s
," ...'. %' ;- -. - _ _ _ _ s
~
se0- .
. .... . .. .: . - :~. . -
s40
&&O . .
9 20 40 00 00 100 WO WO 90 100 300 Time (s)
Fig. 25.
Calculated and measured hot- and cold-leg temperatures for loop B.
t l
44 1
Figures 26 and 27 compare actual and calculated secondary pressures for
! loop A and loop B, respectively. The pressure peaks at about 6s were also caused by TSV closure. TRAC calculated lower peak pressures because of a modeling error in the location of the TBVs. 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 I
pressures cycled between the TBV open and close setpoints. The actual' pressures l decreased below the TBV setpoints and remained below the setpoints for the duration of the transient. The reason for this difference is not clear. )
Possibly the Oconee-3 plant had different TBV setpoints and rate characteristics from those modeled in the calculation.
Figures 28 and 29 show the secondary-side water levels for SGs A and B, respectively. There was excellent agreement between calculated and measured water levels in SG A, but considerable difference in SG B. Subsequently it was found that the measured water level and MFW flow in SG B were inconsistent. The steam generator was refilling at a much faster rate than could be accounted for by just the MFW flow. This indicates that auxiliary feedwater may have been inadvertently delivered to SG B. This inconsistency was not mentioned nor were auxiliary-feedwater flow data given in Ref. 4.
Table XV compares the calculated and measured minimum pressures, temperatures, and pressurizer water level reached in the transient. Addit ional minimum-value results are presented for another TRAC calculation in which it was assumed that the measured Oconee-3 thermal power (Fig. 21) was the decay power.
The Oconee-3 measured thermal power did not include gamma-ray heat ing.
The results do show the sensitivity of the transient to decay power.
- 4. Summary. The Oconee-3 turbine-trip and SG overfeed t rans ient of March 14, 1980, was modeled using the TRAC-PF1 code. The calculated results compared very well with the measured data. Differences between calculated and measured results were minor and consistent throughout the transient. The actual transient had slightly more overcooling than the TRAC calculation. The calculated pressures, loop temperatures, and pressurizer water level were slight 1!y higher than the measured values.
45
7AC' ,
Colcuteied ,w the*. . . geeeeerog .
7,4.g e. ,
( - 'M e
=o 12 9 , ,
T.
b E 3.. [vN M .
- 2' Y 8
E i
'"5 E.
l 2 s.e-e*- .
E l l l
4 6 C'-- . 963 g a.c.. . SSO sy;* 900 t 20 40 6C 80 WC UC w0 eC 143 200 Time (s)
Fig. 26.
Calculated and measured SG A secondary pressures.
< 7,e t'
- Coteviated ,,
74c'. ...... .... age ,,,, e d *
'W 3.4.co . ,
f q , .e.c =e
?
E E
?.DC'- -
l
. E a
S 6 C'- * *EI 6 e C'- . 950 u.c' eso o so ao se ao oc uo wo se so neo Time (s)
Fig. 27.
Calculated and measured SC B secondary pressures.
46
e .
Cercureted 30
- ====**** Nessur ed
- s. -
3s
~ n
~ . . - .
.E., -
~
.. . so \
. 1 1 .- 5 6
s 3
2 a. - -
$ . 3
. .g .
j ..
3 -
.s e 8 o ao ao u so no no wo oc as too Time (s)
Fig. 28.
Calculated and measured water levels in SG A.
t cercutee.d se
... . .... u. a a , .e s.
2S
- ^
E -
w -
.. . so
. a a "
, ~# 6
- s. .***.....~...........-. 2 .
o a 3
, . . . . . . . . , , . . . . . * . e. < a* .
- g. :wv.._ -
-5 e . -
8 3 3e es to 80 30 US WO 90 50 300 Time (s)
Fig. 29.
Calculated and measured water levels in SG B.
47 L
TABLE.XV COMPARISON OF TRAC AND OCONEE-3 RESULTS
- 1. With ANS Decay Power in TRAC Calculation TRAC Oconee-3a 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) l Minimum cold-leg temperature (Loop A) 560.6 K
{ (549 0F) i
- 2. With Reduced Decay Power in TRAC Calculation A TRAC calculation was performed using the Oconee-3 measured reactor j power as a decay power curve input to TRAC. The Oconee-3 reactor power curve I 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) ;
j Minimum pressurizer water level 2.12 m I (6.96 ft)
Minimum hot-leg temperature (Loop A) 557.8 K
- (544 0F) j Minimum cold-leg temperature (Loop A) 558.0 K (545 0F)
- aReference 4 (proprietary data).
j An inconsistency in the measured data given for SG B was found. The steam generator was refilling at a much faster rate than could be accounted for by 1
just the MFW flow. This indicates that auxiliary-feedwater flow may have been inadvertently delivered to SC B in the actual transient. - If it was assumed in l
the TRAC calculation that auxiliary feedwater was -delivered to SC B, the calculation would have agreed better with the data. Other factors that could t
48 P
. , _ . . , . , . - , . - - - .- ~ - * - , , ,e-,. -,.y, - , , , . - - , - - , , , - . , - ,--~y-y
. wm----c--.w,-.y.w-- ,-,-4 =.rr-~4 -+- - -
affect the degree of overcooling are decay power, HPI flow, TBV setpoints, and valve-rate characteristics.
B. The MSLB l 1. Introduction and Summary. For this transient, the overcooling of the l
primary side of the plant is caused by a seve re depressurization of the secondary side. The secondary-side depressurization is caused by a full double-ended steam-line break in SC 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 wain forcing function for the overcooling of the primary l side is the delay by the operator in isolating the main feedwater and emergency feedwater 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 (Case 1) had all of the ICS, protection, and emergency systems operate as designed. In Case 1, the operator is assumed to isolate all feedwater to both steam generators 600 s into the transient and then to restore the unaffected steam generator at 900 s. Also, the operator restarts one RCP in each loop after attaining 42 K subcooling and throttles the HPI to maintain 42 12.5 K fluid subcooling.
Three parametric cases were analyzed in addition to the base case. Case 2 was identical ta 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 failing to actuate as designed, the RCPs never restarted af ter the subcooling margin was reached because of input deck error. In Case 4, the MFW pump was tripped at 0.5 s and the subcooling monitor for restarting the RCPs was coved 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 could occur during a MSLB transient.
- In terms of downcomer fluid temperatures and primary system repressurization considerations, the MSLB base case was one of the most severe of all the specified ORNL transients. A relatively cold downcomer liquid temperature of ~405 K, and hence a small margin against the NDT limit were calculated for the base-case MSLB transient. Repressurization of the primary cystem to the PORV setpoint (~16.9 MPa) was also calculated for the base case.
49
- 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 diagrams are shown in Figs. I 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 open, 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 as follows:
- 1. Full reactor power.
- 2. Nominal temperatures and pressures in primary / secondary.
- 3. Decay heat--1.0 times ANS standard.
- 4. Reactor and turbine trips coincident with MSLB.
- 5. Operator fails to isolate feedwater to both steam generators until 600 s.
- 6. Operator restores the unaffected steam 8enerator (SG B) at 900 s.
- 7. RCPs restarted after 42 K subcooling reached.
- 8. HPI throttled to maintain 42112.5 K subcooling.
i I
l 50
i l
l j 3. Results. ,
l 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) l 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 i
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.
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 the condenser feed from the turbine. The turbine and reactor were then tripped followed by a feedwater-heater flow / drain trip. At ~5 s, the TSV in loop B opened af ter the TBV setpoint. (7.06 MPa) was reached. The TSV 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 MPs. At ~29 s, the ICS detected a low-level !
limit in SG A and the EN pump was started. The EN valve to SG A was opened [
and EN flow was initiated. At ~47 s, the MN pump tripped on low-suction !
t pressure. The ICS detected a low water level in SG B at ~48 s, and EN flow was initiated af ter the EN valve in loop B opened. At ~51 s, the RCPs were tripped (30 s after HPI initiation), and the feedwater realignment trip occurred (all feedwater directed through the EN 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 EN valve to SG B was shut. All of the E N 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 51 were restarted at this time because the subcooling margin was sufficient. At ~530 s, the primary systea had depressurised 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. As a result of the transient specifications,3 the SG secondary sides were isolated at 600 s.
Because of reverse heat transfer in SG B (heat transfer from the secondary to the 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 EN and TSV systems to operate if needed. However, because the 51
_ _ ~ . . - _ _ _ - - _ _ - - . . _ . - - _ - - - --.- ~_ - _~ ,
l
?
l TABLE XVI SEQUENCE OF EVENTS l EVENT TIME (s).
- 1. MSLB--loop-A steam line 0.0
- 2. Turbine and reactor trip; TSV 0.5 loop B closes
- 3. TSV loop B opens (setpoint 7.063 MPa) 5.0
- 4. KPI initiation (setpoint 10.44 MPa) 21.2 r
- 5. SG A low-level limit reached; EFW 29.4 pump starts; loop-A SG EFW flow initiated i 6. TBV loop B closes 39.9 i l 7. MYW pump trip on low- 47.8 I suction pressure l
, 8. SG B low-level limit reached; 48.7 loop-B EFW flow initiated
! 9. RCPs trip (30 s after HPI initiation) 51.2 i 10. CB pump trip (low-suction pressure 53.9
{ 11. SG B level at 50%; loop B 346.7
! EFW valve closed l 12. RCPs ( A1, B1) restart (42 K subcooling 526.0
. reached); HPI throttled 2
- 13. Loops A and B accumulator setpoints 530.9 [
reached (setpoint 4.17 MPa) '
i 14. Loops A and B accumulators off 537.9 l 15. SGs A and B isolated; EFW pump and 600.0
) hotwell pump tripped off '
i 16. SG B restored 900.0
- 17. PORV setpoint reached (setpoint 4678-7200.0 I l
- 16.9 MPa)--PORV opens and !
closes for remainder of calculation !
- 18. TSV loop B opens (setpoint 5462-7200.0 !
7.063 MPa)--TBV opens and [
- closes for remainder of calculation
! to maintain setpoint pressure ;
i 19. SC B level drops below 50% 6121-7200.0 operating range; EFW initiated-- t EFW pump on/off for remainder of '
calculation to maintain level at 50% ,
- 20. Calculation terminated 7200.0 I SC B level was at the 90% limit and the secondary pressure was low, these systems did not actuate untti auch later in the transient. Following isolation of the steam generators, the primary system began to repressurise. At ~4678 s, j the PORV setpoint was reached (16.9 MPa), ard the PORV cycle opened and closed j for the remainder of the transient. The secondary side of SG B repressurised to 1
- 52 t
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 SG B water level dropped below 50% (operating range), and the EFW 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 A 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 llPI after the RCPs were restarted caused the primary-system pressure to decrease to the minimum value for the transient (~3.5 MPa). Af ter the steam generators were isolated at 600 s, the primary system repressurized to the PORV setpoint (16.9 MPa) at ~4678 s. The PORV then cycled for the 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 f completely emptied by ~40 s because of liquid contraction resulting from the severe overcooling of the primary system. Af ter 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-system liquid expanded because the fluid heated up and the pressurizer slowly refilled.
Downcomer liquid temperatures for the base case are at the top axial downcomer level just below the cold-leg inlet nozzles (Figs. 34 and 35).
Because of the nevere overcooling in the affected loop (loop A), asynnetrical liquid temperatures are calculated for the vessel downcomer. The fluid temperatures in the downcomer cells associated with the loop-A cold legs were calculated to be ~20 K colder than the cells on the loop-B side (Fig. 34). The minimum downcomer fluid temperature calculated was ~405 K at ~526 s when the RCPs were restarted. While the RCPs were tripped off, the downcomer fluid
$3
es , , , , , , , ,
.m me. -
.m a... . see I go - .
ene T g
me lse<. .
mos n-no 1 m. & -
. . sos as , , , , , , . ,
e me suo see me ese see no ese ese M (a)
Fig. 30.
Pressurizer pressure (0-900 s)--base case.
es , , , , .
sees
- -"cs nos i see - .
1 . . . yee j
i ase I. es. .
see t
seo ee. .
no
- m. .
. see i
se , , . . , ,
e ese asse sees mee sees esse vues esse MW Fig. 31.
Pressurizer pressure (0-7200 s)--base case.
l l 54
s . . . . . . . .
e- -
s 4 .
1 #
E s<' -
l , .g l
t< .
- e. A / . .e
-t . . . . . . . .
e me see see see see see me ses ese MW Fig. 32.
Pressurizer water level (0-900 s)--base case.
u . , , , , ,
a- - -ee
- n. .
3.
I ..
E e- -
- se 4< .
3< .
e- -
e
-t e mes sees seine seise se'ee ee'ee mise sees MW Fig. 33.
Pressurizer water level (0-7200 s)--base case.
55
eee . . . . . . . .
e,..
R TH Z .
, 216
.. 226 .
{ 236 S see. 246 .
=256 266 -
l Isee. so-es- -
es- -
me a -
I spe . . . . . . . .
e es see see me ese ese me ese oss MW Fig. 34.
Downcomer liquid temperatures (0-900 s) at vessel axial level 6 (all azimuthal sectors)--base case.
l m , , , , , , ,
J j see; ^~i35 see. .
E m. RNZ .
216 I m. 226 -
I l
es-236 246
=256 5 m- 244 -
j l m- -
- m. \ -
i m . . .
l e use asse sees mee sees esse mee eues l
MW Fig. 35.
Downcomer liquid temperatures (0-7200 s) at vessel axial level 6 (all azimuthal sectors)--base case.
1 l
l 56
temperatures were af fected 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. [
The hot-leg liquid subcooling in each hot les is shown in Figs. 37 and 38.
i Figures 39 and 40 show the hot-leg mass flows and Fig. 41 shows the candy-cane void fractions. The subcooling margin calculated in loop A is significantly j larger than in loop B for auch of the transient because of the MSLB in loop A i
- and the resulting enhanced heat transfer in the af fected steam generator. The ,
4 i
?
enhanced heat transfer caused higher natural circulation flows in loop A i compared with loop B (Fiss. 39 and 40) from ~150 s to the time the RCPs (Al and I f B1) were restarted (~526 s). The 42 K subcooling margin was reached in loop A I at ~275 s, but the RCPs could not be restarted until this margin was reached in i
) all loops. The flow in loop 5 stagnated at ~150 s because the candy cane in f this loop reached saturation and voided (Fig. 41). The subcooling margin in I
loop B was not reached until ~526 s, at which time the RCPs were restarted.
i j
Af ter the RCPs were restarted, the void in the loop-B candy cane condensed and ,
i was swept out (Fig. 41), and the subcooling margin in both loops equalised. It
) should be noted that, when 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 rematader of the transient; j thus the HPI was never turned back on nor were the RCPs tripped again. !
l As discussed in the preceding paragraph, the loop-B candy cane voided at j ~150 s (Fig. 41) and the void was swept out after the RCPs were restarted. Even i
i though the candy cane in loop B voided, which is the highest point in the i
primary system, the vessel did not void, as shown in Fig. 42. Figure 42 shows t
i the volume fraction of the upper plenum liquid, which is the region above the ,
j 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 f legs. Figures 43 and '45 show the cold-les mass flows for loop A and loop 5, I i 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 affected steam generator. The loop-B flows before j RCP restart were essentially stagnant, as shown in Fig. 45, because the loop-5 candy cane voided. When the RCPs (A1, 81) were restarted at ~526 s, essentially i
57 l
I
88 . , , , , , , ,
I me. .
me. .
i 1
me. .
i me. .
g me. .
g .. . ,
.e. .
ee. .
- o. ~
-ee e see aio aie see ese ses s ese see MW Fig. 36.
Total vent-valve flow into downcomer--base case.
m . , , , , . , ,
ee LOOP A 00LD) ,
toor a paso es.
. me I
m es.
\./,,.***,,,............... me
- m. -
, .i . se
/*
es
- a. .
\.
1 .,
,.**....=
, ,3e
{ y,.**** ' , . . . . ~
t.- le
- e. - . e
-ee
-m . . . . . . . .
o us see see me see ese no ese see i MW Fig. 37.
Hot-leg liquid subcooling (0-900 s)-base case.
i l
\
l l
58
as , , , ,
. . - Loop A , 3,,
J
Loop B so- . .
.ps l
l N- -
.se g .
j
. ~ . . . . . . . . ..
[ as Igg. m-m.
l se ll m-
! 7s T !
f se ae- .
,: se l
(f
.. . - e
-n , . .
eer i e noe snee snee neee seee enee seee MW Fig. 38.
Hot-leg liquid subcooling (0-7200 s)--base case.
unse , , , , , , , .
- Loop A asses m, ----- Loop B ,
seeme enee- \ -
\ mean t
I enes-
- "M m .se . - \. p.... .
g sees
{
sees- .
\g -
.e e
'.5
.. . s, ................................. . ..
\
-sese
-*ese ,
e as see see see see ese see eso see MW Fig. 39.
Ilot-leg mass flows (0-900 s)--base case.
1 1
59
l sese , , , ,
. Loop A ames
., ----- Loop B ,
1 l sneee l l 1
eene- .
j i seen l
! enes. .
g
_ e. . .
.seen g
sees- , .
..q..: . .e
-asec , , .
e eso anos anoe esso seso seso me sees MW Fig. 40.
Hot-leg mass tiows (0-7200 s)--base case.
u , , , , , , ,
- s. .......................... Loop A ,
! ! ----- Loop B sa- I .
lu-I t .
s4- .
I sa< g i .
I e
=4.1 . , , , . . . .
e me see see me ese eso suo ese see MW Fig. 41.
Candy-cano void f ractions--base case.
60 l
u , , , , , , , ,
l 1 -
l u. .
l u. .
i u. .
l l
- u. .
e- -
-4.1 , . , , , . . .
e ne see see see see one see ene one Min) i Fig. 42.
, Upper plenum liquid volume fraction--base case.
sees ,
enes. LOOP Al -
Loop A2 ,,,,,
enee- -
asse. . . sees Neo- -
4 , . sees &
Ma -
. asse mee. -
- e. .
- e
$ .s, Io i
.g,g . n, l'-
F l ,
..seee
,,,,,, .. ..... .....l ) !..! L-.I .
- . esse
-seec , . . . . . - .
e es see see aos ese see see aos ese Mim)
Fig. 43.
Loop-A cold-les maea flows--base cane.
61
srs . . . . . .
ses Loop Al ase- ----- Loop A2 -
sus- .
, age
- E
. . . se l.se.a. .
. nee ase g _. .
g
. t.. ase
- g. -
i . No
- m. -
i . ane srs . . . . . . . .
a se ano ano aso neo eso see see see MW Fig. 44.
Loop-A cold-leg liquid temperatures--base case.
nec . . . . . . .
Loop B1 7 ,
-- --- Loop B2
.asse esse-asse. -
. sees e asse<
eees I
H- ~
.aome see. .
,, ..~.................= , ,,
j i - i l ,'
.\ ..so .
I. ! ,
,g,
- i. ...... .f .,)
g l
4 .J i..l
.. esse esset . .
e ao neo ase ano see eso veo see ese MW Fig. 45.
Loop-B cold-leg mase flows--base case.
62
steady-state flows were calculated in cold legs Al and B1, but the flows in cold i legs A2 and B2 reversed (flowed out from the vessel), as shown in Figs. 43 and
- 45. For the remainder of the transient, the cold-leg flow directions and i
magnitudes remained essentially the same af ter ~526 s, as shown in Figs. 43 and
- 45. The cold-leg fluid temperatures are shown in Figs. 44 and 46 for loops A l and B, respectively. The loop-A fluid temperatures were significantly colder than those of loop B during the time the RCPs were tripped (~50-526 s).
4 Detailed results of key system parameters in the steam generator are shown in Figs. 47 through 54. Figures 47 through 50 show details for the affected steam generator (SG A), and Figs. 51 through 54 show details 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 sido 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 of the broken steam line during the course of the transient. After the secondary side of SG A had depressurized sufficiently ( in ~ 100 s) , the EFW penetration increased, and the water inventory began increasing (Fig. 47) as less emergency feedwater was bypassed out the broken steam line.
Figure 50 presents some detailed plots to help explain the EFW penetration phenomenon. The top plot in Fig. 50 compares the TRAC-calculated vapor velocity at the EFW injection point with the complete flooding curve predicted by the Wallis-Kutateladze correlation (K = 3.2) for various pressures in the SG A secondary side. This plot shows that EFW penetration will not occur untti the vapor velocity is less than ~8 m/s, and this velocity is not reached in the TRAC calculation until the secondary side depressurizes to ~0.5 MPa. The bottom 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 e and did not change significantly until ~350 s. As will be discussed later, the reason for the increasing inventory af ter ~350 s was because the emergency feedwater to ;
SG B was terminated and all emergency feedwater was directed to SG A. The inventory decreased following RCP restart at ~526 a because of the enhanced heat transfer from forced convection on the primary side. Then, at 600 s, the steam l
l 63 1
see . . . . . . .
see m' Loop B1
~
Loop B2 ses g tes. "
{
. =
l e..
ne. .
~ '**e ee Iage.
.e I
ass.
I x .....,,** . ,,'...j "
ee. ~
.see "o k k k k k O S m m MW Fig. 46.
Loop-B cold-leg liquid temperaturee--base case.
wooo . . > > >
1 8' "
- m. -
neses goes. "
seoes 2 . sos. "
mese I
es.o. -
..see
..e see .. "
-ese
- e. '
-= ; ,, .se e eso .se see .se =
.= dees i
MW Fig. 47.
t SG A secondary-side water inventory--base case.
64
To , , , , , , , ,
en
- e. ,
se.
28 a- .
'ene l
e- -
. me i
.I o< - . ..
-m . . . . . . . .
e se see ano me ese see noe ese eso MW Fig. 48.
SG A secondary-side pressure--base case.
use , , , , ,
mee . . see see .
. . sm i
see. o . asse see. . . .e i see. . . .m I
. . . . sos 1
see< - . see e< - e l
- We . . . , , . . .
e me see see aa ese ese see ese ese MW lig. 49.
SG A steam-11ao flow--hane case.
65
m-3=
me %ser i
w
a k i i is k k
- e. wm (mm) j ..
I:
- s. ,
/
- 5 v.p.ck.wnf ten /s)U A Fig. 50.
Counter current flow limiting phenomenon in affected steam generator (SG A)--base case.
essee , , ,
l , ,
noses ease. .
l essee. .
seeme sesse- .
2 asese- .
E
,essee g
. . sos sees < .
so m seen-sees sees . .
e se see see see see ese see ese see MW Fig. 51.
SG B secondary-side water inventory (0-900 s)--base case.
66 i
generators were isolated and emergency feedwater was terminated. The water inventory in SG A then decreased to zero and remained empty for the remainder of the transient.
I Figures 51 through 54 show some key parameter plots for SG B. Figures 51
[
l and 52 show the secondary-side water inventory calculated for SG B. At ~ 48 s ,
the low-level limit was reached following TBV operation in SG B, and emergency feedwater was initiated. The water level continued to inc rease to the 50%
operating range, at which time the EFW valve to loop B was shut (~350 s) . The SG B level remained essentially constant at the 50% level until the RCPs were restarted at ~526 s. After the RCPs were restarted, the SG 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). The level rose until the steam generators were isolated at 600 s. The secondary-side liquid 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 this time (no emergency feedwater to SG A after 600 s). At ~6121 s the SG B secondary-side level decreased to 50% of the operating range, and emergency feedwater was initiated. For the remainder of the transient, the emergency feedwater was controlled to maintain the 50%
operating-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 loo p-B steam line. The pressure then began to decreane af ter emergency feedwater was initiated, following the low-level limit trip at ~50 s. The pressure continued to decrease because of condensation from the cold emergency feedwater until the 50% operating-range level was reached at
~350 s and emergency feedwater was terminated. The pressure remained essentially constant at ~3.5 MPa until the RCPs were restarted at ~526 s. After l this time, as discussed previously, condensation of the steam in the secondary side on the SG tubes because of secondary-to primary heat transfer caused the pressure to decrease further to ~1.5 MPa. The pressure remained at ~1.5 MPa untti ~900 s. Af ter SG B was restored at 900 s, the secondary-side pressure increased slowly (Fig. 54) until the TBV setpoint was reached at ~5462 s. The 67 i
essee , , , . , , ,
. . masse
.sosse sa. .
2 ,/ 2 g
I.see. . sesse 3 g_ .
.-o g
. sos. '
seees mese.
. .mese sees . .
e see ases ases ases sees sees peos esse M (4 Fig. 52.
SG B secondary-side water inventory (0-7200 s)--base case.
Se . . . > > > .
. .ges es. .
M- -
Wee ee. .
I go.
- m. .
.i go. .
. .see i
go. .
.ges e- .
e , , . . . . . . e e me see see ano see eso see ese ese Mie)
Fig. 53.
SG B secondary-side pressure (0-900 s)--
base case.
68
i 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 temperatures are shown in Fig. 56. The MFW pump tripped on low-suction pressure at ~48 s. The loop-A MFW liquid temperature (Fig. 56) essentially followed the
, saturation temperature corresponding to the SG A secondary-side pressure; thus this temperature was much cooler than for SG B. Figure 57 shows the MFW mass flows into each SG. 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 steam generator are shown in Fig. 58. Emergency feedwater was started at ~30 s to SG A and at ~50 s to SG B. Again, because of the low secondary pressure in SG A, more EFW flow was directed to that steam generator. At ~350 s, the EFW valve to SG B was shut (50% level in SG B reached), and all of the emergency feedwater was directed to SG A. At 600 s, the steam generators were isolated and the EFW pump was tripped off. Figure 59 shows that, at ~6121 s, emergency feedwater was initiated again to SG B af ter the level had dropped below 50% of the operating range. The EFW temperature at the pump discharge is shown in Fig. 60, and the EFW 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 m 3 of accumulator liquid were discharged into the primary system. Figure 66 shows the PORV 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 case (Case 1) except the EFW 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 l 69
88 , , , , , , ,
- mes y,,,.-
M-l
.see i a- .
I .. - .
1 w
l.. - .
=
- . "#8 as-
-soo
- m. - .
8 - . . . . . . .
-#o e see aseo anos moo soec esso woo esoo M (4 Fig. 54.
SG B secondary-side pressure (0-7200 s)--
base case.
ese , , , , ,
- . sono see- .
~ ' #88 ag.
I I
" ' 3**
see- .
I .s.
. ..e I me- - . wee e . . , , , , e e se as se se me se me MW t.
Fig. 55.
MFW pump speed--base case.
70
a a:I- , , , , , , -es
.ao ese- Loop A ~
. ----- Loop B as des- .
. E
.m me- -
i 1
seo l i
1 ans. -
i ass as-
~
aeo
.m as , , , , , ,
9 se e to ao to 30 too MW Fig. 56.
MW liquid temperature-base case.
Sese , , , , , ,
see-
'j Loop A - anoo
, J ---- Lo0P B _
nn- .
soo one- .
30o 300- : -
i moo me- I .
i.
.i 7eo see- : -
.: .See see- : -
30- *i . see A_..............: ,
3- .
-o
-no . , , , , ,
e ao se se as no se ses !
MN -l Fig. 57.
MW mass flows--base case. I 1
l 71 1
~- - - .
tes . . . . . . . .
~
~
Loop A '"'
so-
=- Loop B .
ase so- .
.soo l
es. . :p
.wo g se- .
- .. .... ~i:: no l
.=*..~.. l
- se l
- i :
I :
m-. t . */ i -
g- J k. !......................I o I
-m . . . . . . . .
e no aso aso ao soo soo no soo ese M (4 Fig. 58.
EW mass flows (0-900 s)--base case.
4 we
- L Loop A .seo es- --- Loop B -
aso no- .
.soo
{
4 ee- .
Q
. me 8
- 8 so-
.'l.: :.i
.. no m.
.: ::::[:::: -
t:
.:: -: :: so
- i :: :: :
as- ,i E:i: ! :: -
.m
- ::i:::e*::j; si e- . . .
. .o
-se '
- o men aseo sooo moo sooo sooo woo sooo MW Fig. 59.
E W mass flows (0-7200 s)--base case.
72
seu , , , , , . . . i
- asso seu. _.
. so.see sens. -
E E seu- ,
,ensu I 304 3-a m.
se.rse 3 eat-
. ensae see. J -
ses.e , . . . ease m
e no ano ano ano soo ese eso eco MW Fig. 60.
EW liquid temperature at pump discharge--base case.
eso , , , , . . , ,
. soo 1,00p A see- ---- Loop B .
... \ soo ses- : -
E : e L
i; j . ... ... ......
aee en-I .
i j aos age. :. .
I i 3
- soo ase- i -
'.......................................=* me amo m
e me ano ano ago see eso ese ese MW Fig. 61.
EW liquid temperatures at injection locations--base case.
73
ms , , , , , , , ,
so- N, .
1 wa-
-m I e- Loop Al -
( . Loop A2 -ao g m- -
f
- 3. .
.so g
u- -
S' ~
.a u- .
e- -
o
-u . , , , , , , .
e so ano soo ao soo soc no soo soo M (s)
Fig. 62.
Loop-A HPI flows--base case.
es , , , , , , , ,
.as
- s. %
-se a.s - -
Loop B1 ,,
Loop B2
- e. .
So 12- ~
c .
.s c s-, -
u-. -
s e-- o
-u , , , , . , , .
--s e so soo ano ano soo eso no soo soo M (s)
Fig. 63.
Loop-B HPI flows--base case.
74
asse , , , , , , , , meet Loop A I ases- ~
Loop B l . nnen asse- -
$ M- .
s annetu ,
m- -
N . - sees 3
4 ans. .
I noe. -
. EJsPI as,e . -
Asso , , , , , , , , mJuss.
e so zoo ano me soo eso no eso soo MW Fig. 64.
- Accumulator water levels--base case.
4 y g y , , p 1 i
. m Loop A s- ----
Loop B :--------~~------
O 0 30 4 .
- s lu- as.
es-se as- -
i . ..
e
-e2 , , , , , ,
no e se ano ano me eso eso see see MW Fig. 65.
Accumulator liquid volume discharged--base case. l 1
i 1
75 l
l______-_-- - , __ .. -- -. . - - . -- - - - - - - - - - - ~ _ . - -
1 that the EFW system did not actuate). At ~400 s, the RCPs were restarted af ter l 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 I 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 I
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 steam generators were isolated at 600 s, the pressurizer water level began to increase again because of fluid expansion.
Figure 69 shows the downcomer fluid temperatures near the cold-leg I connections in the vessel. The minimum downcomer fluid temperature calculated was ~475 K and occurred when the RCPs were restarted at ~400 s. Asymmetrical fluid temperatures in the downcomer were calculated in Case 2 in a manner TABLE XVII I,
MSLB" (CASE 2) SEQUENCE OF EVENTS i
EVENT TIME (s) 1-14. Approximately same as base case 0-53.9
- 16. SGs A and B isolated 600.0
- 17. Calculation terminated 900.0 i
- EFW pump never started.
76
e , , , , , , .
n- -
8- .
3 3.- -
m- -
I
.. . .e 5 .
5 l -
.. l
. l' s s- g' e b8 bb -
- d. A m m' a TIME W Fig. 66.
PORV mass flow--base case.
me . . .
-2200 M- "
2000 no.
) '
._ i g
n00 $
" ~
g l .0
. 0.
t 1 . . . . . ,
-800 de . 800 200 300 400 H0 000 700 .00 .00 T.E (*)
Fig. 67.
Pressurizer pressure-Case 2.
l 77
s- -
-s 4 -
2 --a E
s.
h 2- -
E 1 8 5 s
g._ --0
~' o no ao no 4co soo eso s soo soo M (5)
Fig. 68.
Pressurizer water level--Case 2.
sao , , .
s,1 R TH Z uo 216 -
- 226
' .236 m E 90- ) 246 -
=256 **
gno -
266
-.m e
-a
- g. .
- et!
m o O & k 5 S & Y $ m M (s)
Fig. 69.
Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors)--
Case 2.
78
l l
similar to that of the base case, with the colder temperatures calculated for the loop-A side of the downcomer.
Figure 70 shows the hot-leg mass 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 much higher than in loop B because of the enhanced heat transfer through the affected steam generator. At ~400 s, the RCPs (Al, B1) were restarted and forced
! circulation through the primary system was calculated. Figure 71 shows that the l 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 after the steam generators were isolated at ~600 s (Fig. 71).
The cold-leg loop flows for loops A and B are shown in Figs. 72 and 73, respectively. These flows were similar to those calculated for the base case, except that the loo p-A flows were somewhat higher in the base case. The cold-leg 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 minimum temperature in loop B was ~400 K. In Case 2, the loop-B cold-leg temperatures were colder than for loop A; however, this was not true in the base case. This is because the EFW system was operational in the base case and caused the loop-A cold-leg temperatures to be less than the loop-B temperatures.
Details of key SG parameters are shown in Figs. 76 through 80. Figure 76 shows the secondary-side water inventory for SG A. Because the EFW system did not work and because the MFW pump tripped at ~50 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 ~85 s, SG A had depressurized to essentially atmospheric pressure. The steam-line flow out of the broken steam line is shown in Fig. 78. Figure 79 shows the SG B secondary-side water inventory and Fig. 80 shows the secondary-side pressure in SG 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 after the RCPs were restarted at ~400 s because the heat-transfer rate from the secondary to the primary increased.
Other system parameters are shown in ' Figs. 81 and 82. The MFW pump was tripped at ~50 s on low-suction pressure, and the feedwater flow decayed to zero 79
aseo , , , , , , ,
Loop A ameo
. - -- Loop B .
-seaso eneo- \
\ eeno I
seso- 1 i
j - Seen
?
.. i r.. - - _ -
} -eseo i
\
seso- i\ -.
\. . I
- g. '. ,e_ . ...I ..g
\.:
~~***
= Boo . , , , . .
a no zoo soo ao soo eso no eco soo M (s)
Fig. 70.
Hot-leg mass flows-Case 2.
u , , , , , ,
Loop A
, t- ,,. ----- Loop B .
u- ! i .
i u- l -
I u- ,
)
u- l -
I on - -i
=&2 , , , , , , ,
o no soo ano ao soo eso x, eso see M (s)
Fig. 71.
Candy-cane void fractions--Case 2.
l l
f i
l 80
l l
me , , , , , ,
.m seco- '- -
.neoo sooo- -
.. Loop Al ..seoo asso- -
, , f f
- m. -
..ooo h
y m- -
- e. ,- o I $
l l : .:
- i : ; i = i L ..ooo
.acoo-rL! L..!/ .!!: ._..:L ' ::" -
..tooo
.Jooo . , . . . - -
o no ano aoo 4eo soo eso m soo soo M (s)
Fig. 72.
Loop-A cold-leg mass flows-Case 2.
mo . , , , ,
.m sooo- _
+-
.noco sooo- -
4000-Loop B1 -~"
T ----- Loop B2 g g sooo- -
,,,,, g h
e m-
.sooo f
m j m- -
g
.. . --o
- l.
g l
-moo- ii ! g E , i !!- --8000
.:i .. r.. .ii.s.: i..si i ....!! . :~t.
- i
s ..
.sooo- -
-aeoo
--enes e so soo aeo soo soo eso m soo soo M (s)
Fig. 73.
Loop-B cold-leg mass flows-Case 2.
81
m , , , , , , ,
seo eso- Loop Al .
Loop A2 -seo sno- .
j -sso g "~ ~
E
~~
5 no-' 4 I gg.
sto-
"E
--aeo
..o I 400- -
So , ; , , , , , ,
o no ano aoo .co eco eco no soo soo w (s)
Fig. 74.
Loop-A cold-leg liquid temperatures--Case 2.
eoo , , , , ,
Loop B1 ~.ses
"~ , - Loop B2 sSo-
-~
m." .
E W
{< m. .~** 2 9 , , .- '.,
me- ,
y~ .'"
~"
i m , , , . . , ,
o no ano ano 400 soo soo no soo soo M (s)
Fig. 75.
Loop-B cold-leg liquid temperatures--Case 2.
82
wooo , , , . . . . .
-38000 9000- ~
-3e000
-38809
} enco-
., 1 N
sooo- -
1 moo.
-.eooo 3000 ~
4000 0- 0
--e en 0 WO 200 300 M too 800 M0 M M M 4)
Fig. 76.
SG A secondary-side water inventory-Case 2.
n , , , , , , , ,
gg.. -ooO t
so. -
no ao- -'**
i
,, .-300
,_ - -90 q, -
O
~*
e no aim ano ao sie sio me eso ese M 4)
Fig. 77.
SG A secondary-side pressure-Case 2.
83
see , , , , , ,
anse 900-aseo 350-asse
$ moo-T K sees 4 M'
300 h l 3 se0- -_g 2 e.
( .
..e.0
~~***
.aB0 O tes aic 350 e50 edo e40 M she see M (s)
Fig. 78.
SG A steam-line flow-Case 2.
assec , , , , ,
a3eoo- -
40e00 i
- g. a 6
- 2. . -assos a g_. .
g assee tete 0-see00
.gagge "e sie ano see eso se0 M (s) eio m see see Fig. 79.
SG B secondary-side water inventory-Case 2.
84
l so , , , , , , ,
Mao ;
- g. .
so- .
a so.
m m ooo
- n. -ngo E , , , , , , ,
-3eo e no soo no ao soo soo m eso soo M (s)
Fig. 80.
SC B secondary-side pressure-Case 2.
t
- o , . , ,
ego.
i aoen 1 ,eo. .
._ 1 i
i aso.- .-seso i
go.- .- sete i 3" o eet . . . .
e no me ano ao eso soo m eso see w (e)
Fig. 81.
MFW pump speed--Case 2.
85
, - _ _ .. . _ _ ._ _ - _ _ _ _ _ _ _ _ - _ _ _ ~ .. _ .. .. _ . _ . _ . . - __ __ ._
l-l
! by ~150 s. Figure 82 shows the HPI flows into each cold leg. At ~400 s, the
] HPI was throttled after adequate subcooling was reached in the loops.
- c. Parametric Case (Case 3). Parametric Case 3 was identical to Case 2 except that the RCPs did not restart as intended because of input errors.
Another significant difference between Case 3 and Case 2 is that the MFW pump
)
i did not trip until ~330 s in Case 3 because of errors in the ICS modeling.
j Because the MFW pump did not trip until late in the transient, this calculation l can be considered as a MSLB with runaway main feedwater. The sequence of events calculated for Case 3 is given in Table XV111. The events that occurred during the first 54 s were approximately 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 c
E TABLE XVIll j MSLB (CASE 4) SEQUENCE OF EVENTS Event Time (s) i MSLB-loop A steam line 0.0
- EFW to both steam generators 11.5 4
j HPI initiated 22.4 RCPs trip $2.4 4
Loop-B candy cane 140 2 voided HPI throttled 302 Emergency feedwater terminated 311 at SG B i
SG A isolated 602
! SG B restored 900 Calculation terminated 2100
!. -86 I-l L . . - . .- . _ _... _. . _ _ . _ _ _ _ _ _ . _ _ . _ _ _ . . _ _ - . _ _ _
~330 s, the MFW pump was tripped because of a high water level in SG B. At
~356 s, the HPI was throttled af ter the 42 K subcooling margin was reached. The RCPs failed to restart at this time. At ~465 s, the HPI was turned back on again because the 42 K subcooling margin was lost. At 600 s, the steam generators were isolated, and at ~693 s, the HPI was again throttled. The calculation was terminated at 1260 s.
The primary-system pressure is shown in Fig. 83 and the pressurizer water level is shown in Fig. 84. The minimum pressure calculated was ~5.5 MPa and occurred at ~125 s. Af ter natural circulation flows were established (~150 s),
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 s because the HPI was turned on at ~465 s and throttled again at ~693 s (Table XVIII). The primary system repressurized more than in Case 2 because the RCPs were not restarted, and natural circulation flows existed for most of the transient. The pressurizer completely emptied by ~30 s and began to refill at ~175 s after the primary system began to repressurize. The pressurizer water level decrease at ~350 s and subsequent increase at ~460 s can be attributed to HP1 throttling (Table XVIII). The change in slope in pressurizer water level at ~700 s can also be attributed to HPI throttling.
Figure 85 shows the downcomer fluid temperatures for Case 3. The calculated minimum downcomer fluid temperature was ~450 K and occurred at
~650 s. The downcomer temperatures for Case 3 were colder than for Case 2 because of the runaway MFW flow. Asymmetrical temperatures were also calculated similar to the base case. The downcomer fluid temperatures increased after
~650 s because the steam generators were isolated at 600 s.
Figure 86 shows the hot-leg mass flows for 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 in loop A was higher than in loop B. Natural circulation flows continued for the remainder of the transient because the RCPs never restarted. Figure 87 shows that the candy canes in both loops never voided. The candy canes in Case 3 never voided because the main feedwater provided enough cooling to the loo p-B steam generator to prevent any secondary-to primary heat transfer, which occurred in the base case and Case 2.
87
I i
i 1
m4 . , , , , , , ,
i b(sold) LOOP At(dmal$
si(shnee9 LDop at LDOP A2 93- -n s- -
se .
.3 -
g k
m e. -
h 74- -
s-
-e
- u. -
... e e s 6 a 4 a 4 k a m M (s)
Fig. 82.
HPI flows--Case 2.
no . . .
asse I- --
I es. .
mee
. .e.
- g. .
g
, so- -
see 40 . . . . . .
e ano ne see ese see see sees MW Fig. 83.
Pressurizer pressure-Case 3.
! 88 l
7 . . . -
. se s- -
s f a. .
g h .. . . l l s. .
1 s
i.
e- -
e
~' s soo eso see eso neo uso mee Fig. 84.
Pressurizer water level--Case 3.
soo ,
p RNZ sn' g 216 226 .w I .236 g W- 246 -
se g a256 h ,,.. 266 . m m
lsee. m- -
.a 3sl 40s. .
am 4eo . . . .
e Me 400 ese ese 100e use tage 1st M Fig. 85. !
Downcomer liquid temperatures at vessel axial level 6 (azimuthal sectors)-Case !
- 3. )
89
u , , . . .
Loop A i- -----
Loop B -
sa. .
SA<
e,a - -
an. .
e -
T y y y V y e see aos ese ese use use esse Tat 60 Fig. 86.
Hot-leg mass flows-Case 3.
ame , , ,
Loop A assee mese. ---
Loop B .
seee. i 1
- sees esee- .
..e.e y
y ease- .
- eee, Neo- ..m e- '--~~~. ,,_,,,, Q -
-.o
-seen . . . . . .
-eee e see me see see see see e Fig. 87.
Candy-cane void f ractions-Case 3.
90
1 l
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 l legs had natural circulation flows up until the time the steam generators were isolated. These flows then decayed to approximately zero because of significant j loss of heat transf er _ 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 for Case 2 before 600 s because of the runaway 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 runaway MFW filled SG B to the 90% operating range, and then the MFW pump was tripped (Fig. 97). The 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 main feedwater until the MFW pump was tripped at ~330 s. After SG B was Laolated 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 provided by the Duke Power Co. resulced in several clarifications in the location and operation of various instruments.
This information resulted in the following changes to the model:
e The MFW pumps automatically trip at 0.5 s instead of at 47 s (based on the RELAP-5 calculation in Re f . 5 ) . This is because of uncertainties in the measurement of the SG liquid level during transient conditions (6F vs collapsed level).
- The HPI throttling is based on subcooling at the exit of the core instead of at the hot leg when the RCPs are not operating. HP1 throttling is based on subcooling in the hot leg only when the RCPs are operating.
91
i sees , , , , , ,
Loop A1 -me esse- -----
Loop A2 -
. sees asse- .
g
.seen y
M asse- . m g mee g
use- ..,,,, !
e_ .
.e
.mee . . . , ,
-ases e see me ese ese see ese mee Fig. 88.
Loop-A cold-leg mass flows--Case 3.
J p I g I ,
j Loop B1 '"*"
Loop B2 see.
. asse- .
eeen y m- .
.meo see. ..,see
- 19 0s , ,
'*8888 e see me see ese see ese wee Fig. 89.
Loop-B cold-leg mass flows-Case 3.
92 l
l I
l l
_ _ _ . _ _ __ _ . . - - . . , , - - -~. ,
eso , , , , , ,
Loop B1 seo eso- ----- Loop B2 I -een see. -
- g. . me eeo.
h -
see l\
l 1\
s s
\.
f .see an. -
aae ao . . , . . .
e ano ao see eso eso neo wee Fig. 90.
Loop-A cold-leg liquid temperatures--Case 3.
eoo , , ,
~"
- ~
Loop Al '
Loop A2 .,,
-uo uo-. -
se eso -- - ***
ses- f'. ,- -
- I ,\ r'.) ee
- .. \ / -
) see
/
~
.see ano . . .
e see ese see see see see ines Fig. 91.
Loop-B cold-leg liquid temperatures--Case 3.
93
useo . . . . . . l we . . seen meno. . -sees Weso. . smo8 2 a sooo- - anos sooo- . unoo
- ..e.oo i asoo- ..
e-- ..o
-sooo , , , , ,
o so so no ano ano soa ano ao Fig. 92.
SG A secondary-side water inventory-Case 3.
M i . . . , ,
a*
so. .
me so. .
soo I. so-
. as I
E so.i . soo e- . .
e- . .o e son eo see see moo eso use Fig. 93.
SG A secondary-side pressure-Case 3.
94
l l
uso , , , , , ,
l
. o0 300- -
moo woo- -
-Noe Q one. .
?
E J -noo d
. co me- I
-.ooo l
aso-_ , -- co e_
( ..o
-ano , . . . . .
e ano ao ooo con moo noo woo Fig. 94.
SG A steam-line flow--Case 3.
escoe , , , ,
asooo-sooon- -
moooo seco- -
2 moco- -
. soooo g 88 " o-
-.m l_- .
. ._o g
Neco- -
so m .- .aooo i
scoo- -
-soooo Wooo . . . . .
o ano me eso soo moo soo mee Fig. 95.
SG B secondary-side water inventory--Case 3.
95
so . . , , ,
e 3 --neo n- -
as- -
co- -
so- - -
- I gg..
.mo
- g. -
soo a- -
J ss. . -.
ao . ,
e ano ao soo soo moo soo woo Fig. 96.
SG B secondary-side pressure-Case 3.
soo , , , , , .
L . so me-i -
1
- 3. -'"'
g.. ..
g I.. .._ 1 e.. ...
-Bo , . . . . .
e ase me see ese moo use wee Fig. 97.
MW pump speed-Case 3.
96
- The correct instrument location for measuring subcooling for RCP restart is located 3.0 m below the top of the candy-cane centerline instead of at the horizontal part of the hot leg.
These changes did not dramatically affect the minimum downconer liquid temperature; it reached a minimum of ~420 K, as opposed to ~405 K in the base case.
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 steam line. Both steam generators blew down momentarily, but the TSV quickly isolated the loop-B generator. At the same time, the reactor tripped, the MFW pumps were
) TABLE XIX 3 MSLBa (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
- 17. HP1 thrcttled after 42 K subcooling 356.0 reached; RCPs (A1, B1) fail to restart
- 18. HP1 on (42 K subcooling margin lost) 465.0
- 19. SGs A and B isolated 600.0
- 20. HP1 off 693.0
- 21. Calculation terminated 1260.0 aEFW pump never started;- RCPs never 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 similar to the base case.
97
5 2
tripped, and the feedwater-heater drain tank flow was ramped to zero. Closure
! of the TSV pressurized I the loop-B steam generator and the TBV began cycling between open and closed to relieve the pressure. This lasted only ~40 s because
- the EFW flow into the steam space at the top of the generator lowered the pressure in SG B. *1: 3 EFW flow began at ~11.5 s because of the low liquid level in SG A.
i The steam-line break initially caused rapid overcooling and depressuri-zation .of the primary side.. At ~22.4 s, low primary-system pressure actuated the HP1 system and,.30 s later, the RCPs tripped. The candy cane in loop B voided after a 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 HP1
- throttling; conditions were met at ~302 s. Adequate subcooling for the restart ;
of the 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 operatinr, range and the EFW flow was j
rerouted to SG A.
At 600 s, all feedwater 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 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 and 99. Differences in the i
system . pressure cannot be seen until ~302 s, when the HPI was throttled in Case 4; the system 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; thus little energy was deposited from the secondary side. In the base case, restart L
of the RCPs leads to significant secondary-to primary heat transfer and a faster repressurization rate. The downcomer . liquid temperatures differ slightly for
!- the two cases; the warming effect of terminating ilP L flow ~225 s earlier in Case 4 is compensated for by the cooling effect of not' restarting the RCPs.
. . Plots illustrating key events on the secondary side are shown in Figs. 100-107. -The break flow and SG A pressure history (Figs.100 and 101) -
I indicate' the rapid blowdown of SG'A; by 80 s, the generator had almost 1
98 i
[
l
. . , . . . . ~ _ y., _ . - . - . _ . - , , , . , ,,.m, , ._,---.,,,m, _m - .
we . . . . - - -
. esse tes- -
,g Gua4RDRM as- 6 dew 9 am CE . .ses
- I .. . .
I as- -
.mee N- ,,***""s', -
no
- m. ,..- .
see se . . . . . . .
e see see no see see see eso sees l MW Fig. 98.
System pressure for Case 4 and base case.
see . . . . . . .
.gge ses- -
.see see- ! W punm -
g - MMM me
.ao see- ,,
1 p
- ,e**** as me. .
- p*,,,, see no. *. .
ase en- .
l -
. V se as. .: .
t ase me . . . . . . .
e ano soo no see see see see sees MN Fig. 99.
Downcomer liquid temperatures at vessel axial level 6 for Case 4 and base case.
99 I
soo . . . . . . . .
. .an.
I .
eso. .
seen eno-
. I 3_ - .
-no 3
ano.
. -ano
^^
e-- ~- e
. -.aae
-aeo . , , , , .
o ano seo no moo neo eso sno anco azio MW 4 Fig. 100.
Break flow for Case 4.
n . . . . . . .
gg. .~M go-' -
Me I
ooo
- a. .
I so- -- eBo an-
- soo E m-- ..no L_
o
-m . , , , , , .
o me aeo me neo sao use ses seen anno MW Fig. 101.
SG A secondary pressure for Case 4.
I 100 i
so . . . . . . .
3 M- "
mos so. -
I
-soo I
g m. --soo g
E --.
.,oo e
- a. -
. Sco
- a. -
m-. --soo "o ano ano no isoo nao moo veo meo ano M (s)
Fig. 102.
SG B secondary pressure for Case 4.
I w . . . , , . .
(sesc0 LDOP A -20e no- M toop a -
-neo me so-c Go Q
M- * -
- /j: -se l *- ,l
/'
-.oo
!3 R
y so- ,' --e o- ... ..o
=38 , , . . , , . ,
o ano soo no uso uno eso see anoo ano M (P)
Fig. 103.
EFW flow (SG A and SG B) for Case 4.
101 l
9 me . . . . . . . .
no goo. -
' ~
-Soo son. -
g
-aoo 5 )
f so- -
-aos f h me-
-ano so- -
-wo I --o
-so . . . . . . . .
--=
o no too no moo tuo too son mac ano MW Fig. 104.
SG A total feedwater flow for Case 4.
no . . . . . . . .
, -35 as- -.m
.. .-zas s
A n_ .
E h .
ion h
g w. .
g 3 . -oo g as-
- o. --o
-4
'*' o aio soo no sio mio sie see ao'so amo MW Fig. 105.
SG B total feedwater flow for Case 4.
l l
102
l seco . . . . , , , .
-ameo eseo. .
seen.
. eeen Mee-,,
--See0 M- -
-seso I l
ecos. - -esso h
seen.
.esco 300- -
seen see. .
g.- .-o
-e00 , , , , , , , ,
o no aeo no uso son eso see see0 aseo 4
M@
Fig. 106.
SG A tube-bundle-region mass inventory for Case 4.
ano00 . . . , . . . .
-neee ao00o- -
-60000 3e000- -
e .
scoso
$ 3 .0o . .
-40000 ,<
9000- -
3 esso E". "
-2000o gego. .
.gggg o , , , , , , , , e o me See no 300 soo see geo se00 23eo MW Fig. 107.
SG B tube-bundle-region mass inventory for Case 4.
103
depressurized to atmospheric pressure. After 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 pressurization, and the TBV cycled to relieve the pressure.
Emergency feedwater began at ~11.5 s and the cold emergency feedwater 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 the emergency feedwater to SG B was terminated, flow was rerouted to SG A. ORNL specified that the emergency feedwater be terminated at 600 s; the total feedwater delivered to the generators is shown in Figs. 104 and 105. Even with the MFW pumps tripped at 0.5 s, the main feedwater was still delivered by the CB l and hotwell pumps.
The mass ' inventory in the tube-bundle region is shown in Figs.106 and 107. Af ter SG A blew down, the feedwater flow equaled the break flow until the emergency feedwater to SG B was terminated. At this point, the mass in SG A increased until feedwater terminated at 600 s; the generator boiled dry at
~1875 s. In SG B, the emergency feedwater and main feedwater 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, 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 HP1 flow for the transient; the flow was decreased to zero when adequate subcooling was reached at the core exit.
Figures 115 and 116 show the total vent-valve flow and the average upper-plenum liquid temperature. The vent valves make a significant contribution to warming the downcomer liquid. 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 downcomer liquid temperature of
~420 K and indicates the system will repressurize similar to the base case. The changes made to the base-case calculation for Case 4 gave a more accurate i prediction of the postulated accident at Oconee-1 and still indicate that this 104 .
useo , , , , , , ,
ameo Weso-
! Loop A -sesse
. --- Loop B .
meno
- m. .
I 1 5 g-.; i
. sees g
- seso-.! -.
% A + _
o.
(r - _ -
.o V
-asco , . .
e ano eso no woo ano moo sea aseo sano MW j Fig. 108.
Hot-leg A and B mass flows for Case 4.
soo . . . . . . .
soo soo- -
~"
Loop A
- ~ ,i --- Loop B -
E 'l E teo- ,
}
see I.' {\
,,. %, . ese I
t see- g ' .~. s! ~~-
so ago. .
we
- ~ *
.sno 44o . . . .
e ano eso me neo ano neo see asas asse MW Fig. 109.
Hot-leg A and B liquid temperatures for Case 4.
105
seeo , , , , , , . .
-asso Se00-
- l
[ Loop Al -asse
,_ ----- Loop A2 .
.seco m..
5
-asso
.t
-=
0 ase 400 30 900 See 300 .s0 3e00 23e0 MM Fig. 110.
Loop-A cold-leg mass flows for Case 4.
m- , , , , , , ,
I
-seo soo- -
88-
! Loop Al
Loop A2 '-*oo E E sea.
g< ..
g f Q m.- . aos
.?'.
l =-
\, , / -
~"g
'"- s# '
.\ (M seo 5 j
(/ ano
. . .e. . --
MW Fig. 111.
Loop-A cold-leg liquid temperatures for Case 4.
106 l
l
\ __ _. . _ - - _ . - _ . ,
i 8888 . . . . . . . .
2000 5
4000- .
Loop B1 8880 l - Loop B2 i 3080-oooo 5-. .
-. ...e00 4- .ky.n._m -?_ e- -- --a
-moo , , , , , .
--anoo 0 ano Goo no m00 see see sea sono 2:e4 w(s) i Fig. 112.
Loop-B cold-leg mass flows for Case 4.
eso . . . . . . .
.a sys-, -
sec
.. Loop B1 .
Loop B2 mo g
us-. -.
800-- -
440
- c. -
f' , , ,
Ky, jf 7%) .
See 3:e
- m. .
. 33g eso . .
0 20 See me 200 Sec 800 see 3000 3500 MW Fig. 113.
Loop-B cold-leg liquid temperatures for Case 4.
107
_ -_m _ __. ._
eo , , , , , , , -yi M' *
-to Go- -
as to- -
-mo do- -
-M so. .
-no 5 ,
an- -
I
- m. . m 4
o- -
o
-m , , , , , , , ,
e ano ano no moo uso moo ano anoo ano ima:(s)
Fig. 114.
Total HPI flow for Case 4.
eo , , , , , ,
p 400- j. -
I .
me. .
aso- -
ato- ..
n t{i g zoo- fi -
~
\,p #W' .
g- f a
-se , , , , , ,
e no ano aso eso eso eso no eso see 7tdE(s)
Fig. 115.
Total mass flow through vent valves for Case 4.
i 108 t
transient could pose a threat of PTS to the reactor vessel. However, these changes had no significant impact on the overall conclusions stated for the base case, and the minimum downcomer fluid temperature remained approximately the same (~405 K for the base case and ~420 K for Case 4).
- 4. Conclusions. The overcooling of the primary side of the Oconee-1 plant caused by a full double-ended steam-line break in one of the steam lines 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 i protection and control systems operate as deeigned. The minimum downcomer fluid 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 MPa.
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 that prevented the RCPs from restarting once adequate subcooling was achieved.
ene . . . . . , , .
ass <
ane ans.
8 E
.e.
see.
- . me
. .es ano<
. .ae me. .
. aos aos ede . . . . . . . ,
e aos eso me see une see see sees see lhs W Fig. 116.
Upper plenum liqufd temperature for Case 4.
109
For these parametric cases, the downcomer fluid temperatures were considerably higher than for the base case (Case 2 at ~475 K and Case 3 at ~450 K); thus greater margins against PTS were calculated.
l In the last parametric case (Case 4) other changes were made to the model I 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 downcomer fluid temperature 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 presents the Oconee-1 plant response to a small primary-system break (the failure of the PORV in the full-open position). The PORV was assumed to open at transient initiation and remain 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 FORV f ailure, the ICS f ailed to run back the main feedwater. As a result, the steam generators continued to fill until the MFW pumps were tripped on a high SG level signal. The PORV and ICS failures were the only assumed system-related failures. The RCPs tripped 30 s af ter HPI initiation, and this was the only specified operator action.
TRAC calculated a minimum vessel downcomer liquid temperature of ~528 K between 600 and 700 s into the transient. The primary system was calculated to repressurize to ~11.5 MPa af ter 800 s.
- 2. Model Description and Assumptions. A complete description of the primary-side, secondary-side, and ICS modeling can be found in Sec.11. The steady-state operating conditions are also presented in that section.
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 presented in Table XX. To ensure a MFW pump trip on a high SG level signal, the low-suction and high-discharge pressure trips that could prematurely trip the MFW pump were overridden. Also, the MFW pump speed was increased 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 steam generators. The MFW pump maintained the maximum speed setting until the trip at a high SG level.
110
J, d,emA - - - . 2-. .- 43__ _,.eA +L,_,s,,.__ _L___J w___.:_._.4_ 4_A4 a ,a.. _.,3A 4 .i .,m._ --x.A_, - ali. '4.em
--i a-
! cn a high SG level signal, the low-suction and high-discharge pressure trips that could prematurely trip the MW pump were overridden. Also, the MW pump speed was increased to its cated maximum speed (595.8 rad /s), whereas the loop-A and -B MFCV flow areas were maintained at their steady-state operating values l until the MFCV overriding trip at ~100 s. At this point, the SUFCVs were opened I
l by the ICS to continue filling the steam generators. The MW pump maintained l the maximum speed setting until the trip at a high SG 1evel.
- 3. Transient _ Cal _c_ulation. Figures 117 and 118 present the secondary-side j pressures for loops A and B, respectively. Immediately following the reactor i
and turbine trips, the TSV closures produced an increase in secondary-side pressure. The TBVs for both loops were repeatedly activated between ~4 and j ~ 100 s to relieve increases in secondary pressure. The differences in the i pressure distributions between ~150 and ~250 s can be attributed to differences
- in MW mass flows to each steam generator. After the RCPs were tripped, the i MFCV override trip closed the MFCVs , and the MW was realigned to the upper header of the steam generators. The SUFCVs were opened fully by the ICS to lI
j TABLE XX i
l PORV LOCA EVENT SEQUENCE i 1
Event. T_im,e _( s)
- 1. PORV opens 0.0 ;
l 2. Turbine and reactor trip 0.5
- 3. TSVs close 0.5 1 4. Secondary-side heater and heater drain trip 1.1 ,
i 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 t j 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 i
- 16. TSV loop B opens for last time 330.0
- 17. Minimum primary pressure (~7.2 MPa) attained 550.0 [
j 18. Pressurizer water solid 600.0 [
{ 19. Maximum primary repressurization (~ll.5 MPa) 850.0 4 20. End of calculation 1000.0 i
P j 111 1
1 1
-.- .~. _._,_._ _-, . , . . , _ -
- , - , , , - ~ . . ~ . , . _ . - . . _.----,-....--.m.-~m,--..,...,,m, - - , - - . - . . , , . , ,
a , , , , , , , , . -noo 3 --w7s
~"*
n- -
-w2s ?
n- .
j f ..
- woo g
-ws k m_ _
-96o 64 - -
-325 s2 . , , . . . .
soo O mo 2o0 Soo ooo soo Goo 7oo Soo 90o woo Time (s)
Fig. 117.
SG A secondary pressure.
so , ,
- Hoo
- 78- -
' -"8 3_ .
g ,
- moo 3
f mso n_ .
f
-~ * *
- 10- \fl u_ .-990 64
_-Mo 64 - - ~e30
$2 . , , , , , , , , too o no zoo soo 4oo soo Goo too soo soo moo Time (s)
Fig. 118.
SG B secondary pressure.
112
temperatures for loops A and B are shown in Figs.119 and 120, respectively.
- Figures 121 and 122 show the MFW (realignmeat) mass flows supplied to the SG upper header of both loops. The additional mass flow to the loop-A steam generator produced the lower loop-A secondary-side pressure by condensing a portion of the steam located in the upper levels of the steam generator 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 EFW pumps were not activated because the MFW pump was able to attain the high SG 1evel.
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 it 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. After this, the pressurizer vapor f raction decreased rapidly and, correspondingly, the PORV mass flow increased. By ~1000 s, the primary-system pressure was in equilibrium.
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 those of 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 ~625 s 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 and ~650 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 downcomer liquid temperature was calculated to be ~7.2 MPa.
113
l eco , , , , , , , , , _ ,,o i
Loop A l
"'~ l ---- Loop B
. .so l
goo _
- uno
"' ~
I R g -moo 4eo- .
! -no soo- .
'; -soo soo- :
f.
too- .- 25o t 1 o- c-- o
-mo - . . , , , , , ;
o no soo soo 4o0 soo eco no eco soo ioco Time (s)
Fig. 119.
MFW flow--loops A and B.
5 us - , , , , , ,
_yo Soo -
-Se Loop A su- ----- Loop B -
g -
-in g
, 53o-
-4eo nas- , , . . . * * * ' ,-
. -aeo No-
.... ** -an so- ..- -
.....,. . eso
- ~-.' 1, *
~
' - **o sos , , , , , , .
o no soo soo soo too soo no soo soo moc Time (s)
Fig. 120.
MFW liquid temperature--loops A and B.
114
ooo , . . *' CELL m- .
isoa l
soo_ -
-us)
- soo- !
l 6
-100D m- -
n -m n g a -
g m 3o< t 2o0- -
G
,_ _ - 2s<
o- - ^o 7
-mo - -
. 224 SG nr2
-200. . . .
o 2oo e soo soo sooo Time (s)
Fig. 121.
Realignment mass flow-loop A (negative flow is into steam generator).
2mD bblL 2
800- -.p53 soO-tot) 4o0- -
W -
.m 2co- -
I 3st D- -0 SG
.3o utz
=20o , , ,
We2 o 2co .oo soo eco sooo Time (s)
Fig. 122.
Realignment mass flow--loop B (negative flow is into steam generator).
l 115
soooo , , , , , , , , ,
-secoo
- a. -
Soooo
- m. ,
l } anooo-7eo00 $
m- .
.ooooo Mooo- .
zoooo. .- mooo I
Sooo- -
- wooo moco . . , , , , , ,
o 20 anc Soc 400 300 soo 700 ooo soo 1000 Time (s) i Fig. 123.
SG oecondary inventory-loop A.
m oo , , , , , , ,
_ -ococo 3oooo- -
1 Attoo- .
10000 3oooo- .
moo- .. ,,,,, @
m .
- Soo- .-soooo
- 0000 m_ _
-soooo Bloo , , . , , , , , ,
i o no soo am aos soo soo voo ooo soo moo Time (s)
Fig. 124.
i SG occondary inventory--loop B.
116
__ - - . _ , - - _ - - - = _ - , . , - . . . - - .. ,
we , , , , , , , ,
-2200 me- -
gg. .
I- :..
I g
I me.
- teco f n- -
i
-noo so- -
m- -
-- woo ao , , , ; , , .
no o no soo soo eco soo no soo soo moo Time (s)
Fig. 125.
Pressurizer pressure.
1
- m . . , , . . ,
a u- - -do I m- -
4 so I E e- -
ga ...o g
4- -
-90
- s. .
e ,
, , , , , . , , o e no soo ano soo soo soo no soo soo moo Time (s)
Fig. 126.
Pressurizer water level.
117
.w - - - - w -w-m ,y,, wmw~ n-- - - ---w
._. . . .- -. _ _ _ = . _
oo , , , , , , .
mo j So-me ao- -
oo 3- -
I ao- -
~-
I e- -
se o- -
o
=e . .
o no soo soo soo soo ooo no ooo soo moo Time (s)
Fig. 127.
! PORV mass flow.
i t i
i 12 , ,
t ,
1 o.o - -
o.s - -
E o.a - -
o2- -
I I
o-
- o.2 . , , , , , , , ,
o no soo soo aoo soo ooo me ooo eso ese Time (s) l Fig. 128.
PORV vapor fraction.
i 118 ,
l i
sooo , , , , , , , , .
- -moo moo- Loop A1 -
Loop A2 m_ .-sooo aseo- .
. ,,oo j aooo- -
r -eooo Boo-g C ,_ .- moo C i _ -
1 l
noo. -
- - moo too-
- o- --o
-soo o no soo aio eco soo soo 76o soo soo moo i
Time (s)
Fig. 129.
Hot-leg mass flows--loops A and B.
i tooo , 1 , , ,
moo
. Loop B1 _
M ----- Loop B2 mo .-sooo seco- -
7,oo 3eco_ .
-6000 m- -
gooo_ . asco eoo .. -
- 200 sog. -
w o- --o
-noo , , , . , , . . .
o no soo ano .oo neo eso Foo aeo see inoo Time (s)
Fig. 130.
i Hot-leg liquid temperatures--loops A and B.
119
mooo . . , , , , . . .
teso son- toop A -
Loop B . ,
asse- -
Moo - -
Goto- -
geog. .
l 4
me . - sooo asoo- -
sooo- -
sooo soo- -
-mon . , , , , , . . . .
o so no soo soo soo soo no soo ooo 200 l Time (s)
Fig. 131.
Cold-leg mass flows-loops Al and A2.
1
! sos 1 . , , , , r .
4C i
soo- Loop A -
- ----- Loop B
~*
! ses- -
t seo l g m- -
seo sn- -
po I gyg.
ass -
- eno
- o. s - I
. . .o m_-
- son - -
m ede . . . . . . . . .
- o no soo see aeo neo ese no eso soo moo Time (s) j Fig. 132.
Cold-leg mass flows--loops B1 and B2.
l 120 1
soo , , , , , , , , ,
Loop Al sm- ----- Loop A2 .
sm ses soo-g .
. s.o E soo- .
l **0- -
ses I 33o- . 406 i
Mo- -
ass se , , , , , , , , ,
o no soo 3o0 doo son eco no soo soo moo Time (s)
Fig. 133.
Cold-leg liquid temperatures--loops Al and A2.
m , , , , ,
Loop B1
Loop B2 sm sm- -
i
. see m- .
E s4o ato - -
, sas 6*0 - -
.so. .o.
doo NO . , , , , , , , ,
o no soo soo ano son eso no soo eso meo Time (s)
Fig. 134.
Cold-leg liquid temperatures--loops B1 and B2.
i 121
1 , , , ,
R TH Z 118 o.e - el28 -
- 138
+148 a.. . = 15 8 .
. 168 l
os- -
o.2 -
o, -- _-
TIME (s)
Fig. 135.
Candy-cane void fractions-loops A and B.
12 , , , ,
Loop A t- ---- Loop B .
c.s - .
c.s - -
E j
os-e.2 - .
o
- o.2 o no zoo 300 m sio eo roo a6o eo moo Time (s)
Fig. 136.
Vessel upper plenum void fractions--all azimuthal cells.
I 122
- 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 reverse the main feedwater, which resulted in a MFW pump trip on a high SG 1evel. The only specified operator action included a RCP trip 30 s af ter HPI actuation. TRAC calculated a minimum downcomer liquid temperature of ~528 K at ~600 s into the i transient. The primary system was calculated to repressurize to ~11.5 MPa af ter
~800 s.
D ._ TBV_ Failures
- 1. One Ba_nk 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 return to a closed position. This occurred af ter they first opened 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 f ailure of the operator to throttle the HPI 575 , , ,
~*
s,0 R TH Z .
216 _,,,
- e226 -
- 236 ;33, E 246 3_ 256 . -se l_- .' - "
646 -_ -
_g
-,to T m S35- % .4 e MO - '
- 490 j S25 . . . ,
l 0 200 400 600 000 1000 TWC (s)
Fig. 137.
. Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors).
123
system. The lowest downcomer liquid temperature (~458 K), and hence the smallest margin against the NDT limit, were 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 j 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 pumps deliver the condensate to the feedwater heaters. The main feed pumps then pump the feedwater to the steam generators.
The significant features of the TBV failure transient (one bank of two TBVs) are the following:
e Reactor and tarbine trips cause the TBVs to open.
- Failure of one bank of TBVs to close causes a secondary-side depressurization through the affected loop.
e Failure of the SG liquid-level control in the affected loop follows initiation of emergency feedwater.
e The operator does not restart the RCPs.
9 e The operator does not throttle the HPI.
Two parametric cases were also calculated. The SG liquid-level control in the affected loop operates correctly in Case 1. The SG liquid-level control also operates correctly in Car.e 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 124
- c. Results.
i ._ Base Case. Table XXI presents the calculated event times for the base case. Following the reactor and turbine trips, the TSVs closed (at 0.5 s),
l j secondary pressures rose, and the TBVs opened for the first time at ~ 4 s. The secondary pressure peaked and then decreased to permit the loop-B TBV to resent at ~45 s. However, the loop-A TBV failed to reseat, which caused the secondary pressure to depressurize at a faster rate than that of loop B.
l
- The pressuriser pressure is presented in Fig.138. The PORV opened at
! ~1037 s when its pressure setpoint of 16.99 MPs was exceeded. The PORV then cycled 'for the -remainder of the calculated transient to maintain the primary-system pressure at or below the PORV setpoint.
l The loop-A (affected loop) and loop-B secondary pressures are shown in Figs.139 and 140, respectively. The open TBV in loop A caused that loop to depressurize more rapidly than loop A and to a lower level. MFW flows for both
.l loops A and B are shown in Figs.141 and 142, respectively. As specified,3 the 1
i
! TABLE XXI t TBV EVENT SEQUENCE (BASE CASE)
Event Time (s)_
i
- 1. Turbine sad reactor trip 0.5
- 4. .TBV loop B opens 4.3
- 6. HPI . started following trip on low pressure 153.1
. 7. RCPs trip on 30-s delay af ter HPI actuation 183.0
- 8. Feedwater realignment trip 183.0
- 9. Main-flow control. valves overriding trips 183.0
- 10. EFW pump on 209.1
- 12. PORV opens 1036.7 i
i 125 i
, - - ~
+ . - . _ , . - , - , , - , . . . , - , , , . . - . , _ - - . . . . .
30 ,
-29o0 f* t AMSWVA
.0 ,- .
y
./ BASE arso PAR &WCVNC t too -
- Q PARMETmc 2 --2000
/*,.
- -U50 30- -
l 1
,/". "
WO-
- t e0- x, ,
. .,0 l
- 200
.0 , -
-750 40 , .
0 200 400 600 800 1000 SCO 4s00 # 30 M (s$ prwsww (bors)
Fig. 138.
Pressurizer pressure.
e9 , ,
-neo 3;- .
- n. MK .
PARAM[TWC 1 - 900 PMmETmc2 o0 .
. 00 30- -
40 . .
So9 v.s 20
'h %.* .
-no to-
.soo gD - ~ -
o . . . . . . .
e o aos eso soo see moo noo noO meo w (p)
Fig. 139.
I SG secondary pressure--loop A.
l l
l 126
-_ e-- -----, e
so , , , , ,
.noe so- -
saw PuaWETinC 1 uso 70- PARM C ec 2 -
oo.
.soo T N
3 s .,'
1oo. ,
\ .,'- .. . . .
y-
\ .
800 so-
's'N , ..
x ~~ .
So . . , ; . .
'M e zoo ao soo soo woo noo woo woo WE (s)
Fig. 140.
SG secondary pressure-loop B.
soo , , , , , , , -eso no- -_,
ooo. sAst '
PARAWETWIC 1 - BSc 4 PARecac 2 I h soo- -
i m- -
l soo ago. .
. . me
- o. :
..e
-mo , ,
e ano ao eso soo moo eso woo moo MM Fig. 141.
MFW flow--loop A.
t 127
ICS failed to reverse the main feedwater to the loop-A steam generator, and thus the flow did not decrease until the MFW trip on a high liquid level in the loop-A steam generator occurred at ~61 s. The ICS reversed the 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 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 feedwater 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 Fig s. 145 and 146, I 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 the TBV at loop A was open and produced 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 loop-B I valve. A small flow of hotter fluid through the SUFCV produced the temperature j 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 steam generator rose before ~60 s because the ICS f ailed to reverse the main feedwater. The MFW trip occurred at ~60 s, and the liquid inventory boiled' off until the EFW pump actuated t ~210 s. With no liquid-level control, the SG secondary continued to fill and oscillations developed as the liquid level reached ~12.4 m. The :
j transient history of the loop-B steam generator was very dif ferent. There was no initial increase in the SG inventory because of the ICS reduced flow.
Closure of the loop-B SUFCV by ICS action at ~600 s was evident.
Mase flows through the primary-loop hot legs are shown in Figs.151 and ,
' 152, respec:tively. Following the RCP trips at ~183 s, the flows coasted down L 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 the TBV at loop A to reseat. Similar phenomena were observed in the cold legs, as shown in Figs.153 ; through 156. The corresponding cold-leg 1 temperatures are presented in Figs.157. through 160. Vapor fractions in the 128
l l
l eso , , , , , -sno !
l i
no-. -
.go. ;
l ooo_ sast .
. mutAuffitC 1 geo ,
mWUKTIuc 2 '
.aos 5-
.. 5 3oe. -
too aso- -
,. .-m O.
w -- --0 e ago eso too Goo IBoo Soo M Soo M (s)
Fig. 142.
MW flow--loop B.
T 5 sast PARAWETmc 1 -d45 se- ,
mutAMETitC 2 .
N.-
- mo
""~ \
E N, 'N, -
= E l eso-_
\ .,
N,, \~s . mo 40o-
..e i _ .
-- 1 m
doo- .
3eo ato ,
e soo o soo soo moo soo ice.
M (s)
Fig. 143.
MW liquid temperatures--loop A.
129
ses , , , ,
seg m
m- mRA6Cfmc t maAupac r- -
,f, . . . -
.0 sas- ,
E se0-
\, -
-= E as-
\ s -
.e0 <
gg.- * .-4e0
.- . m0 se0-. \, -
4e N
as .
O soO =0 eOO e00 =00 n00 =00 w00 ThC (s)
Fig. 144.
MW liquid temperatures--loop B.
300 ,
-600 pso. -
.Sco 200- s BASE -
dos PARAWCTRIC 1
.a tQ PARA 6CTRIC 2 g e,
R E
300 W g ~-- __...-...... , ,
g .0 g
g -
./. 2*
g sO- -. ,,o l
J 0 --O
.g - -10 0 0 200 400 000 000 1000 900 #00 1000 ThE (s)
Fig. 145.
Flow through EW header--loop A.
.130 s
co . . . . ,
-36 4
80- - e gast .-95 NRaWETRtc 1 l
m9 Panautac 2 so- , .
s
./V -e5 h
- .e- , m l.
-a Q 2
i 3
- \' ., '
35 0
--O
-20 . , ,
o ano .oe soo soo moo .oo .oo w (s)
Fig. 146.
Flow through EFW header--loop B.
ii Soo
- c. ' N ,.. W-w . _
^-405 l
- ="=====..
450 f
/ 3e0 E .
j
'/ d %.c, Iphnesh) MRAMETRIC 2
.->. E
.-M I.co-
- /
m-M 350-~ .
20 90 300- -
-'5 m.' .
=o e
M (s)
Fig. 147.
Liquid temperatures in the EFW header-loop A.
131
ses , , , ,
O eco- -
f w._. . ..
' 400 l
/ .
/ I E "' h putscmei ""
, / F w / Futacme t 2 as-.
, /
l
-.soo f
~-25o m- i (h -
-2oo @
me- l l
- 8
- , -so m- jf -
J i -m no. .
50
+
m .
o ao eco soo soo woo aco woo woo w (s)
Fig. 148.
Liquid temperatures in the EFW header-loop B.
ococo
-woooo foooo- -
\ l .-
/' vp, i -uoooo soooo. l
// /
PAR AWETmC 1
-..oooopk g sowo- ,
j f rutmome 2 2 **000 I E o. p // ,/ .
s
/ / -cocco i meno. , .
-ooooo Nooo- -
nono . , . .
e ano eso eso soo woo ano woo see w (s)
. Fig. 149.
, SG secondary inventory-loop A.
l 132 i
soooo . , , ,
-folooo
_f oco**
.coe.-
i l -
/ utTINC 1 ~-75000 $
g g
2 aoooo-g
- l/
/ l/ (chnas9 PAnantac 2 5
. m oo @
s mooo.
o
//l E
mee .
l .-eooo Sooo- -
-SooCo moeo o ano ao soo soo moo noo woo woo TBE (s)
Fig. 150.
SC secondary inventory--loop B.
4 I a
- .uoco ocoo- .
sAst -8000 Scoo- mutAutTmc 1 PARAhCTIoC 2
- sooo m- .
._ i f
sooo- ,
a I --- , -..
0 m
1ooo-. .
.~
e-- ...
-ioco , , , .
. --sooo ao e aos eso soo woo noo woo o 79 4 (s)
Fig. 151.
Hot-leg flow-loop A.
133
mo , , , , ,
f seopo oooo- -
east -"
3000 PiutAMETHic 1 -
N2 .moco
$ deso- -
S 4 .eco, g j f3000 '
6
.oooo C sooo- h e
-., 2 woo- .
_ oo, o- --o
-moo .
--2000 o ano 4co eco ooo woo noo woo soc M (s)
Fig. 152.
Hot-leg flow-loop B.
mo
-vooo acoo. _. .- --- -
sAst -"
Scoo- PnNAu[TRIC 1 -
a PARAaCTRIC 2 f -eooo
{
t
- m. -
, . coco 1~
M- -
ocu zooo-_ -..
soo- '
-.me o- -
o
. coa , ,
--seos o zoo 4eo ooo ooo woo woo woo soo M (s)
Fig. 153.
. Cold-leg flow--loop A1.
134
sooo , ,
r
-sooo oo. .
-oooo l sast
! sooo- h) Panautme t .
t PARAhpoc 2 -sono g
.a h woo-.
L -.mo f I o-- '
--o e 2
.l
-moo.
.-- m o s - - - - --
___ _..sooo
..oooo
-mo .
o m 4o0 soo ooo moo uoo woo woo M (s)
Fig. 154.
Cold-leg flow--loop A2.
mo ,
weoo sooo- - -- -
pst "
seco- sh) P4RaMETmc 1 -
. PAmpec 2 soon. -
-ooCo h m-h
~
.mo h sono. _
icoo- '
-. ooo o- --o
-moo .
--sooo o 3Do - doo oco ooo tooo Soo seco Goo M (s)
Fig. 155'.
Cold-leg flow--loop Bl.
135
sooo moco
~
4000- .
- Sooo BASE Sooo- PARAWCTEC 1 -
PARAhE110C 2 8000 maa
{
h soo-. -.m. f m
E '4 m o-- - o 3
-Woo.- .-=2oo0
. 4o00
-sooo o aoo 400 soo soo moo nao woo woo M (s)
Fig. 156.
Cold-leg flow--loop B2.
ooo ,
soo SP5 - -
sso ano- m .
PmRAmtflec 1
- ~*
g nas. PARAtETitC 2 .
w ,,
N s , % - -- E 5 soo. '. --_ .- * *
- g s,..
g.- .
.-doo me-- ..no
..,,, 1
.o 375- -
No , , , , . , ,
e ano ao soo soo moo moo woo see w(s) l Fig. 157.
Cold-leg liquid temperatures-loop Bl.
136
l eco .
soo i
s>s - -
soo m- un -
m AutTmC1 3,,
g ,, . numenac 2 .
-.s
--. E I.. g*
..oo Go- -
36o
- e. . ..,,, I 3P5- -
3oo ato , , . . , . .
o ano ao eso soo woo soo woo woo w (s)
Fig. 158.
Cold-leg liquid temperatures-loop B2.
eco , , , , ,
soo neo- -
-54 sAu Seo- PARAWCTmC 1 .
g PAAMENRC 2 'W Mo-- -.,e go- . ano soo-
\.- . .
-no
~~~.--
oo eao- ,'.~.... -
doo- -
me 4eo . . . ,
e ano eso eso eso eso soo woo soo MW Fig. 159.
Cold-leg liquid temperatures--loop A1.
137
i 1
4 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.
Downcomer liquid temperatures for the base case are presented in Fig.164 at the top axial downcomer 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 level 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 will not restart the RCPs and not throttle the HPI. The event sequence for this case is presented in Table XX11. 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 ~60 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 compared directly with the base case. The general trends of Case 1 were similar 1
to those of the base case. The major differences appeared in the secondary side j of loop A and were caused by shutting the loop-A EFW valve on high SG liquid level at ~290 s. The reduced EFW flow affected the primary side also. Compared with the base case, the pressurizer pressure increased more rapidly to the PORV setpoint as shown in Fig. 138. This was caused by reduced primary-to-secondary heat transfer associated with the reduced loop-A SG secondary water inventory (Fig. 149). The SG secondary pressure histories for loops A and B (Figs. 139 TABLE XXII TBV EVENT SEQUENCE, PARAMETRIC CASE 1 Event Time (s) i
- 1-10. Same as base case 0-209.1
- 138 l
eco . . . , , .
soo Ste- -
vo sAst seo- MAWTRC1 -
mad 7IDC 2 W '
g See- .
-Se l
See- --4e0
-no W
' \ ~-- ' N . h.
sec. -
s,
- sao neo- . .
seo eeo- -
ado . , . , , ,
e ano eso eso soo sooo soo woo moo w (s)
Fig. 160.
Cold-leg liquid temperatures--loop A2.
au , . , ,
tes- .
sAst MAWTWC 1 MAdTitC 2 i
88" -
e _ .
-em- -
4 es- -
4 e8 . . . ,
e ano ano eso soo uso soo woo see M (s)
Fig. 161.
Candy-cane vapor fraction--loop A.
139
tas , , , , ,
Sns- .
BASE PaRAnotC 1 MRAaGBC 2 tes- .
4 80 .
1
-sas- .
-ELes - -
-60s .
e age ee0 a00 300 geno 300 me0 300 w (e)
Fig. 162.
Candy-cano vapor fraction--loop B.
n .
BASE
) PanAutme 1 -a
- m. 39 PaRAutTInc 2 -
-30 8- -
25 s-
,/ . 20
- f. ,
s 2- *,',
-* * , , . ' f~~.,
.3
-2 , ,
9 300 400 800 800 1000 000 WOO S00 m (ss =* w (<m) l Fig. 163.
! Pressurizer water level..
140
and 140) were similar because , the level control operated on both steam generators. The closure of the loop-A EW valve limits the flow through the EW header, as shown in Fig. 145. The remaining flow through the header comes through the loop-A SUFCV. The loop-B EW header flow (Fig.146) decreased 'to i 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 axial 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 downconer temperature for Case 1 was caused by the reduced heat transfer to the loop-A SG secondary with its reduced liquid inventory.
11_1. - Parametric Case 2. The specifications for this case differed from j those for 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 i subcooling; and throttling the HPI was permitted on attainment of 75 12.5 F i
i subcooling. The event sequence for this transient is presented in Table XXIII.
I Events 1-10 were identical to the base case. - One RCP in each loop was restarted at ~383 s af ter a 30-s delay following the subcooled monitor trip. The HPI was throttled at ~485 s after a second subcooled-monitor trip at 75 12.5 F j subcooling.
I s
TABLE XXIII 1
l' i'
TBV EVENT SEQUENCE, PARAMETRIC. CASE 2 j Event Time (s) i 1-10. Same as base case 0-209.1
]- 11. - Loop-A EW valve shuts 290.2 f 12. Restart RCPs in one loop after subcooling 383.5 I
monitor trip
- 13. Loop-B EW valve shuts on high SG liquid level 395.1
- 14. HPI throttled after subcooling monitor trip 484.7
- 141 t
,e - , , , , - - - -.- - , . . . - , w,, - - - - -
SM- eTHETA = 1 o a THETA = 2
+ THETA = 3 550- = THETA = 4
+ THETA = 5 h . THETA = 6 - 510 p M 530 w a: a:
E -470 8 m 510- 5 E -430
- H 490- N
- g. -390 450 , , , , , , , , 350 0 150 300 450 600 750 900 1050 1200 1350 1500 Fig. 164.
Downcomer liquid temperatures (base case) at vessel axial level 6 (all azirithal sectors).
580
- THETA = 1
- THETA - 2 -564 i
+ THETA = 3
~
- THETA = 4
+ THETA = 5
{ ' THETA = 6 -524g a 540- s w
m -
m D D D '
4
- a. 520-
-484$ c.
E
- r N g_ \ -444 1
4a0 , , , , , , , , , T 4 04 0 100 200 300 400 500 600 700 000 900 1000 1100 Fig. 165.
Downcomer liquid. temperatures (parametric case 1) at vessel axial level 6 (all azimuthal sectors).
142
l Results for Case 2 are presented in Figs. 138 through 163 and may be compared directly with the results of base case. Case 2 displayed significant l
differences from the base case. A major consequence of throttling the HPI was l that primary-sys tem repressurization did not occur and thus the PORV did not open. The absence of repressurization is seen in Fig. 138. Restart of the loop-Al and -B1 RCPs can be observed in Figs. 153 and 155, respectively.
l Operation of the RCPs induced a reverse flow through the cold legs with nonoperating pumps. The influence of RCP restart on heat transfer to the loop-A SG secondary can be observed in Fig. 139. The pressure increased with RCP l 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 SG secondary water inventory (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 SG 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 loop-A and -B steam generators is less than in the base case, and this is evident in the primary-system temperature. Downcomer liquid temperatures for Case 2 are presented in Fig. 166 at the top axial downcomer level (just below the cold-leg nozzles). At the end of the calculated transient (1500 s), the minimum temperature was ~491 K. This was ~33 K higher than the temperature of 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 depressurization transient was simulated using TRAC-PFl. The transient studied was the failure of one bank of TBVs (two valves) to close after initially opening af ter the reactor and turbine trips. The base-case transient included additional failures caused by f ailure of the level control in the affected steam generator, no operator restart of the reactor-coolant pumps , and no operator throttling of the HPI system. A minimum liquid temperature in the downcomer of
~458 K at 1500 s was calculated. If correct operation of the SG 1evel control, operator restart of the reactor-coolant pumps, and throttling of the HPI flow were assumed, the primary system would not repressurize and a minimum downcomer liquid temperature of ~491 K would be calculated at 1500 s.
j 143 1
4
~
i . THETA = 3
- THETA = 3 NO-
- THETA = 3 ,,
= THETA = 4
-
E =- - -- -dd n
g.*00- -asu$
E e- -age 300- -2223 i
WO- , , , , , , , 1523 8 3000 3100 3100 4000 8000 0000 1000 Fig. 166.
Downcomer liquid temperatures (parametric case 2) at vessel axial level 6 (all azimuthal sectors).
- 2. Two Banks of Two TBVs
- a. Introduction and Summary. This case differs from the previous case (Sec. III.D) by assuming that two banks of TBVs will f ail instead of one bank.
For this study, the performance of the Oconee-1 plant after a secondary-side depressurization was predicted. The base case analyzed was the failure of two banks of TBVs to reseat af ter initially opening af ter the reactor and turbine f trips from full power. Additional failures assumed for the base case were f
failure of the level control in the af fected steam generators, failure of the operator to restart the RCPs, and failure of the operator to throttle the HP1 system. The lowest downcomer liquid temperature (~445 K), and hence the smallest margin against the NDT limit, were again calculated for the base case.
Repressurization of the primary system to the PORV setpoint was also predicted 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 f ailure transient (two banks of two TBVs) are the following:
144
_ . . . - . , . . . .._ . - . - . _ . - - . . . _ . .~.
- Reactor and turbine trips cause the TBVs to open.
e Failure of two banks of TBVs to close causes a secondary-side depressurization through both loops.
- Failure of the SG liquid-level control in the affected loops follows initiation of the emergency feedwater.
! i e ~ The operator does not restart the RCPs.
- The operator does not throttle the HPI.
( Two parametric cases were also calculated. The SG liquid-level controls in the affected loops operate correctly in Case 1. The SG liquid-level controls also j operate correctly in Case 2. In addition, operator. actions to restart ths RCPs i and throttle the HPI are permitted if the primary-system subcooling monitor trip 1
points are exceeded.
j .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 (at 0.5 s),
secondary pressures rose, and the TBVs opened for the first time at ~4 s. The i secondary pressure peaked and then decreased, but both banks of TBVs ( four valves, two on each line) failed to reseat. Continued flow through the TBVs l ,
resulted in a secondary-side depressurization. ,
l The pressurizer pressure is presented in . Fig. 167. The PORV opened at
~1175 s when its pressure setpoint was exceeded. The PORV .then cycled for the f
remainder of the calculated transient to maintain the primary-system pressure at 5
or below the PORV 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. As specified,3 'the ICS failed to return the main feedwater to the j steam generators and thus the flows did not decrease until' the main feedwater tripped on a high SG B liquid level at ~91 s. A higher ' mass flow was predicted
- j. for loop B before the MFW pump was tripped because the MFCV area under ICS l control' was .~6% greater - than in loop A. Loop-A and -B MFW temperatures are presented in Fig. 172 and 173, respectively, i
l t 145 i.
TABLE XXIV TBV EVENT SEQUENCE, BASE CASE i
Event Time (s)
- 1. Turbine and reactor trip 0.5
- 2. TSVs close 0.5
- 3. Loop-A TBV opens (fails to reseat thereaf ter) 4.1
- 4. Loop-B TBV opens (fails to reseat thereaf ter) 4.3
- 5. HPl begins following trip on low system pressure 87.5
- 7. RCPs trip on 30-s delay af ter HPl actuation 117.4 -
- 8. Feedwater realignment trip 117.4
- 9. Main-flow control valves trip 117.4
- 10. Emergency feedwater pump on 147.0
! 12. PORV opens 1175.7 I EFW flows through loops A and B are shown in Figs.174 and 175. These flows were initiated at ~150 s. Because the EFW level-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 f with the emergency feedwater before entering the steam generators. 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 j more rapidly before the MFW trip because, as previously discussed, the loo p-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 level reached ~12.4 m (top of the generators).
t i
146
so , . , , , .- ,
-seco
- - j
~
.sseo r
- ./, /
we- muwste t /
.seso mumpuc 2
./
no- j - '"
. ; -moo me. ,.-
t .-
l j
..* aso i
se. -
/ '. ..- neo
- n. ~
-~
no s-
././. '. .
soo 3o . . . . . . .
o zoo ao eso soo woo soo woo woo M (s)
Fig. 167.
Pressurizer pressure.
so , , , , , , ,
.ano so- -
m- .
%Mnowatc, 2 -
moo ee. .
I 90-ooo
~
as- - '
\, '
-me as- -
, .. \ ~...
........ ~ ~ . , ,y i
n- _
e . . . . . . e ao e ano eso ese moo see wee meo M(4 Fig. 168.
SG secondary pressure--loop A.
147
se , , , , , , ,
. see
- n. -
N- .
me.
9 PiutMGec 2 n- -
I g
. een 1 ee- -
- ~***
- a. -
se. .
. .noe 3- -
. .\. .~..w. ....
- m. -
e . . , , . . , e e ano aos ese eeo uso see wee neo MW Fig. 169.
SG secondary pressure-loop B.
Me , , , , , , ,
. wee see. .
g .nso see- PiutAutfmc 1 -
m2 ,,,,
$ age. - p g
4 .
ase- -
399- -
. .m me- .
.ese
- e. . A- -
. - e
-me "
--see
- e see see see see see see wee ses
! MW Fig. 170.
MMW flow--loop A.
148
me , , , , , , ,
. uso see. -
g neo see. PiutAML1 tic 1 -
, (men)m 2 ,,,,
{ .o . ; . :p 4 -[ eee f gg . f .
E i soo E ase- -
se .
soo h..sr e- =_ __ _- .
l
-me '
--ase e see me eso soo neo neo teso sco MW Fig. 171.
MFW flow--loop B.
ano , , , , , , ,
en. .
me east see. PiutAhpIIC 1 -
g - , PlWtAhCBC 2 436 88- . \ -
ee
% d. .
] me. - ' . , N
. as
...'% , K,
' ano es. ..., -N, .
m 5 me . .
l -
seo me. -
Set i
480 , , , , , , ,
e see me eso eso neo uso weo see MW Fig. 172.
i MFW liquid temperatures--loop A.
1 149 l
seo , , , , , , ,
. des so. .
.go aAst age- PARAETIBC 1 -
PWtAKTIBC 2 8 L 4ae E
me . 7 s .
-me I
.e. .
...N
. . ..se
..h '
g ,,
... .' ..h 'g, ano 5 -
% m age. .
3eo se ,
1' 440 , . . . . . .
e aos ano eso ano moo noo woo woo MW Fig. 173.
MFW liquid temperatures--loop B.
ano , , , , . . .
noo ase- -
. mo BASE
, es. PWtAWTmc 1 PapKTmc 2.,u aos 9 g
- p. aw :. : - -
- 1 --. -
..o i I .. .
..e i
- e. w - e
-se
. --ee o sie aio ein eie s 0 tese see MW Fig. 174.
EFW flow--loop A.
I 150
ano . . , , , , ,
. 3e0 go. -
. me aASE go. Piutedttmc 1 .
.y e v~g2, w
- aos SP g
ag. -
E , ano C
.o. . -
..e I
e- " - o
-ee e aos me eso soo eso eso woo moo MW Fig. 175.
EW flow--loop B.
ens , , , , , , ,
ego. .
30 F
as. -
k . me g
~'%-
g -- -
J4JL.C, -
o E
88- ,
W2 -
.aos me. -
35o I s Q ms. % .
g -
ano me . .
se ass- -
ses-d me ae as , , , , , , ,
9 age me ese soo neo ese teos see MN Fig. 176.
EW liquid temperatures--loop A.
l 151 l
1
see , , , , , , ,
gg.
as-
.%.~~N - -
ase E .e .
%"g, .
E me.
\ .
I ..
\. ase 1 .e.
g
- -538
- m. .
ase-
/ -
.e.
a m . . . . . , ,
e ano eso eso eso uso neo isso isoo MW Fig. 177.
EFW liquid temperatures--loop 3.
esooo , . . , , , . .
Bast
- mutAwmc i .mosso
- nome- RWLAdTisc 2 .
1; .-
eseen. ,
t .
~
- oeene- .
~" l
- / neone l_. - e
/ .
.aosso nl aeone- ,
/ .
/ sosso acons- .
, e,,,
sees . . . . . .
aos aso eso aeo eso me ese eso eso see e me MW Fig. 178.
SG secondary inventory-loop A.
l l
152
Mass flows through the pcimary-loop hot legs are shown in Figs.180 and 181. Following the RCP trips at ~117 s, the flows coasted down and natural I circulation flows through both loops were established. Similar phenomena were observed in the cold legs, as shown in Figs.182 and 185. The corresponding loop cold-leg temperatures are presented in Figs. 186 through 189. Vapor fractions in the loop-A and -B candy-cane sections are shown in Figs.190 and 191, and it can be seen that no voiding occurred during the transient. The i pressurizer water level is presented in Fig. 192.
Downcomer liquid temperatures (Fig.193) for the base case are at the top axial downcomer level (just below the cold-leg nozzles). At the end of the calculated transient (1320 s), the minimum temperature was ~445 K. The minimum downcomer . 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 1
compared with one, steam generators.
ii. Param_etric Case 1. A single specification was changed for this parametric study. The loop-A and -B SG 1evel controls following EFW 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 HPl. The event sequence for this case is presented in Table XXV.
The event sequence was identical to the base case through event 9. At ~ 147 s the loop-B EFW valve shut on high SG-B secondary liquid level. The loop-A EFW valve shut on high SG A secondary liquid level at ~373 s. The PORV opened ~13 s early in Case 1 because of reduced heat transfer to the two SGs.
TABLE XXV TBV EVENT SEQUENCE, PARAMETRIC CASE 1 Event Time (s) 1-9. Same as base case 0-117.4
- 12. PORV opens 1062.1 153 4
7 5
soooo , , , , , , , ,
BASE esooo-PMLA6ETRIC 1
-sooos 5nast) PWLAndCTRC 2
. oooo mose- -
^
"Y '.
/s
'\ - secos g googe.
l g' /;"IT\ Y. '.. j n$
( -
-nocoo y essee. .
L 3
, -1000oo
/
moeo-j
/ -
-accoo geogo. .
.ooooo anooo-, -
mooo , , , ,
o ao no aoo ao soo soo no soo eco moo veo TIME (s)
Fig. 179.
SG secondary inventory--loop B.
mac , , , , , , ,
enee.
l .
moea
. EASE . gogo PMtAdTmc 1 seee- FMtadTinc 2 -
. .moso
- m. .
.e.oo g .
y
~
w- - - -
. .aooo asee- .
e i mes- .
sees
^
~ :. , .
i e . . . . . . . e s ase me eso eso neo use neo moo l
ltE ($
, Fig. 180.
Hot-leg flow--loop A.
154
mo , , , . . . .
[ woon asso. .
g -Gooo seco. PARAuttmc 1 -
PARA wTec 2 '
$ seso. -
Q 4 . .oooo f l ._ - - - -
. coco 1
asse- ,
8 -
ese-. ,
.oooo 1
... . .o
-Soo
~~ M o ano ano eso soo isoo moo woo see MW Fig. 181.
Hot-leg flow--loop B.
mo , , . . . . .
g noon sooo- PARA WTRC 1 -
N2 2o00
/
esso. .
7 8
oooo g 1 -- .
.oooo I asso- -
isso- .
.oooo e- . -
o
" '*3
- Woo , , , , . . .
q aos ano eso oso woo son woo moo MW Fig. 182.
Cold-leg flow--loop A1.
155
esso . . . . , , .
g moso seee. .
.eSoS aAst ages- PMtAETIBC 1 .
mutAETsuc 2 seco asse- ,
s- .
~
.. g
.. . . .o i . -anoo
- --- . . . ._ -4ooo
. ooo
-aeos , , , , , , ,
e ano ano eso eso moo non seco uno MW Fig. 183.
Cold-leg flow--loop A2.
sees , , , , , , ,
.- . . .. wo
. m noso sees- . mutAwTsue 1 -
2 g
seen- .
.geen m- .
m- .
e neo- .
I e- - .
o
-mos , , , , , , ,
-.asos e ano ano eso aos moo uso meno moo MW Fig. 184.
Cold-leg flow--loop Bl.
156
l 8888 . . . . , , .
y moso neee. .
m Bast asse- mutmETsuc1 .
hWtAW.TiuC 2 ecco asse- .
moo 888' -
asoo e- - . e o.
. . --aeoo
. moo
.o
-seos , , , , , , ,
e ano 4eo eso soo woo neo woo neo M (e)
Fig. 185.
Cold-leg flow--loop B2.
soo ,
l ses too- -
geo- BAtt .
PAR AWITINC 1 g --
PARAMETIUC 2
.m E
h -
soo W
< m. .
R,
\ E me g soo-o.
\g\ ~~.
g
. .. -.......... ~ ~ ' m .. -ns eso. .
-mo aso-se M - , . . . ,
o ano eso soo soo moo soo moo moo Ted (s)
Fig. 186.
Cold-leg liquid temperature--loop A1.
I 157 l
l
, . . . . ~ - - ,-
soo , , . , , .
ses noo. .
-See geo. BASE .
PM AtotflhC 1 g neo-Pumacnec 2 m p doo
- m. .
\ #
-ee g
\pN 'N. .'.
-m W
ago.
--........ --.m.. .
-m
-3eo
, 440- -
se m
, o ano 4co soo soo moo noo woo moo l
M (s)
Fig. 187.
Cold-leg liquid temperature--loop A2.
soo , ,
SOS Seo- -
sAst
~"
3.o . -
PAR AndETRIC 1 g m-9 PAaaacec 2 m
-E 5 ..so y m- ,
N "
ass l
W
.4~ '
4eo. *..,% _.
....~
_" m seo 4.o. -
-se m , .
e zoo ano soo soo moo moo woo neo M (s)
Fig. 188.
J Cold-leg liquid temperature-loop Bl.
158
&M . , ,
Ees- -
W BASE i dosFV RWlAWCT1 tlc 1 I mutA6ETIUC 2
, e.or- -
E too -
-on -
-4.oe - -
-&M ,
a 300 4o0 000 000 80o 0 Goo 14 o 0 15 o 0 M (s)
Fig. 189.
Cold-leg liquid temperature--loop B2.
cM . . . .
E04 - -
sAst dos mutAutfitiC 1 PARA 6CTitC 2 0m- -
E &m -
I -em - -
-e.o - -
-&M .
e ano 4e9 soo soO moo soo isoO moo M (s)
Fig. 190.
Candy-cane vapor fraction--loop A.
159
em , , , . . -
i su. .
sast Pe wtTmc1 PWtANETleC 2 sat- -
ano -
1
-em - -
-su - -
-&M ,
e ano eso soo soo moo soo woo moo M (4 Fig. 191.
Candy-cane vapor fraction--loop B.
a , , , , , , ,
.m e- .
SAR 3o do PWtAnoluc 1 e- -
.m E g g
4- \ .
.i .e B
' ~
['N,~.%. s
~J . ..
..s
-t . . . . . , ,
9 ano aos eso eso inoo ano isso use MW Fig. 192.
Pressurizer water level.
160
Results for Case 1 are presented in Figs. 167 through 192 and may be compared directly with the base case. The general trends of Case 1 were similar to those of the base case. The 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 with that of the base case, the pressurizer pressure increased more rapidly to 1
the PORV setpoint, as shown in Fig. 167. This was caused by reduced primary-to-secondary heat transfer associated with reduced loop-A and -B SG secondary inventories (Figs. 178 and 179).
Downcomer liquid temperatures for Case 1 (Fig.194) are 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 was ~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 1
subcooling, and throttling the HPI was permitted on attainment of 75 1 12.5"F subcooling. The event sequence for this transient is presented in Table XXVI.
Events 1-9 were identical to the same events for the base case. The HP1 was throttled at ~421 s af ter 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.
I Results for Case 2 are presented in Figs. 167 through 192 and may be
- compared directly with 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 4
legs with nonoperating pumps. The influence of RCP restart on heat transfer to
! the loop-A and -B SG secondaries can be observed in Figs. 168 and 169, i respectively. The pressure increased with RCP restart and continued to increase s
to the end of the calculated transient. The increase of the primary system pressure (Fig.167) was terminated af ter the HPI was throttled at ~420 s. The 1
primary pressure decayed rapidly to the accumulator setpoint of 4.168 MPa at
{
j ~565 s. Water at 305 K was then injected into the primary for ~20 s. The j 161
- ,. ,._ - = - . , _ - . . - - - , - , _ _ ~ _ . . _ _ - _ _ _
800
- THETA = 1 - 572 j
- THETA = 2
+ THETA = 3 I
- thera = 4 - 532 I
- THk7, A = 6 m 640 g
- THETA = 8
-492 h I
- 680- N j D
t--
-452 g 800- W o.
g 2 F g, ~ 412 N
- g. - 372 440 , , , , , , , , , 332 0 ISO 300 450 000 750 900 1060 1200 1360 Fig. 193.
- Downcomer liquid temperatures (base case) at vessel axial level 6 (all azimuthal sectors).
J Sao e THETA = 1 - 572 a
- THETA = 2 000-
+ THETA = 3 a THETA = 4 - 532 m 640-
- THL7A = 6 ;
-F
- THETA = 4
_4gg -
b
$Sao- $
g o 4 - 452 N 000-6 n.
l
- 412 e-5 l m-
~ ;
. -372 440 , , , , , , , , 332 0 ISO 300 460 000 760 900 1060 1300 1350 Fig. 194.
Downcomer liquid temperatures (parametric case 1) at vessel axial level 6 (all azimuthal sectors).
162
TABLE XXVI TBV EVENT SEQUENCE, PARAMETRIC CASE 2 Event Time (s) 1-9. Same as base case 0-117.4
- 12. HPI turned off 421.0
- 13. Restart RCPs Al and B1 af ter subcooled 517.0 monitor trip
- 14. Loop-A accumulator begins discharging 565.5
- 15. Loop-B accumulator begins discharging 565.5 primary pressure then increased until it was turned around by the increased heat transfer to the secondaries by RCP operation. The accumulators again discharged near the end of the calculated transient, thereby increasing the rate of l d owncomer temperature decrease. This was the only TBV transient that experienced accumulator discharge and it markedly influenced the extrapolated downcomer temperatures at 7200 s.
Downcomer liquid temperatures for Case 2 (Fig.195) are at the top axial downcomer level (just below the cold-leg nozzles). At the end of the calculated transient (1500 s), the minimum temperature was ~467 K. At 1062 s, the minimum temperature was ~477 K, which compares with ~465 K for Case 1 and ~453 K for the base case at the same time.
- c. Conclusions. The response of the Oconee-1 plant to a secondary-system
- depressurization transient was simulated using TRAC-PF1. The transient studied was failure of two banks of TBVs (four valves, two on each steam line) to close after initially opening after the 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 no operator throttling of the HPI system. A minimum liquid temperature in 163
the downcomer of ~495 K at 1320 s was calculated. If correct operation of the SG level control, operator restart of the reactor coolant pumps, and throttling of the HPI flow were assumed, the primary system would not repressurize, and a minimum downcomer liquid temperature of ~472 K would be calculated at 1320 s.
E. Hot-Leg Break LOCAs
- 1. TVo-Inch 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 and turbine tripped from full power. Reactor decay heat was specified as 1.0 t imes the ANS standard. Fo r this transient calculation, the ICS and all key system components were asstned 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 investigated. One calculation investigated the effects of HPI throttling to 42 12.5 K subcooling and the other investigated the ef fects of no HPI throttling. The throttled HPI case was not run because the subcooling margin was never achieved.
800
- THETA = 1 - 572 i ' e THETA = 2 800- + THETA = 3
= THETA = 4
- 532 640-
- THETA = 6 m
E rurra -
- _ 4,2 t W
a:
I 85-800-
- 452 ,
W n.
3
- 412 %
( ,
l l
- 372 400-440 , , , , , , , , , 332 i
0 10 0 300 450 000 750 000 1060 1300 1360 1600
! Fig. 195.
Downcomer liquid temperatures (parametric Case 2) at vessel axial level 6 (all azimuthal sectors).
164 (
l TRAC calculated a relative minimum in temperature at approximately 1000 s into the transient; the pressure at 1000 s was -6.2 MPa. The temperature and j~ pressure were both decreasing at the end of the calculation at 3760 s and had
{ values of ~450 K and ~2.1 MPa, respectively.
Because the reactor should not have been tripped until the low-pressure reactor - trip setpoint was reached, and because the calculation was not run to 7200 s, it is recommended that the RELAP5 calculations be used for the ORNL study.
- b. Model Description and Assumptions. A complete description of the i primary-side, secondary-side, and ICS modeling can be found in Section II. The
. steady-state operating conditions are also presented in that section.
j The 2-in. break LOCA specification containing the initial conditions, I event sequence, and assumed failures is presented in Ref. 3. The TRAC transient event sequence is presented in Table XXVII. Because the ICS and major system f components were specified to function correctly, it was not necessary to make j any overriding assumptions.
a 1
l TABLE XXVII i HOT-LEG BREAK LOCA--2-IN. BREAK
! SEQUENCE OF EVENTS I
i Event Time (s) 1
- 1. Break opens 0.0
- 2. Reactor and turbine trips 0.5 I 0.5
- 5. HPI actuation 43+1
! 8. MFW realignment 73.1 i 9. TBVs open/close (both loops) 75.7 4
- 10. Vent valves open ~100
- 11. ICS closes SUFCVs ~350
- 12. Candy canes remain voided ~500
- 13. Minimum downconer temperature (470 K; 6.2 MPa) ~750 j 14. Loop oscillations begin ~1200
- 15. Accumulator injection begins ~1750
< 16. End of calculation 3670 1
i 165
- . - - . ~ __ _ - . _ . . . _ - . - _ _ - _ _ - = _ _ . - . - --
l l 1 l
Transient Calculation. - Figures 196
~
c.- and 197 the SG present secondary-side pressures for loops A and B. Following the initiation of the .
1 break, the reactor and turbine tripped from full power and the TSVs closed, causing a temporary increase in secondary-side pressure. The TBVs for both l- loops were activated ~4.2 s to relieve the initial secondary-side pressure i
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 main feedwater was realigned to the EFW header. 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 and ~187 s and increased the SG inventories. The ICS continued to 4
supply main feedwater to- the SG upper header until ~350 s. At this time, - the
, SUFCVs were closed by the ICS, based on the SG 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.
j The pressurizer pressure and water level are presented in Figs. 204 and
! 205, respectively. The primary system depressurized rapidly until HPI actuation at ^43 s and remained above ~6.0 MPa until ~1200 s. From 1200 s to the end of ;
the calculation, the pressure decreased steadily to ~2.1 MPa. The pressurizer
, e
- water level also dropped rapidly and was zero by ~50 s. Figures 206 and 207 I
present the break mass flow and void fraction. The break mass flow was greater *
- than ~100 kg/s for ~1000 s into the transient until the primary slowly voided.
f 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 HPI fluid began to flow toward the vessel and fell into the -
l downconer, as shown in the cold-leg mass flows and temperature profiles.
Figures 211 and 212 present the cold-leg liquid temperatures for loops A and B.
The 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 j temperature fluctuations ~ that occurred af ter ~1000 s will be discussed later.
j_ TRAC calculated a minimum cold-leg temperature of ~420 and ~440 K for loops A 166 1
A
so . -
use Go- -
meo n_ _
I. -
I so. _
~*
So- -'"
So . . . .
-3co o soo woo soo sooo mo mo mo sooo M (s)
Fig. 196.
SG secondary-side pressure-loop A.
l
-n so. _
l 4
j N" =
co- .
so.' .
~*
soo as- -~#
ao . . . . . .
3co o soo woo neo sooo anoo 3ooo me M (s)
Fig. 197.
SG secondary-side pressure--loop B.
167
l l
l ooooo ,
i l moose nooo- -
useos goeSo- -
}
-nosos g 4 seaso-p
-teoooo j hasooo- I -
E a .
osooo j asseo- -
l
- e. _
9 .
-doooo toooo . . .
o soo woo moo sooo asoo aooo anoo .ooo M (s) '
Fig. 198.
SG secondary-side inventory--loop A.
ococo nocoe foooo- .
=0000 Goooo-
- tmooo
~ ~"' I E,oooo.
esooo l
.oooo.- . soooo riooo aoooo- -
-ooose soooo .
. . . <enee o soo woo moo neoo aise aseo neo M (s)
Fig. 199.
SG secondary-side inventory--loop B.
=
168
000 , , , , -WBO
,.~ Loop A Loop B .'"
. -me 300- -
l .
I_ a. .
-- I 2 s00 2
- 3. .
go. [ .'N 0-- "- 4 7 .W ^2 -7M , f el -
e
.0 . , . . . . .
0 900 1000 900 3000 2600 3000 3000 4000 M (s)
Fig. 200.
MW mass flows--loops A and B.
1 300 , ,
i 940
- m. .
se0-Loop A .
E N ----- Loop B m-
\, 00 E y
s.
R*
l p
see_
~
j 8 sig . .. ... . .. , T . 400 8 '.~,
s00- '.,
..=
400-. -
400 . . . . . .
0 000 1000 See 3000 2000 3000 3000 4000 M (0)
Fig. 201.
MW liquid temperatures--loops A and B.
169
y I
se e-, 7 7
- o Loop A ~
2 ~
Loop B .
-ee
} }
h -m- ---so h
--so
.e. -
=00-. -
.go
\.' +
-.re
~*
o een eco eco mo mo acco asoc 4o00 M (s)
Fig. 202.
Emergency / realigned mass flows--loops A and B.
i see m
m- Loop A -
,, ) ----- Loop B .uo
- ~ ~
g srs s.o- E g
4 Sac- -
4 mo- --es 1 g
- o. ,, .
- m. -'"
o .
0 900 - =00 soo 2000 2000 3000 3600 6 Mfs)
Fig. 203.
Emergency / realigned liquid temperatures--loops A and B.
f 170 e.
,g- - - - . , , , - ~ - - , , , > ~ < - - - - - -
e , ,,,
mo , o.asasseo me- Loop A -
Loop B .
"' ~
[ mm l
h so- -
8 . .
.-.f ae-
- 8.coeMS So-o . . . . esoooooo O too woo eso tooo 2Soo Jooo J6eo dooo M (s) i Fig. 204.
Pressurizer pressure.
t I
6 s- .
.s a- .
s- - e E
., I
- t. .
e- J-- -A = -
o et . . . . . ,
e aeo woo eco sooo seco aooo anoo oo M (*)
Fig. 205.
Pressurizer water level.
171
200 ,
gyg . .
3- .
3 i--. ey',"9mngp:..i .
o 2S-q -
0- --@
0 500 1000 600 2000 2500 3000 3600 4000 NE(s)
Fig. 206.
Break mass flow.
u , , , , , , ,
- i. .
?
u- .
l l-- , %i 1 / I u-a w '
I e- .
e s s s s s s s m MW Fig. 207.
Break void fraction.
172
L' u , , ,
i 1
)
. u- LOOP A .
i ----- Loop B l
g o.e-l t
E i' :
g u- ,
. o2- 5 o-i .oJ , ,
o soo woo moo sooo noo seco anos 4eso M (e)
Ftp 'n8.
Candy-cane void frw .ons--loops A and B.
t l
seco l
Loop A *'
asoo. ----- Loop B .
seeo j asso- .
i seco j
h m- .
. so i Soo- .
4
.o
.- L gW#gg . . .
o e ese neo eso sees noe asse sees eseo M (e) l Fig. 209.
Cold-leg mass flows--loops Al and A2.
l 173
seso , , , , , , ,
,. Loop A '"
- w. ---- Loop B -
-seso
-eeso a m- .
._ i i _. .seen i
l ~~
.a .
e neo eso moo so o me asco anoo aseo ins (s)
Fig. 210.
Cold-leg maso flows--loops B1 and B2.
eco .
' eee gy,_, Loop A _
, ----- Loop B see see- -
E 2- E
-aeo
, /i g ese- t: g ., -4eo 1 es - '*
l ass-(
W -
ne
- see "o m m m m m m m m ins (s)
Fig. 211.
Cold-leg liquid temperatures--loops Al and A2.
174
- n. n .-m . .,e-- c- . , -,~w ----r. --,-.--.n- -
ese , ,
mm -m one-mn -
-seo see. -
E E
\g .m m- - -
leso.
3 J
I me
.as g g age.
{
g
. 5 see es0- \t ,
seo 4,o. -
-se "o m n n m m m m e M (s)
Fig. 212.
Cold-leg liquid temperatures--loops B1 and B2.
wooo .
(seed) LOOP A
- (p snoop e eseo- -
yges see,. ' * *
.seos m- -
! I m.
~~
l seso
^
9- -' -
--9 e see o . seee am aeos me esse MM j
Fig. 213.
Hot-leg mass flows--loops A and B.
175
4 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 t,he vent valves located j around the upper plenum of the vessel (level 7 in the TRl.C model), which provide the upper plenum region access to the downcomer. Af ter the RCPs have tripped,
- the vent valves are capable of providing a source of hot fluid for mixing with cold HP1 fluid that'may flow toward the downcomer during these stagnant periods I in the cold legs. Between ~200 and ~1000 s of this calculation, HPI fluid did flow toward the downcomer and was mixed with the vent-valve flow in the downconer at the cold-leg junction.
Downcomer liquid temperatures at the top axial downcomer level (just below the cold-leg no.:zles) are presented in Fig. 215. TRAC calculated a minimum j downcomer 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 HP1 flow as the system pressure decreased and accumulator flow once
! again decreased the liquid temperatures. At the end of the calculation, the j downco1ter temperature was approximately 450 K. -
From the results of a previous calculation in which the vent valves were accidently isolated from the downcomer (input error), the importance of the vent j valves for this particular transient was determined. Figure 216 presents an azimuthal comparison of selected downcomer liquid temperatures at axial level 6 for both calculations. When the vent valves were modeled properly, TRAC j calculated downcomer liquid temperatures that were at least 25 K warmer. From j the PTS viewpoint, the vent valves were an asset in maintaininE " warmer" f downcomer 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.
l 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 l negative mass flow (away from the vessel) in loop B2. Between ~1200 and
~1250 s, a similar occurrence happened in loop A. Cool HPI fluid from the loop Al Sold leg dropped into the loop seal in loop A. Immediately following this i 176
ooo , ,
Loop A .soe
,,,_ ----- Loop B .
. s'o g *** ~
.a E seo- -
se m- - ~'**
I too- -
.. I eto ato-mo o Soo loco too tooo 2Soo Sooo Moo dooo NC (s)
Fig. 214.
Hot-leg liquid temperatures--loops A and B.
eco 8t TH Z soo- 216 -
226 .q
- ~ 236 -
g -
246 m
- ~ 256 -
m-266 .
-=
- e
'I .a goo _
j ""
- o. .
Mi eso -
.w 44o-39 40o . . . . .
o soo woo noe mo me mo mee mo TWC (s)
Fig. 215.
Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors).
i 1
177
s,. . .
,q
~
wov.v.isoc caoug ~
- . k '
,, ? ROPER V.V. les0G" $ hts $ m.
! E '.
E
~
) !. ,J
. %,'1r. .
- g. i ,i'd.
fs .I2 ty i1 gb.. g ag i
m 11
,Lg -
l . .m.
i
- s I W .
m
- . = . . .
ThE(4 Fig. 216.
Downcomer liquid temperature comparison for 2-in. break case (vent valves vs no
- vent valves).
u . . . . . .
I E" "
l 5 .. ,
1 .. .
-. t
< I -
.. J -
N Fig. 217.
Total vent-valve mass flow.
178
__ _ _ _ ~ . _ _
l occurrence, the loop-A and -B mass flows and liquid temperatures came in phase
, and began oscillating. Initially, the loop-B oscillations were regular at ~17 s and an amplitude of 400 kg/s. The loop-A oscillations were not initially I
regular.
l
- To explain the loop oscillations, several case studies were conducted.
These cases included the following: turn the HPI of f before the oscillations i begin, turn the core power off with the SG heat transfer on, turn the SG heat 1
transfer off with the core power on, close the break, and renode the j middle-to-upper levels of the vessel and steam generators. The oscillations presented in this report were calculated using the renoded vessel and the steam generators.
l Briefly, the results of the parametric case studies indicated that the HPI and the secondary-to primary heat transfer in the steam generators were the j
4 forcing functions that caused the oscillations to persist once initiated. These were manometer-type oscillations and the columns of liquid that oscillated included the legs of the loop seals, lower half of the steam generators, and the i 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 up the lower half of
) the steam generator. Because of reverse SG heat transfer (secondary to primary), this additional liquid was heated, thus changing the effective elevation head in the steam generator. Thus, because of these changing elevation heads in the loop seals and steam generators, the oscillations i persisted.
1 l e. Summary. The Oconee-1 plant response to a 2-in. break in the surge line was calculated using TRAC-PF1. 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 minimum vessel downcomer itquid temperature of 470 K at 1000 s.
The primary system pressure at this minimum liquid temperature was calculated to be 4.2 MPa. HPI flow and accumulator injection reduced the temperature at the end of the calculation at 3670 s to -450 K.
The calculated' minimum downcomer liquid temperatures never approached the current NDT value of Oconee-1 for two reasons: (1) vent-valve flow mixing with the fluid in the downcomer region and (2) calculated loop oscillations.
i However, if the calculation were continued, the LPIS will actuate as a result of 179
.__ _ _ . _ _. . . . . . . . . _ _ _ . _ _ . _ _ . _ _ . = ,
, the- depressurization. The addition of low-pressure injection (LPI) would i probably result in downconer liquid temperatures that would 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 downcomer liquid temperature falls below the NDT limit.
- 2. Four-Inch Break f a. Introduction and Summary. This report presents the Oconee-1 plant j response to a 4-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 and turbine tripped f rom 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 that the RCPs trip - 30 s after HPI actuation. Two cases involving HPI throttling to j system subcooling were to be investigated for this transient. At the end of the t base-case calculation (~1433 s), the subcooling condition had not been achieved; j therefore, only one calculation was required.
f TRAC calculated a minimum vessel downcomer liquid temperature of ~350 K; i
i the primary system pressure at this minimum temperature was ~1.0 MPa.
- b. Model Description and Assumptions. A complete description of the primary-side, secondary-side, and ICS modeling can be found in Sec. II.. The j steady-state operating conditions are also presented in that section.
f j The 4-in. break specification containing the initial conditions, event ,
k
- 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 not l necessary to make any overriding assumptions. !
I
- c. Transient Calculation. Following the initiation of the break, the l reactor and turbine tripped from full power and the TSVs closed, causing an increase in secondary-side pressure. Between ~17 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 steam generators remained momentarily constant just below 4.0 MPa after the SUFCVs . closed
! because of the increasing secondary-side inventories. After ~400 s, the loop-A I secondary-side pressure decreased at a much faster rate as the primary cooled .
the secondary. This heat-transfer mechanism lowered the loop-A secondary 180
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. TSVs 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. HPI 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/close 92.1
- 14. Candy canes void 125.0
- 15. ICS closes SUFCVs 300.0
- 16. Accumulator injection loop A (first time) 540.7
- 17. Accumulator injection loop B (first time) 541.0
- 18. Accumulator injections (both loops) 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. LPIS actuation on low primary system pressure ~1236.0
- 28. End of calculation ~1400.0 pressure and produced a larger loop-A secondary-side inventory as a result of condensation. Figures 220 and 221 show the loop-A and -B SG secondary-side inventories.
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 main feedwata 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 181
. ~ . - ._ - . - - _ .
l
-noo oo. .
70- -
-Woo to-
-a00 Q so- - 8
-Soo
- m. -
30- _,
20-
- -200
- g. .
o , , , . O o too 400 soo soo e coo woo itoo TIME (s)
Fig. 218.
Steam generator secondary-side pressure--loop A.
as , , , , , , ,
- -noo 1
en. .
- '*o
- n. .
70- -_,
I --
go.
..oo I ss- - - -800 go. .
. -poo
- e. .
-Soo do- -
35 . , , , , . ,
n zoo 4o0 soo soo coo eco woo moo TIWC (s)
Fig. 219.
SG secondary-side pressure--loop B. i 1
182
l ao000 . , , . . .
95000- -
.goooo 30000- -
-moooo 4 000- -
k anooo. -
-socco g l3e000- E
.m a i
"- ~
00 25000- -
20000- .-4S000
' ~
-30000 10000 . .
0 200 400 600 800 1000 0 00 WOD 8600 TlWE (s)
Fig. 220 SG secondary-side ir.ventory-loop A.
20000 . , , . .
-105000 goog. .
40000-~ -
$ "~_ ~
_75ao, l s g Y soo00- -
B E - Doo 5 1 ,, . .
3 20000-
- '# 000 sooo- -
-30000 10000 , , ,
0 700 400 600 800 1000 000 woo 9600 TIME (s)
Fig. 221.
SG secondary-side inventory-loop B.
f 183 i
400 , , , , , , , -1750 700- Loop A ,
Loop B - 500 eno. -
-1250 m s00- -
-1000 d,
- 750 Y $
2 -500 2 go . -
100- / f
-}
1 0- * -My :,
-O
- 10 0 .
0 200 400 600 800 1000 120 0 MSO 1600 TWE (s)
Fig. 222.
MFW flows--loops A and B.
540 , ,
-500 s30- _,.., \..........,,,
. -no s20- -
E Loop A F v
siO ~
~_ ,o
Loop B 300- -- 440
- 0-_ -- 420 $
8 4s0- -
470 - -
- 380 460 0 200 400 600 800 1000 1200 u0o 1600 TWE (s)
Fig. 223.
MFW liquid temperatures--loops A and B.
l 184
zo . . - ' ' '
-m s
T '
~ ~
1
. --m 3o. -
..a Loop A '
-do- -----
Loop B
..uo 4 .
---as
--29
.wo . -
--ze I
-no .
a 200 doo 600 800 Woo Cao Woo 400 1WE(s)
Fig. 224.
Realigned mass flows--loops A and B.
soo , . . . . .
-Soo
- Loop A
Loop B ,
.soo
.I
$ goo. -
.... ~ . . ~ - . . . . . . . . . . . . . . . . . .. . . .
{
-noo l- .
..oo 400- -
4 ..oo me. .
P.
-mo 3oo . , . ,
o zoo 400 soo soo woo woo woo woo TIME (s)
Fig. 225.
Realigned liquid temperatures--loops A and B.
185
temperatures. The EFW pump was not activated in this transient because it was not needed.
Pressurizer pressure and water level are shown in Figs. 226 and 221, respectively. The primary system pressure fell very rapidly until the HPI was actuated at ~17 s. The system pressure was almost. stabilized after ~800 s at
~2.0 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.
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 zero following the RCPs trip. The loop-B hot leg flow became stagnant as the candy cane voided; however, the loop-A hot leg did not stagnate and continued to feed the break with vapor. Hot-leg liquid temperatures decreased in accordance with the primary system depressurization. TRAC calculated a minimum hot-leg temperature of ~455 K. The loop -A and -B cold-leg mass flows (Figs. 233 and 234) exhibited trends similar to those of the hot legs. The corresponding cold-leg liquid temperatures are shown in Figs. 235 and 236. Minimum cold-leg temperatures (before LPI actuation) were calculated to be
~430 K at ~650 a 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), HPI 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 corresponding 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 LPIS at
~1240 s on low primary-system pressure dropped the cold-leg liquid temperatures near or below -400 K. The LPIS injected directly into the vessel downcomer in axial level 7.
186
- - ~ -
90 , , , , , , ,
2100 te0- .
l
~
l 90- -
i I ggg.
30-
~
'E
-1200 I
n- 00- - . -900 40 - - -800 20- - -
300 0 , .
, , , , 0 0 200 400 600 800 1000 C00 WOO 1600 TWE (s)
Fig. 226.
Pressurizer pressure.
S , , , , , , ,
5- -
-5 4- .
4 s- . - iO l
2- 1 -
g
-s 3 1- .
i,
' A-0- --^ - ^ - A -
0 4 -1 . . , .
0 200 400 800 000 1000 t?00 M00 8800
, IEE (S)
Fig. 227.
, Pressurizer water level.
187
800 , , , , , , . -f50 M~
-1500 goo. .
-1250 i l
m- -
2 m. .
k
, "M M-Q
- m. .
N $ i p 0- --O
-100 .
0 200 430 600 800 1000 1200 u00 1600 TIME (s)
Fig. 228.
Break mass flow.
u , , , , , , ,
j gg. .
l3 es-s.t - ,
i
- e. .
e s e a s m m mm TIME W Fig. 229.
Break void fraction.
188
t2 . . . . . .
l l 1- ,
jgr -
0.e - Loop A -
Loop B E
0.4 - ' -
0.2 - -
0- -
-0.2 0 200 400 600 803 1000 1200 WOO 400
- WE (s)
Fig. 230.
Candy-cane void f ractions--loops A and B.
10000 . . . . . .
-21000 Loop A
> 9 Loop B
- a000- -
-risco
-uc00 0000- -
._0 I m- -
g -
- m00 g 5 3 anon. -
-3600 ,
o- - .......... .. . . . .. T e # * --0
--200
-2000 . , , .
0 200 400 Goo 800 1000 1700 WOO 500 WE (s)
Fig. 231.
Hot-leg mass flows--loops A and B.
189
soo . . . . . .
~
Loop A '"
soo~
Loop B
- sro
-540 s40- -
-se m-- - -4ao
-co Soo- -
- 420 4eo- -
-390
~ '
l .
3.o 44o . .
o 200 400 Goo A00 ?c00 P00 54 0 o 8600 TIME (s)
Fig. 232.
Hot-leg liquid temperatures--loops A and B.
SooO , . .
Loop A1 -80000 Loop A2
} _
-sooo q M- -
,- -sooo R
"' g b
~
sn
-4000 c 3
moo- -
-2o00 o-- %'. - pp D ; --o
-moo . , . , .
-2m 0 20o 400 600 soo ".000 1200 1800 1600 Tiut (s)
Fig. 233.
Cold-leg mass flows-loops Al and A2.
190
.000 . . . . . . .
-10000 l Loop B1 4000- ) Loop B2 -
-s000
- m. -
8000 2000- -
-4000 1000-_ ~-2000 h
2 0- _
. --0 9
t .--2000
-4000
-2000 . .
0 200 400 600 800 1000 1200 M00 1600 TIME (s)
Fig. 234.
Cold-leg mass flows--loops B1 and B2.
I 800 . , . . . . .
-600 373 Loop Al _
Loop A2 _,
350 .
. -500 g 525- ,,
A -
g 00 .. .
4 L :.
4 I Nl .-400 c3 460-
,k,
- [h h[;.Ii I .
Ih?[A- *
--350 425-. **
--300 3 O. ~
_ 230 Sys. -
-200 350 . .
0 200 400 600 800 1000 1200 1400 1600
- TIME (s)
Fig. 235. i Cold-leg liquid temperatures--loops Al and A2. I 191
eco , , - ' ' '
_ goo ps. Loop B1 -
- Loop B2 _33o sea. -
-500 g m- - -
E
-O ly."
475 -
I :. .
m- -- 350 o
. 5 425- -
300
~-250 Sys . -
-200 350 .
0 200 400 600 800 1000 000 100 12.1 0 TIME (s)
Fig. 236.
Cold-leg liquid temperatures--loops B1 and B2.
ses , , . . . . ,
see. .
ese- Ys -
\
$ .. \ _
1 \
l -~
\' i y!
J
- e. )
~J)L JA %)s.
e b o 5 5 5 5 m M
Fig. 237.
Total positive vent-valve vapor mass flow.
192
As identified in Sec. I, transients involving system repressurization and )
overcooling have been identified as events that could damage and possibly cause ;
the failure of a PWR vessel. Thus, the key PTS parameters are pressure and liquid temperature in the vessel downcomer region around the weld locations.
For this particular transient, the primary concern is thermal shock because the primary system will not repressurize. In this plant, the welds located in vessel level 6 in the TRAC model are the weld locations that are considered important for this study. TRAC calculated a minimum downcomer liquid temperature of ~350 K in level 6. The system pressure at this minimum downcomer liquid temperature was calculated to be ~1.0 MPa. The minimum temperature values calculated for downcomer level 6 (between ~540 s and ~1200 s) directly correspond to accumulator injections. The LPIS actuation produces the minimum temperatures after ~1200 s. Downcomer liquid temperatures at the top axial downcomer level (just below 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 after HPI actuation. TRAC calculated a minimum downcomer liquid temperature of ~350 K; the system pressure at this minimum temperature was ~1.0 MPa.
Adequate fluid mixing between the vent-valve fluid and the cold-leg (HPI) fluid in the downcomer at the cold-leg junction maintained the downcomer liquid temperatures above ~450 K. However, the actuation of the LPIS at ~1240 s dropped the downcomer temperatures very rapidly and below the current NDT value
(~365 K) for the Oconee-1 plant. Even though the downcomer 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 (SG 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, SG dryout followed by EFW overfeed, has been analyzed. The accident sequence began as a loss of MFW transient (MFW pumps trip). The EFW control valves failed to open on demand but were subsequently manually opened by the operator. Also, the RCPs l remained on during the transient, and the emergency feedwater to the steam 193
000 . + '
-606
\\
SSo-
~
-506
.. o 8 soo- _
4 ~
l -
s mz 216 c
r
) 30, soo_
o226 l _
236 9 4
_ +246 hp]' 200
- m. 256 F _
266 300 30c , . .
0 200 400 600 800 1000 C00 1400 1600 TNE (s)
Fig. 238.
Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors).
Senerators was not terminated until 4200 s. The primary system repressurization was limited to ~13.8 MPa as a result of the operator throttling the llPI system.
TRAC calculated a minimum downcomer liquid temperature of ~452 K at 4200 s. Repressutization of the primary system to ~13.8 MPa was also calculated.
- 2. Model Description and Assumptions. A complete description of the primary system, secondary system, and ICS modeling can be found in Sec. II. The steady-state operating conditions are presented in that section also. The SG dryout followed by EFW overfeed accident specifications containing the assumed plant initial conditions and postulated event sequence are presented in Table XXIX. The calculated event sequence is presented in Table XXX. To ensure the correct plant response to the postulated sequence of events, significant portions of the TRAC ICS model were used.
- 3. Results. After the loss of main feedwater and coincident turbine trip, the reactor tripped from full power on high RCS pressure at ~4.4 s. The PORV functioned properly between ~226 and ~550 s and relieved the pressure increase in the primary system while the steam generators dried out. The 194 1
. . = ~ -
TABLE XXIX I
l 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 fail close
- 6. SG dryout
- 7. PORV (primary) function as designed
- 8. Turbine bypass system operates as designed
- 9. Operator fully _ opens EFW control valves at 9 min
- 10. HPI activates on low pressure
- 11. Accumulator and LPI function as designeda
- 12. Operator fully opens EFW control valves at 9 min
-13. Operator. limits pressurization to 13.8 MPa by throttling HPI flow
- 14. EFW flow terminated at 70 min
- 16. Operator throttles HPI to maintain 42 K subcooling after steam generators are restored to proper level aMay be phenomenologically dependent.
Primary-system pressure, pressurizer 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 loops A and B, respectively. At 540 s, the EFW valves were opened fully and emergency feedwater began to refill the steam generators. EFW mass flow rates and liquid temperatures at the point of injection are presented in Figs. 244 and 245, respectively. The SG inventories !
for loops - A and B are shown in Figs. 246 and 247, respectively. As the 195
(
- - -- -. - - - - - .- = _.
['
i TABLE XXX RA!MHO SECO-TYPE TRANSIENT SEQUENCE OF EVENTS l
l l Event Time (s)
- 1. ~MFW pumps trip, MFCVs and SUFCVs close 0.0
- 2. TSVs close (both loops) 0.0 i 3. Reactor trips on high pressure 4.4 i 4. TBVs actuated ~4.6
- j. 5. PORV actuated 226.4
- 6. Emergency feedwater initiated 540.0 to both steam generators
- 7. HPI actuated on low pressure ~738. 0
- 8. HPI throttled to limit repressurization ~1255.0
- 9. Emergency feedwater terminated 4200.0 to both steam generators
- 10. . Minimum vessel downcomer liquid temperature (~452 K) calculated ~4200.0
- 11. Calculation terminated 4300.0 emergency feedwater was injected into the steam generators,- the secondary-side inventories began to recover and the primary system began to depressurize.
, Cold-leg liquid temperatures for loops A and B are presented in Figs. 248 and 249, and Fig. 250 presents the hot-leg liquid temperatures for both loops.
1 As a result of the depressurization, the HPI system was actuated at ~738 s on low pressure. At ~1255 s, the - HPI system flow was throttled to limit RCS repressurization to ~13.8 MPa, as specified in the . event sequence. The HPI, system flow was continually throttled to limit system repressurization throughout the transient, as indicated by the loop-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 B are presented in Figs. 253 and 254, respectively. Figure 255 presents the hot-leg mass flows for both loops.
The continued . operation of the RCPs provided forced convective heat-transfer on the primary side, which assisted in the rapid cool down of . the j primary system, as indicated by the cold- and hot-leg liquid temperature j
l profiles between ~540 and ~3500 s.
196
50 . . . . , . . . atoo m _
zaoo So-1- -
=~ .
I l no- ~
mon go. -
~
. woo oc - . . - -
c 800 e0' 0 500 2o00 2$00 3000 3600 4000 4 30 M (5)
Fig. 239.
Primary system pressure.
m , , , , ,
e- -
.. - =
s- -
m I ,. ,
- E b .. - = b s- -
i s
4 -
a- - #
3 =
s s , , . . . .
e soo mac Soo sooo sooo anco anoo sooo eso MW Fig. 240.
Pressurizer water level.
197
85 . . . . , , . ,
, m s- .
30 as - .
m e- .
3o 7.s - . $
.e e E s- .
.m 2.s - ~ .
e- -,
. ., 1
-2.5 . , , , , , ,
--s e 500 2 00 too foco 2500 E00 Jeoo 4000 4 00 TWE (s)
Fig. 241.
PORV mass flow.
so , , , , ,
75- -" '80 i M. , h" moo m- .
~
so. .
..oo y
so- -
mo f '
- s. .
.soo
- e. .
as. .
-nao ao , , . , , .
o soo moo soo sooo seco aoso asoo sooo e,mo
, INC (4 Fig. 242.
SG A secondary-side pressure.
! 198
1 l
I ac , , , , , , .
l nso l z.
M- V "
,ggg m.
I ao-soo T9
- -soc m.
ac-
-No a
-soo 4
M , , . . ,
o soo woo soo sooo 2s00 3000 moo 4000 moo M (s)
Fig. 243.
SC B secondary-side pressure.
- o , , , , , ,
- -gg Loop A g,
-- Loop B !i -83 e asnow so- *
! i -
i
, . , . . , , , . . . - .. . . . . . . . * *i *
- so
- 88GD*
j i m. .
4o- - so 4 =
g W u,-
so- -
-se e< W -.
o i
- -.3, i
-n , , . . , ,
o too eco soo sooo 2500 3000 Boo 4000 e00 M (4 )
l Fig. 244.
EFW mass flows.
199 l
l
eso , , , , , , , ,
l Loop A esos ----* ----- Loop B .
I g m- .
m- .
I me - .
I me- .
r-
- m. _ : : : : : : ,
no ,
o saa woo sea 200o soo sooo moc 4eco coo TIME (s)
Fig. 245.
EFW liquid temperatures at injection point.
o0000 . , , , , , ,
I soooo- .
m i mono- .
Goooo
< escoo- . ,
! messo >-
m- .
soooo gj anooo- -
I h nooo @
E ~- -
g scono soooo. .
mooo. , . mooo e- . .o
-moos , ,
o son moo soo 2000 noo 2000 moo sooo moo indt (s)
Fig. 246.
SG A secondary-side inventory.
200
socco , ,
j soooo- -
psooo 7900o- -
sooon.. -
m usooo E soooo-2 wooon $
I 4acoo- -
i 5 noo. g j scooo. -
g
.scono soooo- .
moeo. ; - 2000 e- - o
-900m . , ,
o soo too too sooo asco sooo asco 4o00 ano M (s)
Fig. 247.
SG B secondary-side inventory.
- no , , , , ,
Loop Al "
- D- Loop A2 -
sto aso- -
E E sso- -
se me. - ***
..o soo- -
-ee a me. -
soo aso- -
Ado .
O soo loco too Sooo 2soo 3000 3600 4cCo eco M (s)
Fig. 248.
Loop-A cold-leg liquid temperatures.
201 l
l
o00 , , , , , , , ,
Loop B1
- ' Loop B2
~
spo aso- -
sao 8 E g
g me- . 400 e3 oco- -
Go a 400- -
-soo aso. .
aso 440 . . ,
o 800 9 00 500 2000 2500 3000 . MOD 4000 4 00 TWI(s)
Fig. 249.
Loop-B cold-leg liquid temperatures.
800 , , , , ,
Loop A 800 as0- ----- Loop B -
s70 see. .
340 sac- -
3, go.
\ '
. as0
'g .o soo- N, -
me-
\ ' -
eo s
. ano m- -
~~. -
44o , , , , , . .
o soo moo eso sooo we 3000 2 00 acco moo ThC (a)
Fig. 250.
l Hot-leg liquid temperatures.
l 202
l I
F.s . . . . . . . .
Loop Al as s- - Loop A2 -
e enow ao n.s- '" - a wupw as 3 m. . g 4 -ao E fU y.s. :5
\, !.$ h',
).
.s d ,
O
$ s-h h, .
j
.\ \ ,
s.s- l -
.s 31' l l ec : n' LE L-+- U >-- d - -
e
. .s 1 , ,
-s' o soo moo sco roco asco sooo asco 4 00 moo M 6)
Fig. 251.
Loop-A HPI mass flows.
ts . . . . . . .
Loop B1 8' "
Loop B2 n *m e-
- a wupw
- I mo e- -
s
- s. .' i d $
e'. r'
.. 1- e k k '-
m N I d-
- j l!, !j.
il i 5
i' j 8
.; ML, Ad i ___._ .
L L .,__ ,. _ _ . .o
-2 . . . . , , ,
o soo eco soo sooo asco acoo asco moo moo M 4)
Fig. 252.
Loop-B HPI mass flows.
203 1
Seco . . , , , . . .
Loop Al teco 8 80 ==
Loop A2 -
veco 3000- -
esoo 48o0-4 seco E T
400o- -
sooo
{
3 seco- -
N 5
eco-
-anoo moco- -
asco mo: . , ,
'W o soo 10 0o nos 2000 noo 2 00 2 00 4000 coo M (s)
Fig. 253.
Loop-A cold-leg mass flows.
seco . . , , , , . .
Loop B1 -- - **c o 8'" ' -
Loop B2 ~
toco 3000- -
msoo asoo- -
O Seco b m. $
d .
d mooo N N l $ deco- -
I esoo moo-
-esoa 400o- -
Stoo 3000 , , . . . . .
'M o Soo 1000 900 3Doo 2600 3000 380o 4000 aoo M (s)
Fig. 254.
Loop-B cold-leg mass flows.
204
At ~3500 s into the transient, the loop A secondary side (steam generator and steam lines) had been completely filled with emergency feedwater 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 emergency feedwater continued to feed the system. As a result of the secondary-side repressurization, the primary side began to cool at a slower rate. At 4200 s, the emergency feedwater was terminated to both steam generators (as l specified). At this point in the transient TRAC calculated a minimum vessel downcomer liquid temperature of ~452 K. The system pressure at this calculated minimum downcomer liquid temperature was ~13.8 MPa. Vessel downcomer liquid temperature profiles for all six azimuthal sectors at axial level 6 (at the weld locations) are shown in Fig. 256.
- 4. Conclusions. The overcooling of the primary side of the Oconee-1 Pl ant caused by a SG dryout followed by EFW overfeed (Rancho Seco-type transient) was simulated with TRAC-PFl. The TRAC simulation calculated most of
^
the plant response and occurrences outlined in the postulated sequence of events. A minimum vessel downcomer fluid temperature of -452 K was calculated
- at 4200 s. Repressurization of the primary system to ~13.8 MPa was also I
calculated.
IV. CONCLUSIONS AND RECOMMENDATIONS The response of the Oconee-1 plant for several overcooling transients has been predicted using TRAC-PF1. The complete plant, including the primary and secondary sides, was modeled so that accurate predictions of system thermal-hydraulic conditions could be made. The plant control and protection systems were also modeled in sufficient detail to siiulate actual plant response during these postulated overcooling transients. The results of these calculations are to be used for PTS analyses at OPNL.
Several overcooling transients were analyzed. The transients calculated included a MSLB with a delay in isolating the affected steam generator, a small-break PORV LOCA with failure of the ICS to throttle MFW flow and RCP trip, and TBV transients with SG overfeed. An actual plant trcnsient (Oconee-3 turbine trip) was also simulated by TRAC to compare with actual plant data. Two small hot-leg breaks were also analyzed to investigate the effects of vent-valve flows 205
socc , , , . - - - '
asooo Loop A gooo. - Loop B -
.....- l sooo- -
-sanoo
....,..~
,,- , .seco n . .-o E
$ sooo- -
I '
mooo seco- ! -
I
- . moon oooo.
nooo m:
m o soo moo soo sooo asco Neo moo eco M (s)
Fig. 255.
Hot-leg mass flows.
803 . . , , , ,
R TH 2 soe aso-216 -
226 '"
asoJ 236 -
ses 8 246 g s o- . \ 256 -
se
- m. - 84 g soo- \ -
.a me - -
seg I
, no- -
- ass l
MD , , , , , , ,
o soo moo 300 sooo asco acoo inoo eco moo M (s)
Fig. 256.
Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors).
206
l l
on downcomer fluid mixing. Finally, a Rancho Seco-type transient was j 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 downcomer. The most severe transient in terms of overcooling and system repressurization was the TBV transient.. 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 mininum 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 should 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 runaway MFW flow and all other plant systems operating would possibly lead to a more severe overcooling transient.
ACKNOWLEDCMENTS The authors wish to acknowledge the extraordinary efforts of word processors Jean Martinez 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.
t i
207 1
l 1
l REFERENCES l
- 1. Safety Code Development Group, " TRAC-PFl: An Advanced Best-Estimate Computer Program for Pressurized Water Reactor Analysis," Los Alamos
, National Laboratory report (to be published).
- 2. R. C. Kryter, R. D. Cheverton, S. D. K. Kam, T. J. Burns, R. A. Hedrick, and C. W. May, " Evaluation of Pressurized Thermal Shock," Oak Ridge National Laboratory report ORNL TM-8072, NUREG/CR-2083 (October 1981).
- 3. J.' D. White, " List of Oconee-1 Transients for Thermal-Hydraulic Calculations," Oak Ridge National Laboratory letter (December 1982).
1 4. " Transient Assessment Report for Oconee Nuclear Station Unit III Reactor Trip of March 14, 1980," Duke Power Company report (No date).
- 5. C. D. Fletcher, "RELAP 5 Thermal-Hydraulic Analysis of Pressurized Thermal Shock Sequences for the Oconee-1 Pressurized Water Reactor,"
Idaho National Engineering Laboratory report ECC-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 document 4 LA-UR-82-3294 (November 1982).
t 208 l
2
i APPENDIX A l
OCONEE ICS CONTROLLER FOR LOOP A
, (All signals input are in SI units.) l
'(Initialization of valves is for steady state only.) l
'(Letters indicate boxes in the previous figure.)
BTU LIMITER 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 + 0.037260
- FWTEMP , -1.270 < D < 9.080 E = -16.0 + B + C + D , -10.0 < E < 12.0 F = 0.55555 + 0.055555
- E H = -10.0 + A
- F
- hold initial value for H until 10 s have passed h IF(TIME'.LT. 10.0) H = 8.0
" TOP" 0F LAYOUT
- -20%/ min ramp after trip A1 = TIME - (TIME OF REACTOR TRIP)
B1 = 1.0 -L(0.2/60.0)
- A1 , 0.0 < B1
- feedwater demand function
, C1 = 18.0
- B1 D1 = f(C1) : C1 D1 0.0 204.0 0.562 240.0 3.6 320.0 5.4 356.0 9.36 402.0 18.0 460.0 21.42 483.0 209
- feedwater temperature compensation El = -460.0 - D1 + 1.8
- FWTEMP i_
F1 = 1.0 + 0.0013
- El G1 = F1
- C1
- hold initial value of G1 for 10 s
,IF(TIME .LT. 10.0) G1 = 18.0 T-
- neutron power cross limiter
- initialize signal IF(TIME .LT. 10.0) POWER = 2568.0E6
- bias signal back to zero SP = POWER - 2568.0E6
- first order lag of power with 4.5 s time constant JL = JL + ((SP - JL)/4.5)
- DELT
- remove bias J1 = JL - 2568.0E6
, H1 = 1.6 + 14.4
- B1 Il = -1.0 * (H1 - 6.23053E-9
- J1) , -10.0 < Il < 10.0 Kl= f(II)' : 11 K1
-10.5 -10.0
- 0.5 0.0 0.5 0.0 10.5 10.0
- initialize signal IF(TIME .LT. 10.0)'FWFLOWA = 680.4
- bias signal to zero 210
\
i FWB = FWFLOWA - 680.4
- first-order lag of f.w. flow error with 1.0-s time constant, loop A ,
FSL = FSL + ((FWB - FSL)/1.0)
- DELT l l
- remove bias SL = FSL + 680.4 S1 = 10.0 + R - 0.026455
- high-level limiter, loop A
- operating level scale, 96 to 388 in. (level in meters)
HL1 = f(ALEV) : ALEV HL1 2.438 -10.0 9.855 10.0 P1 = -2.0 * (HL1 - 7.0)
- take the smallest between signals Si and P1 Q1 = min (S1,P1)
- low-level limiter, loop A
- startup level scale, 0.0 to 250 in.
LL1 = f(ALEV) : ALEV LL1 0.0 -10.0 6.350 10.0
- decide which setpoint to use depending on pump trip
- pumps running : .24 in. = 0.61 m = -8.08 V 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,Ql) 211
" BOTTOM END" CF FLOW CONTROL
- choose which constants to use depending on whether the S3 is low-level limited or not
- if P2 < 0, low limit has not been hit IF(P2 .CE. 0.0).CNST1 = 0.12
.IF(P2 .LT. 0.0) CNST1 = 0.1125 IF(P2 .CE. 0.0) CNST2 = 2.4 IF(P2 .LT. 0.0) CNST2 = 0.9
- integrate, T1o is the last time-step value of signal T1 U1 = U1 + CNSTI * (T1 + Tio )/2.0
- DELT , -18.0 < U < 2.0 X11 = U1 + CNST2
- T1 X1 = X11 + 8.0 , -10.0 < X1 < 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 a 0.1
- SUA , 0.0 < SUFVA < 1.0 j
- Main flow control valve function, loop A 4
MFA = 0.5555
- X1 - 4.4444 , -10.0 < MFA < 10.0
- normalized flow area for MFCV--loop A
- this signal sent to valve MFCVA = 0.5 + 0.5
- MFA , 0.0 < MFCVA < 1.0 i
l l
[ . 212 l
i
Oconee ICS Controller for Loop B (Balance of signals comes from loop A sections.)
BTU LIMITER BA = 0.00204083
- RCFLOUB i
BB = -605.4459 + 1.04092
- RCTEMPB , -10.0 < BB < 9.080 BC = 82.549 - (1.16958e-05) *-SGPRESB , -1.0 < BC < 9.080 BE = -16.0 + BB + BC + D , -10.0 < BE < 12.0 BF = 0.55555 +.0.055555
- BE BH = ~10.0 + BA
- BF
- hold value of BH at 8.0 until 10 s pass IF(TIME .LT. 10.0) BH = 8.0
" TOP" 0F LAYOUT SECTION
- initialize signal 1
IF(TIME .LT. 10.0) FWFLOWB = 680.4
- bias signal to zero i FWC = FWFLOWB - 680.4
- DELT
- remove bias BSL = FBL + 680.4
'BS1 = 10.0 + BR - 0.026455
- BSL l
I t
213 l
, , = . - - . . . -
SG OPERATING LEVEL LIMITERS l
- high-level limiter, loop B
- operating level scale, 96 to 388 in. (level in meters)
BHL1 = f(BLEV) : BLEV BHL1 9.855 10.0 2.438 -10.0
'BP1 = -2.0 * (BHL1 - 7.0)
- take the smallest between signals BS1 and BP1 BQ1 = min (BS1,BP1)
- low-level limiter, loop B
- startup level scale, 0.0 to 250 in.
BLLI = f(BLEV) : BLEV BLL1 6.350 10.0 0.0 -10.0
- low-level error function BP2 = -2.0 * (BLL1 - STP)
- take - the largest signal - BP2, BQ1 BT1 = max (BP2,BQ1)
" BOTTOM END" 0F FLOW CONTROL
- choose which constants to use depending on whether SG B
- is low-level' limited or not
- if BP2'< 0, low limit has not been hit IF(BP2 .GE. 0.0) BCNSTL = 0.12 IF(BP2 .LT. 0.0) BCNST1 = 0.1125 IF(BP2 .GE. 0.0) BCNST2 = 2.4 IF(BP2 .LT. 0.0) BCNST2 = 0.9
- Integrate, BT1o is the last time-step value of signal BT1 BU1 = BU1 + BCNST1 * (BT1 + BT1 o )/2.0
- DELT , -18.0 < BU1 < 2.0 BX11 = BU1 + BCNST2
- BT1
.BX1 = BX11 + 8.0 , -10.0 < bX1 < 10.0 214 l
l
- startup control valve function, loop B
- normalized flow area for SUFV--loop B
- this signal sent to valve SUFVB = 0.1
- SUB , 0.0 < SUFVB < 1.0 l
- Main flow-control valve function, loop B MFB = 0.5555
- BX1 - 4.4444 , -10.0 < MFB < 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 FEEDPUMP CONTROL f
s
- DELPA is the pressure drop for the loop A MFCV, DELPB for loop B
- initialize signal IF(TIME .LT. 10.0) DELPA = 3.55ES
- bias signal to zero DPAB = DELFA - 3.55ES
- first-order lag of DELPA with a 1.0-s time constant a
- 40-psi limit on both sides DPAL = DPAL + ((DPAB - DPAL)/1.0)
- DELT , -2.44e5.< DPAL < 2.4E5
- remove bias DPA = DPAL + 3.55ES FA = 2.90074e-5
- DPA - 10.0
- initialize signal IF(TIME .LT. 10.0) DELPB = 3.55ES l
- bias signal to zero DPBB = DELPB - 3.55E5
- first-order lag of DELPB with a 1.0-s time constant
- 40-psi limit'on both sides DPBL = DPBL + (DPBB - DPBL)/1.0)
- DELT , .-2.44e5 < DPBL < 2.44E5 215
- .- . - . .- - . . . ~ .
- remove bias "
DPB = DPBL + 3.55E5 i-FB = 2.90074e-5
FD = FC - 0.2975 s
-FE = 0.2
- inte8 rate once per time step, FE, is last time-step value of signal FE FF = FF + 0.2333 * (FE +.FE o )/2.0
- DELT , -10.0 < FF < 10.0
.FG = FF + FE , -10.0 <-FG < 10.0 4
- 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 is required pump speed (523.6 rad /s = 5000 rpm)
-2.0 370.4
! 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 /s2 l
1 l
f i
[
t I
t l'
l
~
216
RC Flow A RC Temp A SG Pres A FW Temp - RC Flow B RC Temp B SG Pres B 3 a l &Al @ 83 M ex em g
[ rip Power FW Flow S en B Le 1 3
n Nn e *-- mig FW Flow A
' bg "h K1 SG; g
@ tgen A Leve T "M il N "I" NMB mar to mp tp mp sns
-.(gg: N Del A De B
- l custa }+ t, g min Y
to C mp i
to C mp 4
to MFW Pump T
B Fig. A-1.
TRAC-PF1 UCS model for Oconee-l.
INPUT 1 1
[ 2 2
( 3
)~ -*
3 4 4 o
Can Also Output to Have Outp t Other Trip' to Otbr Trips no.
Component no.
that is Affected by this Trip Fig. A-2.
Trip System Legend (for Section II.C).
- 1. the trip no.
- 2. description of trip
- 3. input to trip, four kinds available:
S.V. E signal variable input T.S.E. E trip sigeal expression, a mathematical operation of S.V.'s C.B. E a control block No. E any leading no. means it is a trip-controlled trip defined in the input deck by this no. Following a trip controlled trip No., the condition that the input must meet to change the trip set status is indicated:
[ E summation M E product
- 4. trip output (the trip ISET value) can only be -1 E on-reverse O E off
+1 E on-forward i
218
i-APPENDIX B l t
I EXTRAPOLATIONS i
- Because the MSLB' calculation was run to 7200 s, no extrapolation of the-results is necessary. The pressurizer' pressure, downcomer liquid temperatures, and heat-transfer coefficients are shown in Figs. B.1 through B.3, respectively.
j The. uncertainties in the MSLB calculation are categorized as follows:
1
- e SG secondary-side water level-TRAC used a collapsed-level calculation to approximate the AP acasurement.
e - MFW pump trip-because collapsed liquid level is used, MFW pump is tripped i later than if AP were used.
e HPI throttling--core exit temperatures should have been used af ter RCPs I
tripped. TRAC used hot-leg temperatures. HPI should have been throttled
! at ~275 s instead of ~525 s as calculated.
- e RCP restart-42 K subcooling margin must be reached in all loop locations i before RCPs restarted. TRAC only used one location (hot leg); therefore, RCPs should not have been restarted-at ~525 s.
i The first two uncertainties regarding the liquid-level calculation do not
- appear to . affect the primary-side overcooling calculated by TRAC. This is because- the flow into the affected steam generator before the feedwater 4
realignment trip (~50 s) is only through the ' startup flow-control valve (main
, flow-control valves closed because of reactor trip), which limits the flow to
~15% of normal. So, even if the main feed pumps' are running, the flow is limited to approximately the same rate as if they are not running. Therefore, the primary-side overcooling rate during this period (0-50 s) is essentially .
I' independent of whether or not the main feed pumps are running. Also, the EFW 4
pump operation during the first ~85 s of the transient has no effect on the primary-side overcooling, because all of this liquid is bypassed out- the. break 5.
(refer to Sec. III.B. of report for further details). ;
. 219 f
y 9,, -
- a. - - -.r e -i+v-.g-w - - . -.-,3-,wm-,.e.--.1-4+m- --,.-e n4,., , y-m%,m,.,,7,.: q, ..m_., .+t----.c- - ~ ..,
l so ,
~"
.o.
! s% y1,Y(Y .
- - nso
- '- ; --2000 i
/
} .o- ,. . ..
- aco oo. . .
g ,/ .- c ao ,/ -
E
/
> / - moo oo- -
U/ I
~ 1so
- m. L./ .
soo ao o eco 200o 3000 400o 0000 0o00 Moo oooo TNE (s)
Fig. B.1.
Pressurizer pressure.
f sao , ,
neo. MMM .
t
- m. .
E ,,o . , .
goe. l l
1 R TH2 E . ,
216 g ,
- 226 5 me .
d 236 -
246 m- h ~256 -
l +266
. me- -
aco o eco 200o 30o0 4000 sooo ocoo 7000 sooo TWE (s)
Fig. B.2.
Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors).
l I
220 l
3e000 34000 -
M 3 33000 -
v l E 80000-t 8000- .
8 u
g 8000- F h
z y 4000-h" k e . . . . . . .
e isso anno anno meo sees anno suco seso Tile W Fig. B.3.
Heat-transfer coefficients at vessel axial level 6 (all azimuthal sectors).
The offeet of throttling the HPI on primary-system overcooling at ~275 s instead of at ~525 s is expected to be small. This is because the overcooling caused by the energy removal through the affected steam generator is much greater than the cooling provided by the HPI. Therefore, the effects of HPI throttling in the MSLB transient are believed to be insignificant for this time period.
The effects of restarting or not restarting the RCPs are perhaps the most difficult to estimate. In the TRAC calculation, the RCPs were res ta rted at
~525 s because the subcooling monitor was not modeled correctly. When the results were reviewed, it appeared that the RCPs should not have been restarted at that time and probably would have not been restarted at all if modeled correctly. 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 met. 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 downcomer temperatures would have continued to decrease at the same rate as before ~525 s until Sc isolation at 600 s. Then the downcomer fluid temperatures would begin to increase and continue to 221
increase for the remainder of the transient. Also, the system wnuld repressurize to the PORV setpoint, as shown in Fig. B.1.
Because there is only ~75 s worth of additional cooling if the RCPs had not been restarted, there probably will not be much dif ference in the minimum downcomer fluid temperature calculated. Therefore, if all of these uncertainties were removed, it is estimated that the minimum downcomer fluid temperature would be ~405 25 K.
II. TBV TRANSIENTS The TRAC-PF1 results of six TBV failure transients are presented in the body of the report (Secs. III.D and 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 extrapolated parameters are the system pressures, downcomer liquid temperatures, and the heat-transfer coefficients in the downcomer.
The extrapolated pressure histories are presented in Figs. B.4 and B.5.
Following an initial depressurization, the system repressurized to the PORV setpoint in the following four cases: 5A, SB, 6A, and 6B. The system did not ago
_ -2535 p~
186 -
-222 g iso- -1935 g w deotid) BASE k w
idash) PARAMETRIC 1 ~
W dehndsh) PARAMETRIC 2 g E g, E
-1335 e -- h, 5
-sa35 86-W**.N-.**-
Y
-735 E i i i i . . . 4M o 1000 2000 3000 4000 3000 sooo woo Fig. B.4.
Pressurizer pressure histories for Case 5 (Case 5A-base; Case 5B-paranetric 1; Case SC parametric 2).
222
repressurize in cases SC and 6C because the HPI was throttled upon attainment of sufficient primary-systen subcooling.
The extrapolated downcomer liquid temperatures are presented in Figs. B.6 through B.ll. A discussion of factors expected to influence the transient histories through the extrapolation period are presented below. The extrapolated heat-transfer coefficients in the downcomer are presented in Figs. B.12 through B.17.
A. General For the TBV failure transients, ICS failure to return main feedwater to the affected steam generator (s) was specified. The method chosen to simulate this failure was to fix the MFW 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 proved to be significant. Although the MFW pump is tripped and the EFW pumps are operating, a significant flow from the hotwell and through the HFW pump continues. This flow continues to the affected SG(s), even following tripping of the EFW pumps.
- 1. Case 5A. The EFW pumps begin operation at ~210 s. The two motor-driven and one turbine-driven EFW pumps take suction from the surge tank, which empties at ~2500 s. At this time 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 comes through the MFW pump and SUFCV. After the surge tank is depleted, only the turbine-driven EFW pump operates and delivers ~59 kg/s. Total flow through the EFW header is estimated at ~196 kg/s. After ~2500 s the cooldown rate is reduced by two factors. First, the reduced flow to the steam generator reduces the secondary-side heat-transfer coefficient. Second, reduced flow from the EFW pumps results in a higher mixed-mean temperature for the flow through the EFW header. These two effects reduce the cooldown rate by ~25% after 2500 s. The estimated extrapolated downcomer liquid temperature at 7200 s is ~365 1 30 K.
- 2. Case SB. 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 SC Liquid level. Flow through the SUFCV continues at ~125 kg/s. The cooldown rate between ~800 and ~1100 s is -0.0130 K/s. This rate is used to estimate the downcomer liquid temperature at 7200 s. The estimated temperature is ~440 t 30 K.
223
M-
-2535 7--.
386 - j W
- - 2185 I
I
~
-1835 3 l
w l ll solid BASE idash PARAWETRIC 1 w
[
g .I llchndsh) PARAMETRIC 2 m og 105- / -1485 "
l 80- I
/ - 1135 A-B u
86- *
- r. -785 f\ 45 0 1000 3000 3000 4000 0000 acco wiec rig. B.S.
Pressurizer pressure histories for Case 6 (Case 6A-base; Case 6B-parametric 1; Case 6C parametric 2).
M1 = TIETA = 1
~
- THETA = 2 S40 *THLTA = 3
= THETA = 4 _g
- THETA = 5 k M" w
. THETA = 6 h u -428h
" m- $
$ u M - 3683oc g 4So- M E 5 6 - 308.3H 3g0 -2483 l l l 300 . , . . , 188.3 0 1000 2000 3000 4000 6000 0000 7000 f
Fig. B.6.
Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors) for Case SA.
224 l
980 584 3
= THETA = 1 I
a THFTA = 2 066 -
- THETA = 3 -5393 a THETA = 4
- THETA = 5 _
M S30-
- THETA = s -494 5'- -
W W hSc6- -449%
5 5 A
n.
2 4ao. -404g 456- - 359.3
. l l
430 , , , , , , , 314 3 e 1000 3000 3000 4000 8000 0000 1000 Fig. B.7.
Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors) for Case SB.
pro- = THETA = 1 I
aTHETA = 2 ,
a THETA = 3 645-
- THETA = 4
, = THETA = 5
,4=ag x . THETA . .
- 800-
- u u 2 a:
-446
{486-5 L
5 a.
5 - 396.3 3 F 410- s-445- -3463 i
I 430 . , , , , , , 2963 0 1000 2000 3000 4000 3000 sooo 7000 Fig. B.8.
Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors) for Case SC.
225 l - - __ . _ . _ .
a
, . THEA = 1 j
- THUA = 2 640 -
- THUA = 3 ,g
= THETA = 4
+ THETA = 6 e m
X 800- *1HETA = S 4d W W
" b h480- -3623D E
a
-2iA23 f 4m-330 - 222.3 i
340 , , , , i i i N O 1000 2000 3000 4000 6000 0000 M Fig. B.9.
Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors) for Case 6A.
i
" -533
- THEA = 2 900-
- THETA = 3 ,g a THETA = 4
- THLTA = 6 g 0= . THERA = . _
3E M W E E W - 4733 m m A A 900: - 4 36.3 480= -4013 400 $ , , , , , , -3663 1000 2000 3000 4000 5000 8000 T000 r
Fig. B.10.
Downcomer liquid temperatures at vessel axial level 6 (all azimuthal sectors) for Caee 6B.
226
secco-A M
E
} 33000- *
$ 10000-G E sooo-8
$ 0000-M i
g 4aoo-O y sooo. T e . . , , , , .
e meo sono sono sono sono sono isso sono 71ME(s)
Fig. B.ll.
Downcomer liquid temperatures at vessel axial level 6 (all i
azimuthal sectors) for Case 6C.
800
- 572.3 i *THLTA = 1
- THLTA = 3 540-
- THLTA = 3 ,
a THETA = 4
- THLTA = 6 800-
- THLTA = 8 ., 4 W W a: a:
- 362.3 e
A.
e a.
=
E.430-e -M$
380- -222.3 i
340 , , , , , , , 152.3 0 1000 2000 3000 4000 6000 0000 7000 Fig. B.12.
Heat-transfer coefficients at vessel axial level 6 (all azimuthal sectors) for Case SA.
227
secco ,
i l
l
^ 3400o- i M
3_,
k M~ n E
e soooo-2 m
y 800o-u
=
- ' 000o-a
.- acco.
7 w
I 3000-e sino mion sono eino sino sian m soon TIME (s)
Fig. B.13.
Hea t-t ransf er coefficients at vessel axial level 6 (all azimuthal sectors) for Case 5B.
_ 1400o-22 y "~ s 3inoco-G C.
b Scoo-O ^
u _
5 soco-M 5
g sooo-7 I T 3000-h m
e sono sono sono sono sono sono sono TIME (s)
Fig. B.14.
l Heat-transfer coefficients at vessel axial icvel 6 (all azimuthal sectors) for Case SC.
l 228 1
- 33000 ,
14000-A M
3 '
2 : sono-E 3D000-5 2
b soc 3-o u
$ 0000-az 5
4000-7 y anoo- _
o . . . . . . .
e sono seen anno sono sono seco sono sees TlME(e)
Fig. B.15.
Heat-transfer coefficients at vessel axial level 6 (all azimuthal sectors) for Case 6A. ,
i tendo ,
HoDD-
=
M 1 33000- '
s L
g mano-e o
C 8000-u
, 8000- '
O E
g 400-H
$ "~ L 0 . . . . . . .
e seno sono sono sono seno anno soon esas TlWE (e)
Fig. B.16.
IIca t-t rans f e r coefficients at vessel axial level 6 (all azimuthal sectors) for Case 6B. I 4
229
l 4
l t
n"' , 1 l
l H000-A x
h 88000' ,
Y
.- 10000-5 G
- - sooo-b 8
, 0D00-d w
j 4000-3z 3000-e . . . . . . .
e uso sono sono soon soon sono isso sono TIME (s)
Fig. B.17.
Heat-transfer coefficients at vessel axial level 6 (all azimuthal sectors) for Case 6C.
- 3. Case 5_C. 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 ~120 kg/s. The estimated cooldown rate between ~1200 and ~1500 s is -0.0103 K/s. This rate is used to estimate the downcomer liquid temperature at 7200 s. The estimated temperature is ~430 1 30 K.
- 4. Case 6A. The approach used to estimate the downcomer liquid temperature is similar to that discussed for Case SA. The EFW pumps begin operation at ~150 s. The surge tank empties at ~1750 s. Before ~1750 s, the flow through the EFW header to one stean generator 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 downcomer liquid temperature is ~350 30 K.
- 5. Case 6B. The EFW pump does not operate during this transient because the liquid level in both steam generators is too high after the MFW punp trips.
The downcomer liquid temperature cooldown has stopped by the end of the 230
7 calculated transient. The minimum downcomer liquid temperature of ~465 K occurs at,~950 s.
- 6. CaseJ 6C. The EFW pump _ does not operate during the transient. The estimated cooldown rate between ~750 and ~1500 s is -0.0203 K/s. This rate is '
used to : estimate the downcomer liquid temperature at 7200 s. The estimated temperature is ~350 30 K.
B. Summary The extrapolated results for the TBV transients are summarized in Table B-I.
TABLE B-I EXTRAPOLATED RESULTS FOR TBV TRANSIENTS AT 7200 s Downcomer Liquid Pressure lient-Transfer Coefficient Case- ' Temperature (K) (bars) (W/m2 g) 5A ~365 ! 30 ~170 t 5 ~1200 2 400
~58 ~440 30 ~170 1 5 '~1200 2 400 SC' ~430 1 30 ~40 20- ~6000-7500 400 6A ~350 t 30
~170 t 5 ~1200 2 400 6B ~465a 3 30 ~170 .5 ~1200 i 400 6C ~350 1 30 ~40 t 20 ~6000-7500 1 400 E
aminimum occurs at ~950 s s
231
l III. SB LOCA TRANSIENTS Extrapolations of the key parameters (primary system pressures, vessel downcomer liquid temperatures, and downcomer heat-transfer coefficients) to 7200 s are presented in Figs. B.18 through B.23 for the PORV and 4-in.-diam SBLOCAs. For each. transient, the extrapolation assumptions, modeling assumptions / uncertainties, and effect of the assumptions / uncertainties are described. Also, uncertainties on the extrapolated results are estimated.
! 1. Extrapolation Assumptions.- The extrapolation of the PORV LOCA primary 4 system pressure, vessel downcomer liquid temperatures, and vessel downcomer heat-transfer c.efficients presented in Figs. B.18 through B.20 assume the following:
- a. HPI continues to operate
- b. PORV remains open
- c. Accumulators and LPI will not actuate
- d. No operator actions taken e.- ICS, trips, and system components function correctly.
- 2. Modeling Assumptions / Uncertainties. The following modeling assump-f i tions possibly affected the calculated results:
- a. MFW pump speed increased to maximum rated speed.
- b. 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.
! 3. Effect of Modeling Assumptions / Uncertainties. The above modeling assumptions probably would not significantly affect the calculated final vessel y downcomer liquid temperature results. The uncertainty of the modeling has essentially no effect on the calculated downcomer itquid temperature. . Based on the modeling assumptions described, a downcomer liquid temperature uncertainty i of 115 K has been estimated for the extrapolated results.
i 232
B. Four-Inch-Diameter _SBLOCA Extrapolation
- 1. Extrapolation Assumptions. The extrapolation of the 4-in-diam SBLOCA primary system pressure, vessel downcome r liquid temperatures, and vessel
)
downcomer heat trans fe r coefficients, presented in Figs. B.21 through B.26, ,
assume the following: !
- a. HPI, accumulators, and LPI continue to operate
- b. Break is not isolated (closed)
- c. No operator actions taken
- d. ICS, trips, and system components function correctly.
- 2. Modeling Assumptions / Uncertainties. The following modeling casumption affected the calculated results: total LPI volumetric flow rate of 6000 gal./ min (2 pumps) at 50 0F.
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 perticular accident sequence, were found.
- 3. Effect of Modeling Assumptions / Uncertainties. The LPI modeling assumption does affect L5e calculated final system pressure, vessel downcomer liquid tempecature, and heat-transfer coef ficient results. The LPI volumetric 0
flow of 6000 gal./ min (2 pumps) at 50 F reflects the maximum discharge rste and temperature for the LPI system obtained from the FSAk. Variations in the volumetric flow (as a function of rystem pressure) would somewhat alter the '
slope of the system pressure curve, the downcomer liquid temperature profiles, and calculated heat-transfer coef ficients. The extrapolation of the three key parameters following the LPI actuation (~1265 s) is very difficult and should be recognized as a rough approximation. Based upon the modeling assumptions wade, the following uncertainties were estimated for the extrapolated results:
- a. System pressure is t 2.0 x 105 Pa.
- b. Downcomer liquid temperature is i 30 K.
- c. Downcomer heat-transfer coefficient is t 2000 W/cm2 g, C. SG Dryout Followed by EFW Overfeed (Rancho Seco-Type Transient)
- 1. Extrapolation Assumptions. The extrapolation of the SG dryout followed by EFW overfeed transient primary-system pressure, vessel downcomer liquid temperatures, and vessel downcomer heat-transfer coef ficient (Figs. B.27 through B.29) assumes the following
- a. SG 1evel will not be restored to correct level by 7200 8.
- b. HPI will not be throttled to maintain 50 F subcooling by 7200 s.
- c. No further operator actions taken.
233
- d. ICS, trips, and system components function correctly.
By the time the emergency feedwate r has been terminated to the steam generators (4200 s), the primary system does not have an adequate heat sink. As a result, the primary system would begin to heat up (expand), causing an j increase in primary-system temperature and pressure. The secondary system will follow the primary system by gradually heating up and slowly boiling of f the inventory in the steam generators. The primary pressure will reach the PORV set point, and the primary system temperature will gradually approach the saturation temperature of the secondary system. The primary system will not have an adequate heat sink to lower the primary pressure and temperature until the SG inventory is restored to the proper level (24 in. with RCPs operating). The proper steam generator operating level may not be obtained by 7200 s. The following uncertainties were estimated for the extrapolated results:
- a. System pressure is 2.0 x 105 Pa.
- b. Downcomer liquid temperature is 30 K.
- c. Downcomer heat-transfer coefficient is i 500 W/n2 g,
- 2. Modeling Assumptions / Uncertainties No modelir.g assumptions or uncertainties in the TRAC model were found to af fect the calculated results.
we'. . asse EXTRAPOLATED we'- .
ne'- .
I us'- .
un'- mac
(
E we'- ,
te saw'- .
ass'-~ .
u n'- .
w e* , , ,
8 N00 3000 3000 4000 4000 4000 7000 4000 TWE (s)
Fig. B.18.
PORV LOCA extrapolated system pressure.
234 l
l l
l suo , , , , . , -
seo gag. -
EXTRAPOLATED g ass- 2 2 - S**
1 E
eso. -
-us Ses- -
- gge 8**' '
-se
.. \
soo ug. ~+. .
. moo ass- -
-400 Goo , . , , , , .
e moo soon anoo sooo sooo esoo me sees MM Fig. B.19.
PORV LOCA extrapolated downcomer liquid temperature.
18000 ,
, H000-n
,Y 2
y ""
EXTPAPOLATED E 10000- 2 2 2
2 O
sooo-v 5 son-a 5
5 4000-M anco-1 : : : : 1 0 , , , , , , ,
e sono sono sono ecco soon sono soon anos 71 lit (s)
Fig. B.20.
PORV LOCA extrapolated downcomer heat-transfer coefficient.
235
wg , , . - - ' '
.mo we'- .
! EXTRAPOLATED (NO LPI) ,, I ad. *--s -
EXTRAPOLATED (LPI)
[ ad- .
eas I ud-
,g_ .
. een
~
- d' * '
eee
. _ , _ a _ s -s -
W W- ~
. .see wd- : -
"o moo se'oo anos moo anco seco moo anoo ThE (s)
Fig. B.21.
Four-inch-diameter SBLOCA extrapolated system pressure.
aos , , , , , , ,
Ws - -
ase- -
\ .ges ass- -
- i im s
- 2 m- ,
{
+ - * - * .
... j
. ..so
! So- -
ag.
.3gg
- m. -
"- EXTRAPOLATED (NO LPI) -
s__. .ne.
33g. -
. EXTRAPOLATED (LPI) .
soo- -
m , .
e uso anco sooo acco asco eeoo moo aseo TWC (a)
Fig. B.22.
Four-inch-diameter SBLOCA extrapolated downcomer liquid temperature.
236 l
. _ . = -
l I
38000 - I naooo-EXTRAPOLATED (NO LPI) i
- - -- a A
M 2 14o00-
) EXTRAPOLATED (LPI)
- nacoo- = =
5 i g nocoo-b y sooo-a:
y 8000-x = = 1 : : :
5 4000- l9,1 ... ..-4 -. , _ .
7 Y ****'
i o , , , . . . .
e soon anno sono seco ecco anno 1eco anse TIME (s)
Fig. B.23.
Four-inch-diameter SBLOCA extrapolated downcomer heat-transfer coef ficient.
to ts' . , , , , , .
tees s*W- -
s*W-mW- .
I s*#. EXTRAPOLATED .
{
.see &
u d.
l sa,d.
. eae a*W- -
sed- .
so d- -
et e
isso anco f
anco ecco ecco sooo mm
?
sooo e
TIME (a)
Fig. B.24.
Four-inch-diameter SBLOCA extrapolated system pressure.
237 l
1
soo , , , , , , ,
.oso SPs - -
Soo- -
aas- -
i
- E soo- EXTRAPOLATED -
E es- -
2 . .mo l
me- I .
g as-. -
aos g .o. .
g
..so aso- -
ass- , , , ,
,g, aso- -
aPl , , , , , , ,
e uso soon sooo moo sooo sono noo sono M 60 Fig. B.25.
Four-inch-diameter SBLOCA extrapolated downcomer liquid temperature.
3e000 g H000-
{ EXTRAPOLATED
~
g13000- _
w
$10000-G_.
b 8000-u 8
5 0000-g z
$ "" g q
h sooo- f
- 0. , , , , , , ,
o 3000 anno anno esso oooo ecos soon anos TIME (o)
Fig. B.26.
Four-inch-diameter SBLOCA extrapolated downcomer heat-transfer coefficient.
238
masoceo . . . . . i . .asse I seeeeen esseems 1 asse
- e. .
l g .'
-' ~ ' ~
~'~'
fgegesse. .
.ene g aseenes messee . . .mos assesse .
. . wee seesenc . . , , . . .
e see asas asse esos sees sees noe sees MW Fig. B.27.
Rancho Seco-type transient extrapolated system pressure.
ese , , , , , , ,
- ,eee sen- .
- .e,e see- .
see g
E Itas-l ese-see-een 3 -
.as ese- < .
a
,o ,see a
ee- ," .
a aee 48e . . . . . . .
e see sees asse esse sees esse noe sees MW Fig. B.28.
Rancho Seco-type transient extrapolated downcomer liquid temperature.
i i 239
84000
~
F *,s 5*s 3,IS000- ,,=5 g -
E w
G 84000-C u
8 u
5 san 00-hz 5
7 no00- .===
M * *
.- ' s "-
$M s i a s e a s e asse asse sees sees seen sees ines sees 1RE N Fig. B.29.
Rancho Seco-type transient extrapolated downcomer heat-transfer coefficients.
1 I
l I
1 240
APPENDIX C UNCERTAINTIES IN OCONEE PTS CALCULATIONS I. INTRODUCTION l 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 cap 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, and frictional losses); (2) component models (fuel rod, steam generator, valves, and pumps); (3) initial conditions (operating power, system pressure, primary flow rate, SG inventory, and pressurizer inventory); (4) plant model (noding, combined components, setpoints, control delays, and 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. 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, that is, 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 moment can still be large.
Such nonlinear behavior and the overall nonlinearity of the equations
, being solved make estimating uncertainties extremely difficult. One must also 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 Ge best estimate will probably cause the pressure to behave in a like manner. I 1
l 1
241 '
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 flow rates 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 because the initial conditions were defined.
Our algorithm ignored the nonlinear effects and accounter 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-transfer and choked-flow correlations, 10-20%. We used 20% for this study. When we tested TRAC against data from integral experiments, we were able to predict results to better than 20%, but usually only after adjustment of the input model to obtain better resolution in specific regions of the calculations.
Thus, our algorithm for the temperature uncertainty was 6T = 0.2 ITt - Ts I*
and the pressure uncertainty was 6p = 0.2 lpt Psls where Te and P tare the transient temperature and pressure, respectively, and Ts and ps are the steady-state temperature and pressure, respectively.
As the transient values began to approach the steady-state values, then the maximum uncertainty predicted so far was used. These algorithms were used directly to obtain the uncertainty in the primary-system pressure and the downcomer liquid temperature. The initial primary-system pressure was 15.03 MPa, and the initial downcomer temperature was 563 K. We used a 20%
uncerteinty in heat-trancfer coefficient at all times.
242 l
l
III. UNCERTAINTY EFFECTS ON PTS TRANSIENTS We examined 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 these trips. For example, our best-estimate calculations used 10.44 MPa to trip on the HPI. However, at 11.21 MPa, 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 MPa. 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. MSLB The loop-B TBV was tripped (7.064 MPa) open at 5 s; the loop-B pressure increased so rapidly that its uncertainty should have had 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 have had little effect. Advance or delay. of the RCP trip, which occurred 30 s later, may have had 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 either the advance or delay was only a few seconds. At 526 s, the subcooling margin at 42 K was reached, and the HPI was shut off and the RCPs were*r'estarted. A delay would have given colder downcomer 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.
TABLE C -I SYSTEMS AFFECTED BY UNCERTAINTIES Effective System Setpoint Uncertainty Range TBV 7.064 MPa 6.94-7.26 HPI on 10.44 MPa 9.29-11.21 RCPs 30 s after HPI on Accumulators 4.17 MPa 1.46-5.98 PORV 16.9 MPa 16.59-17.37 LPI 1.0 MPa 0.0-3.34 HPI off -42 + 12.5 K subcooling 243
uncertainty is such that the accumulators could have begun dumping as early as
~70 s. This could have caused more cooling, possibly more subcooling, with and the possibility for the RCPs to be restarted earlier. The PORV setpoint was hit at 4678 s; the results were insensitive to the uncertainty when the PORV opened.
B. PORV LOCA The trips observed in this transient were that the TBV opened at 4.4 s, the HPI came on at 70 s, and the RCPs shut off at 100 s. The transient would be l 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 the transient. The pressure plateau from 180 to 380 s 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 downcomer temperature to increase. Although the accumulators were not predicted to actuate, the pressure uncertainty could cause accumulator actuation at approximately 900 s 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 insensitive 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 accumulators would probably shut off almost immediately.
244 l
1'
-E. Two-Inch SBLOCA The transient would be insensitive to the timing of the TBV opening at 4.1 s. HPI 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 .
l actuation is predicted at 1800 s; uncertainty in pressure could lead to actuation as early as 1200 s.
F. Four-Inch SBLOCA i
l 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 s instead of the predicted 540 s,
. or it could be delayed until approximately 1200 s. LPI could have started as
_early as ~600 s instead of the predicted 1236 s.
G.~ Rancho Seco' The Rancho Seco transient is fairly insensitive to the pressure uncertainties. The timing of HPI initiation, predicted at . 738 s, would be
- changed by the uncertainties, but would have little effect on the overall l transient..
1 i
IV. CONCLUSIONS
$ Overall, it is probable that uncertainties in the timing of the actuation l
of the engineered safety features, arising from thermal-hydraulic uncertainties, would have little effect .on pressurized thermal shock results for these i
transients. Only a more extensive sensitivity study could verify this conclusion.
1 i
i i
f 1
245 i
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- w. -n-
DISTRIBUTION ,
Copies Nuclear Regulatory Commission, R4, Bethesda, Maryland 298 Technical Information Center, Oak Ridge, Tennessee 2 Los Alamos National Laboratory, Los Alamos, New Mexico 50 350 246
U S NUCLEAft KEGULATo%v COMutssioN 1 atPuMT NupetM #Assfwa e, rsoc. saa vos %e. .taar#
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- NUREWCR-3706 72,"3E2 BIBLIOGRAPHIC DATA SHEET LA-10055-MS sit INSTavCTIONS ON TME a(Vta$$
2 TITLE ANo SYSTITLE J LE AVE BLAN=
TRAC Analyses of Sev e Overcooling Transients
- o^Te ataoa' Cow'L* T 5 0 for the Oconee-1 PWR #
veaa oo~T.
j i auvaomse r February 1984 i
- '""55 I Compiled by John R. Irel' d ve*=
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May 1985 i r ea.oau,so oacsamz.rio= =*.a .~o w.,u~o .ooaass , s. <, C , < . PaoacT a asa .o u~.r , .a
/ '"""'"**'"
Los Alamos National Laboratory Los Alamos, NM 87545 A7217 6
io sPo%som%o oaoamzatio= =4ve ~o ua umo *ooness uy~.e.co., n. TvPtoe atPoaf Division of Accident Evaluation Informal Office of Nuclear Regulatory Research U.S. Nuclear Regulatory Commission Washington, D.C. 20555 12 EUPPLtutNT Aav NoTts 13 ASSTR ACT (200 woras er <asso This report describes the results f several Tran ent Reactor Analysis Code (TRAC)-PFl calculations of overcooling tran ents in a Babcoe & Wilcox lowered-loop, pressurized water reactor (Oconee-1). The rpose of this stud is to provide detailed input on thermal-hydraulic data to Oak dge National Laborato for pressurized thermal-shock analyses. The transient cale ations performed were p nt specific in that details of the primary system, the secon ary system, and the plant ntegrated control system of Oconee-1 were included in th TRAC input model. The res ts of the calculations indicate that the turbine-b ass valve failure transient s the most severe in terms of resulting in relatively old liquid temperatures in the owncomer region of the vessel. The power-operat relief valve loss-of-coolant acc ent transient was the least severe in terms of owncomer liquid temperatures becaus of vent-valve fluid mixing and near-saturate conditions in the primary system. I is recommended that future calculations con ider a wider range of operator actions cover the spectra of overcooling transie sequences more completely.
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