ML20115G927
| ML20115G927 | |
| Person / Time | |
|---|---|
| Site: | Calvert Cliffs  |
| Issue date: | 04/30/1985 |
| From: | Koenig J, Rich Smith, Spriggs G LOS ALAMOS NATIONAL LABORATORY |
| To: | NRC OFFICE OF NUCLEAR REGULATORY RESEARCH (RES) |
| References | |
| CON-FIN-A-7315, REF-GTECI-A-49, REF-GTECI-RV, TASK-A-49, TASK-OR LA-10321-MS, NUREG-CR-4109, NUDOCS 8504220382 | |
| Download: ML20115G927 (355) | |
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NUREG/CR-4109 LA-10321-MS R4 TRAC-PF1 Analyses of Potential Pressurized-Thermal-Shock Transients at Calvert Cliffs / Unit 1 A Combustion Engineering PWR l Gregory D. Spriggs Jan E. Koenig Russell C. Smith
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Manuscript subrnitted: June 1984 l Date published: February 1985 i Prepared for I Division of Accident Evaluation Office of Nuclear Regulatory Research ' US Nuclear Regulatory Comrnission Washington, DC 20555 ! NRC FIN No. A7315-4 h I
*Present address: Org. 6425, Sandia National Laboratories, P. O. Box 5800, Albuquerque, NM 87185.
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CONTENTS ACR0NYMS................................................................... vi SUBSCRIPTS................................................................. vi 1 ABSTRACT................................................................... I. EXECUTIVE
SUMMARY
................................................... 2 II. INTL 00UCTION........................................................ 5 A. T h e P TS P r ob l e m . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5 B. T he PT S P r o g ra m . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5 III. THE TRAC-PF1 MODEL OF CALVERT CLIFFS-1.............................. 8 A. Primary Side.................................................... 8
- 1. VesSe1...................................................... 8
- a. Radial Rings........................................... 10
- b. Azimuthal Segments..................................... 10
- c. Ax i a l L ev e l s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10
- d. CEA Shrouds and Bypass Flows. . . . . . . . . . . . . . . . . . . . . . . . . . . 12
- e. Fluid Mixing in Ve s s e1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12
- f. H e a t S la b s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12
- 2. Hot Legs................................................... 12
- 3. P r e s s u r i z e r . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12
- 4. SGs........................................................ 13
- 5. C o l d L e g s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13
- 6. C ha r g i ng F low . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14
- 7. HPI Flow................................................... 14 B. S e co nd a ry S id e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14
- 1. F e e dwa t e r T ra in . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14
- 2. S t e a ml i ne s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18 C. C on t ro l Sys t em and T rips . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 20
- 1. Primary-Side Controllers................................... 22
- a. Power.................................................. 22
- b. Pressure............................................... 23
)
- c. Flow...................................................24
- d. V o l um e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 4
- 2. Secondary-Side Controllers................................. 24
- a. Pressure............................................... 24
- b. Flow................................................... 24
- c. V o 1ume . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 5 IV. STEADY-STATE CALCULATIONS.......................................... 26 A. HZP............................................................ 26 B. FP............................................................. 26 V. MO D E L V E RI F I CAT IO N . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 9 v
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r VI. TRANSIENT ANALYSES METHODOLOGY . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3 8 A. Energy-Transfer Analyses....................................... 38 B. Loop-Flow Stagnation........................................... 42 VII. RUNAWAY FE E D WATE R . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 56 A. Runaway MFW to Two SGs from FP................................. 59 B. Runaway MFW to One SG from FP.................................. 65 C. Runaway AFW to Two SG s f rom FP . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7 9 D. Comparisons.................................................... 89 VIII. STEAMLINE BREAKS................................................... 95 A. 0.1-m2 MSLBs................................................... 96
- 1. From HZP................................................... 96
- 2. From FP.................................................... 108
- 3. With Two Operating RCPs from HZP........................... 113
- 4. C ompa r i s o n s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12 3 B. D ouble-E nd e d MSLB s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 127
- 1. With Unisolated AFW to Broken SG From HZP. . . . . . . . . . . . . . . . . . 127
- 2. With Two Stuck-Open MSIVs from HZP......................... 136
- 3. C omp a r i s o n s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14 0 C. Small Steamline Breaks (Stuck-Open TBV)........................ 145
- 1. F r o m F P . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14 5
- 2. With Stuck-Open MSIV from FP............................... 151
- 3. O cupa r i s on s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 6 IX. S MALL-B RE AK L0 CAS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16 0 A. 0.002-m2 Hot-Leg Break from FP................................. 160 B. S tuck-Open PORV with S tuck-Open ADV f rom FP. . . . . . . . . . . . . . . . . . . . 17 3 C. Stuck-Open PG3V from HZP....................................... 184 D. Comparisons.................................................... 192 X. UNCE RTAI NT I E S . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19 5 A. Unce rtainty Sources Cons idered. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 198
- 1. SG Mass.................................................... 199
- 2. C hoke d-F low Mod e 1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19 9
- 3. D ecay Hea t Following Shutdown. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19 9
- 4. Pres sure-His tory E f f ect . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 203
- 5. Feedwater Temperature...................................... 205
- 6. Condenser /Hotwell Liquid Inventory......................... 206 B. Runaway Fe e d wa t e r . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 20 6
- 1. Runaway MFW to Two SG s f rom FP . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 207
- 2. Runaway MFW to One SG f rom FP . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 207
- 3. Runaway AFW to Two SG s f rom FP . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 207 C. Steamline Breaks............................................... 209
- 1. 0.1-m2 Steamline Break from HZP............................ 209
- 2. 0.1-m2 Steamline Break from FP............................. 211
- 3. 0.1-m2 Steamline Break with Two Operating RCPs
[ 4. from HZP................................................... Double-Ended MSLB with Unisolated AFW to Broken 211 SG from Har................................................ 213 vi
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- 5. Double-Ended MSLB with Stuck-Open MSIVs f rom HZP. . . . . . . . . . . 214
- 6. Stuck-Open TBV from FP..................................... 214
- 7. Stuck-Open TBV with Stuck-Open MSIV from FP................ 214 D. SBL0CAs........................................................ 217
- 1. 0.002-m2 Hot-Leg Break from FP............................. 217
- 2. Stuck-Open PORV with S tuck-Open ADV f rom FP . . . . . . . . . . . . . . . . 218
- 3. S tuck-Open PORV f rom HZP . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 218 E. Pressure Uncertainty........................................... 221 F. Downcomer-Heat-Transfer Coefficient Uncertainty................ 221 XI. CONCLUSIONS AND RECOMMENDATIONS.................................... 223 ACKNOWLE DG MENT S . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 2 9 RE F E RE N C E S . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 3 0 NOTE TO THE APPENDIXES.................................................... 231 APPENDIX A. RUNAWAY-MFW TO TWO SGS FROM FP . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 23 2 APPENDIX B. RUNAWAY-MFW TO ONE SG FROM FP . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 24 0 APPENDIX C. RUNAWAY-AFW TO TWO SGS FROM FP............................... 248 APPENDIX D. 0.1 -m2 MSLB FROM HZP......................................... 256 APPENDIX E. 0.1-m2 MSLB FROM FP.......................................... 265 APPENSIX F. 0.1-m2 MSLB WITH TWO OPERATING RCPS FROM HZP................. 274 APPENDIX G. DOUBLE-ENDED MSLB WITH UNISOLATED AFW TO BROKEN SG FROM HZP.. 283 ,
APPEND 1X H. DOUBLE-ENDED MSLB WITH TWO STUCK-OPEN MSIVS FROM FP. . . . . . . . . . 292 APPENDIX I. ONE STUCK-OPEN TdV FROM FP................................... 301 APPEN1)IX J ., ONE. STUCK-OPEN TBV WITH ONE STUCK-opt,N MSIV FROM FP. . . . . . . . . . 310 APPENDIX K. 0.002-m2 JOT-LEG BREAK FROM FP............................... 319 APPENDIX L. STUCK-OPEN PORV WITH STUCK-OPEN ADV FROM FP.................. 328 APPENDIX M. STUCK-OPEN PROV FROM HZP..................................... 338 vii
ACRONYMS ADV Atmospheric dump valve AFAS Auxiliary feedwater actuation signal AFW Auxiliary feedwater flow BG&E Baltimore Gas and Electric Co. BNL Brookhaven National Laboratory C-E Combustion-Engineering Inc. CEA Control-element assembly FP Full power steady-state conditions FSAR Final safety-analysis report HP High pressure HPI High pressure injection HZP Hot-zero power steady-state conditions LOCA Loss-of-coolant accident LOFW Loss of feedwater LP Low pressure MFBV Main-feedwater bypass valve MFIV Main-feedwater isolation valve MFRV Main-feedwater regulating valve MFW Main feedwater MSIV Main-steam isolation valve MSLB Main-steamline break NRC U.S. Nuclear Regulatory Commission ORNL Oak Ridge National Laboratory PORV Power-operated relief valve PTS Pressurized thermal shock PWR Pressurized water reactor RCP Reactor coolant pump RTD Resistant-temperature device SAI Science Applications, Inc., Oak Ridge, TN SBLOCA Small-break loss-of-coolant accident SG Steam generator SGIS Steam generator isolation signal SIAS Safety-injection actuation signal SIS Safety-injection system SRV Safety relief valve TBV Turbine-bypass valve TRAC-PF1 Transient Reactor Analysis Code TSV Turbine-stop valve SUBSCRIPTS C Convective CONT Containment NC Non-convective PRI Primary-side SEC Secondary-side viii i
TRAC-PF1 ANALYSES OF POTENTIAL PRESSURIZED-THERMAL-SHOCK TRANSIENTS AT CALVERT CLIFFS / UNIT 1 A Combustion Engineering PWR by Gregory D. Spriggs, Jan E. Koenig, and Russell C. Smith ABSTRACT Los Alamos National Laboratory participated in a program to assess the risk of a pressurized thermal shock (PTS) to the reactor vessel during a postulated overcooling transient in a pressurized water reactor (PWR). We provided the thermal-hydraulic analyses of three general accident categories: steamline breaks, runaway-feedwater transients, and small-break loss-of-coolant accidents. These postulated accidents included multiple operator and equipment failures. Results were provided to Oak Ridge National Laboratory (ORNL) who plan to determine the probability of vessel failure and accident occurrence for an overall assessment of PTS' risk. As was specified by ORNL, the postula'ted
- overcooling transients were simulated for 7200 s (2 h) after the transient initiation. Our study was performed for a Combustion Engineering (C-E) PWR, Calvert Cliffs / Unit 1,-
using the ' Transient Reactor Analysis Code (TRAC-PF1). The plant owner, Baltimore Gas and Electric Co., and the plant vendor, C-E, provided extensive information for our model. The analyses identified the phenomena important to the PTS issue. Flow stagnation in all reactor coolant loops, which occurred in one transient, could have severe consequences. We found the results to be very sensitive to the initial conditions of the plant. If the plant was initially at hot-zero power (compared to full power), the decay heat was much less, which made it possible for the same accident initiator to produce significantly lower downcomer _ temperatures. However, routine operator actions may reduce the consequences of any of these simulated accidents if the prescribed pressure-temperature relationships are followed. 4 1 i b . .. .
I. EXECUTIVE
SUMMARY
Los Alamos National Laboratory participated in a program to assess the risk of a pressurized thermal shock (PTS) to the reactor vessel during a postulated overcooling transient in a pressurized water reactor (PWR). We provided the thermal-hydraulic analyses of 13 postulated accidents to Oak Ridge National Laboratory (ORNL) scientists who plan to determine the probability of vessel failure and accident occurrence for an overall assessment of PTS risk. Our study was performed for a Combustion Engineering (C-E) PWR, Calvert Cliffs / Unit 1. The plant owner, Baltimore Gas and Electric Co., and the plant vendor, C-E, provided extensive information for our model. Many operator inactions and equipment failures were specified by ORNL for our analyses to conservatively assess the risk of PTS. On the other hand, our analyses were best-estimate in the sense that every effort was made to model the plant as it actually operates (other than the assumed failures) and that we used the bast-estimate thermal-hydraulics Transient Reactor Analysis Code (TRAC-PF1). As specified by ORNL, the postulated overcooling transients were simulated for 7200 s (2 h) after the transient initiation. We found the results to be sensitive to the initial conditions of the plant. When the decay heat was lower (hot-zero power initial conditions versus full power initial conditions), the same accident initiator produced significantly lower downcomer temperatures because of the lower sustained heat source. The ipitial steam generator (SG) secondary side water mass, which fluctuates during daily operations and is known only to within 10%, also had an important effect on the downcomer temperatures for the steamline break transients.
- A description of the transients and the minimum downcomer liquid temperatures during 7200 s are given in Table I. The current US Nuclear Regulatory Commission's (NRC) screening criterion is 405 K (270 F). Because charging flow was not terminated in any transient, all transients repressurized unless a break in the primary system was present (indicated in Table 1). It is important to remember that not only were multiple operator / equipment failures assumed in the transients but also that the bulk downcomer liquid temperature was not the same as the temperature in the vessel wall. The vessel wall was hot initially and the time constant fer energy removal from the wall was long compared to the time constants associated with changes in the fluid temperature.
2
I The vessel-wall temperature and probability of cracking will be determined from the fracture-mechanics analyses performed by ORNL. Flow stagnation is of possible importance to PTS because the cold high-pressure injection (HP1) fluid may stratify in the stagnant (or low-flow) cold leg, flow along the bottom of the pipe and directly contact the vessel wall. Three-dimensional mixing calculations are being performed at Los Alamos and Purdue to further address cold-leg mixing when loop flow was stagnated. No esiculation predicted flow stagnation in both loops. However, we believe that in a majority of the transients that we analyzed, both loops could produce cold-leg temperature stratification from very low loop flows if the decay heat was low enough. Flow stagnation in one loop, (indicated in Table I) was predicted in some cases and could still be of PTS concern. When the flow stagnated in one loop, the effective total heat capacity of the primary system was reduced so that the downcomer temperature decreased more rapidly per unit of energy removed (energy was removed from the flowing fluid only). The operation of the RCPs increased the primary-to-secondary heat transfer. When the secondary side was a heat source, the RCP operation was advantageous (as in steamline breaks) whereas it was not when the secondary side was a heat sink (as in runaway-feedwater transients). In conclusion, our work addressed the thermal-hydraulic consequences of various overcooling transients in a C-E PWR. The analyses identified the phenomena important to the PTS issue. The majority of the initiators could have led to conditions of PTS concern if the decay-heat level had been lower. We recommend that further analyses be performed for different decay-heat levels, particularly for the small-break loss-of-coolant accidents in which flow stagnation in both loops is highly probable at low decay-heat levels. 3 i
)
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TABLE l a TRANSIENT RESULTS Minimum T Repressuri- Flow Stag- ' Description O K F zation nation Runaway-feedwater Cases:
- 1. Runaway-MFW to two SGs from FP 480 404 yes no
- 2. Runaway-MFW to one SG from FP 490 422 yes one loop
- 3. Runaway-AFW to two SGs 490 422 yes no from FP Steamline Breaks:
- 4. 0.1-m2 MSLB
- a. From HZP 395 251 yes yes
- b. From FP 468 383 yes very low-flow in one loop
- c. With two operating RCPs 446 343 yes no from HZP
- 5. Double-ended MSLB
- a. With failure to isolate 377 219 yes yes AFW to broken SG from HZP
- b. With two stuck-open MSIVs 376 217 yes no from HZP
- 6. Small steamline break (stuck-open TBV)
- a. From FP 530 494 yes no
- b. With one stuck-open MSIV 500 440 yes no; calcula-from FP tion termin-ated at 2500 s SBLOCAs:
- 7. 0.002-m2 hot-leg break 440 332 no one loop from FP
- 8. One stuck open primary PORV 407 273 no one loop with one stuck-open ADV from FP
- 9. One stuck-open primary PORV 350 D 171 D
no both loops from HZP "These transients assume no operator intervention except to trip the RCPs see Table II). { Estimated. 4
s r O II. INTRODUCTION
- Los Alamos National Laboratory participated in a program to assess the risk of a PTS to a reactor vessel. Our role was to provide best-estimate thermal-hydraulic analyses of 13 postulated . overcooling transients using TRAC-PFl.I' These transients included multiple equipment failures and multiple operator failures. Calvert Cliffs / Unit-1, a C-E plant, was the PWR modeled for this study.
A. The PTS Problem The reactor vessels of certain older plants with copper impurities in the vessel welds risk cracking if subjected to a thermal shock with subsequent system repressurization (referred to as PTS). After years of irradiation, the vsssel welds in these plants lose their fracture toughness. Hence, as the fluence to the vessel welds increases, the temperature at which a crack may initiate . or propagate increases. Overcooling transients have been postulated that may lower the vessel wall temperature rapidly but maintain or return to high' system pressure. For this reason, in late 1981, the NRC identified PTS as an unresolved safety issue and developed a task action plan (TAP A-49) to essolve the issue. B. T_he PTS Program (TAP A-49) An effort to study older plants of the three PWR vendors was established. A Westinghouse plant (H. B. Robinson) and a Babcock & Wilcox plant (Oconee-1) ware also analyzed as part of the program. For the C-E plant (Calvert - Cliffs-1), several organizations participated: the plant owner, whice is the Bmitimore Gas and Electric Co. (BG&E); C-E; the NRC; ORNL; Brookhaven National Laboratory (BNL); and Los Alamos National Laboratory. The NRC managed the multi-organizational project. BG&E and C-E supplied } extensive Information about the plant and its operation. Los Alamos used this ! information to prepare a comprehensive TRAC-PF1 model of Calvert Cliffs. ORNL identified 13 postulated overcooling transients that could lead to PTS, and Los Alamos simulated these transients for 7200 s (2 h) after their initiation. < These transients were reviewed by BG&E, C-E, ORNL, and BNL. Our results were [. . provided . to ORNL, who plan to extend these results to other postulated PTS transients using a simplified mass-and-energy balance approach. For each of these postulated transients, ORNL plans to determine the stresses in the vessel wall and calculate the probability of vessel failure. ORNL then plans to 5 Ie i
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publish a report that incorporat9s the entire study and identifies the important event sequences, operator and control actions, and uncertainties. The purpose of these calculations is to aid the NRC in confirming the screening criterion in the proposed PTS rule (proposed 10 CFR 50.61). The NRC will also use these analyses to develop requirements for the licensees' plant-specific PTS safety-analysis reports and the acceptance criteria for proposed PTS preventive actions. We emphasize that the transient calculations specified by ORNL were purely hypothetical and not necessarily probable. The transients were chosen to give as much insight as possible (in a minimum set of calculations) to the effect of certain operator and equipment failures, even when the probability of the com-bination of these failures was extremely low. Unless indicated otherwise, the temperatures, flows, pressures, operator and control system actions are calculated by the TRAC code or postulated in the definition of the accident sequences, and they do not represent observations or occurrences at the plant. The operator and equipment failures for each transient are summarized in Table II. TABLE iia ASSUMED OPERATOR ACTIONS / INACTIONS AND EQUIPMENT FAILURES Operator actions / inactions common to all transients (except as specified):
- a. Operator will turn off all RCPs 30 s after SIAS based on low pressurizer pressure.
- b. Operator fails to turn off charging pumps before full repressurization.
- c. Operator fails to control repressurization.
- d. Operator fails to maintain level in intact SG.
- e. Operator fails to respond to high SG-level alarm at 30".
- f. Operator fails to respond to high SG-level alarm at 50".
Additional operator actions / inactions: Number Description Additional Action / Inaction Runaway Feedwater: 1 Runaway-MFW to both None SGs from FP 2 Runaway-MFW to one None SG from FP aFor definition of acronyms, see list on page vi. 6
TABLE II (cont) Additional operator actions / inactions: Number Description Additional Action / Inaction 3 Runaway-AFW to two SGs Operator unable to supply AFW from FP following LOFW for 1200 s. Steamline Breaks: 4a 0.1-m2 MSLB from HZP None 4b 0.1-m2MSLB from FP None 4c 0.1-m2 MSLB with two operating Two of the four RCPs RCPs from HZP were left in operation 5a Double-ended MSLB Operator failed to manually with failure to isolate isolate AFW to the broken SG AFW to broken SG from HZP Sb Double-ended MSLB with AFW terminated at 480 s two stuck-open MSIVs Operator failed to manually from HZP close the stuck-open MSIVs 6a One stuck-open TBV from FP Operator failed to manually close the stuck-open TBV 6b one stuck-open TBV with one Operator failed to manually stuck-open MSIV from FP close the stuck-open TBV Operator failed to manually close the stuck-open MSIV SBLOCAs: 7 0.002-m2 hot-leg break from FP None
^8 One stuck-open primary PORV with Operator failed to isolate one stuck-open ADV from FP stuck-open PORV; Operator failed to isolate stuck-open ADV 9 one stuck-open primary Operator failed to isolate PORV from HZP stuck-open PORV 7
t 1
III. THE TRAC-PF1 MODEL OF CALVERT CLIFFS-1 TRAC-PFi l is a best-estimate finite-difference computer code capable of modeling thermal-hydraulic transients in both one and three dimensions. The code solves the full set of field equations for mass, momentum and energy conservation for both steam and liquid. The Calvert Cliffs model made full use of the capabilities of TRAC-PFl. Calvert Cliffs / Unit 1, located on the Chesaps ake Bay in Maryland, began operation in January 1975. Unit I has a 2 x 4 loop arrangement: two hot legs and two steam generators with four cold legs and four reactor-coolant pumps. The plant operates at 2700 MW. From a PTS standpoint, the following are important features of Calvert Cliffs:
- 1. the HPI pumps have a shutoff head of 8.9 MPa (1270 psig), which is below the normal operating pressure;
- 2. the charging-flow pumps are positive-displacement pumps and are capable of pressurizing the primary system to above the pressure setpoint of the PORVs;
- 3. auxiliary-feedwater (AFW) flow is valved out to the lower pressure SG when a pressure differential greater than 0.8 MPa (115 psid) exists between the two SGs;
- 4. isolation valves on both the feedwater lines and steamlines isolate both SGs if a low pressure of 4.6 MPa (653 psig) is sensed in either SG; and
- 5. the two SGs have relatively large liquid inventories (102000 kg (225000 lb) at HZP and 63000 kg (138600 lb) at FP).
The Calvert Cliffs-1 TRAC model had several evolutionary steps during its development. Most of the changes resulted from efforts to improve the modeling of various system components such as the SGs and the pressurizer. The following describes the current model of Calvert Cliffs-1. A_ . Primary Side Figure III.A.1 shows the TRAC noding diagram of the primary side. Table III gives the metal masses for the primary system that were used in the TRAC model.
- 1. Ve s s,el,. The reactor vessel of Calvert Cliffs-1 was modeled
- three-dimensionally with twelve axial icvels, two radial rings, and six theta segments. The vessel model totaled 144 calculational-mesh cells.
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I g hh ae65 - #### bd ensi41312i 1i1194 119 I k 33 O HPI Fill 3170 - - - HPI Fill 3 f 3M # U I4' i 3 L M/L 1880 2.184 ' Fig. III.A.l. TRAC noding diagram for the primary-side at Calvert Cliffs-1. e
g . TABLE III PRIMARY SYSTEM METAL MASS Mass Component g lb Vessel 482 954 1 064 431 Hot leg (each) 37 322 82 258 SG-tubes only (each) 161 558 356 074 Cold leg (each) 130 932 288 574 Total 812 766 1 791 33/
- a. Radial Rings. The vessel was divided into two radial rings, as shown in Fig. III.A.2. The inner ring represented the core region, located within the core-support barrel. The outer ring represented the annular downcomer region, located between the vessel wall and the core-support barrel. The vessel wall was modeled as a heat slab that interacted with the fluid in the vessel but it did not occupy any of the volume in the downconer.
- b. Azimuthal Segments. The vessel was divided into six symmetric azimuthal segments - one segment for each penetration (four cold legs and two hot legs). All six penetrations are located at the same elevation (level 9 in the TRAC noding diagram shown in Fig. III.A.2).
- c. Axial Levels. The vessel was divided into 12 axial levels. Using the botton of the vessel as a reference point, the top of the first level corresponded to the bottom of the core-support barrel. The top of the second level corresponded to the bottom of the fuel column. Hence, the bottom end fitting of each fuel assembly was located in level 2. The active core height, 3.5 m (11.4 ft), was divided into five equal axial sections. The gas plenum of each fuel rod and the top end-fitting of each fuel assembly were located in level 8. The top of level 8 was at the same height as the bottom of the hot-leg penetrations and the top of level 9 corresponded to the top of the hot-leg penetrations. The top of level 10 was at an elevation slightly above the top of the control-element assembly (CEA) grid-support plate. The top of level 11 corresponded to the top of the CEA shrouds. The CEA shrouds extended approximately 1m (3.28 ft) into the upper head. The top of level 12 corresponded to the total vessel height of 12.54 m (41.1 ft).
10
.__~. - - _ -
12.540 11.261 IEA \ O 10.3W r/11 7 ,_ y_ _;3 LEVEL 9
- NOZZLE LEVEL 1 8"A** 10 COLD LEGS 2 5 8.971 ~ _.
- 7. .
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7 8 S 11 6.642 - O g_94g _ /7T//1//// 1l 2 5 l4 5.254 - Y///!/// ~ 8 2 5 h O O 4_gg9 _ 'A///9nV// ~~ 3 4 gg , ',4//////,/ u 3.170 - 885/8' / 4\ HOT 9 10 HT 2 shSMg"7T&Ey, 9 LEG 3 COLD LEGS 6 LEG 1.8S2 1 na arme Lower Plenum Mixing Pipes 1.880 _ 2.184 Fig. III.A.2. TRAC noding diagram for the reactor vessel at Calvert Cliffs-1. Z
s
- d. _CE_A_Shrou_ds and Bypass Flows _. In the reactor vessel, a small portion of the total flow (~1.9%) goes through the CEA shrouds located in the upper plenum into the upper-head region and is referred to as bypass flow. The flow recirculates back into the upper plenum through 19 small holes located in the CEA grid-support plate. The CEA shrouds were modeled in the TRAC input deck using six pipes - one for each of the six azimuthal segments in the upcomer region.
A small bypass flow (~0.1%) occurs between the core-support barrel and the upper head at the keyways at the upper-head mating surface. In addition, a small bypass flow (~0.6%) occurs at the mating surface between the hot-leg nozzles and the core-support barrel. Both of these bypass flows were modeled in the vessel model of Calvert Cliffs-1.
- e. Fluid __Mixi_n tin Vessel _. C-E performed a series of experiments to measure the amount of fluid entering a vessel similar to the Calvert Cliffs vessel via one loop and exiting via the other loop. Under the conditions of uniform flow in each cold leg and equal cold-leg temperatures, C-E measured a mixing f raction of 27%. To force the TRAC model to predict this mixing, four pipe components were placed in level 1 to induce flow from one loop to the other. The flow area of each pipe was adjusted until the mixing f raction of approximately 27% was obtained.
- f. Heat Slabs _. The heat slabs for the Calvert Cliffs reactor vessel were formulated in accordance with the method presented in Ref. 2. The volume and characteristic thickness of each component were calculated using the nominal dimensions obtained from the drawings supplied by C-E.
- 2. Hot Leg. The hot legs have an inner diameter of 1.2 m (48 in.) at the core barrel converging to 1.0 m (42 in.) outside the vessel. The surge line to the pressurizer was connected to the hot leg in Loop B. Both 5.8-m (18.9 ft) hot legs are divided into five calculational cells.
3_._ _ P r e s s u r iz_e r_. The pressurizer was represented by three TRAC components. The first, PIPE component 10, simulated the part of the pressurizer containing the proportional and backup heaters. Control of these heaters is described in Sec. III.C.I. TEE component 47 was the major part of the pressurizer with a connection to the power-operated relief valves (PORVs) and primary safety relief valves (SRVs). This component contained six cells, which were found to be adequate for modeling the liquid / steam interface. PRIZER component 9, the third of these, fixed the pressure at 15.51 MPa (2250 psia) 12
during a steady-state calculation. The liquid level in the pressurizer was controlled by makeup / letdown (also known as charging) flow during steady state. The liquid level was measured with pressure taps located in cell 1 of component 47 and cell 3 of component 10.
- 4. SG s_. Calvert Cliffs has two U-tube SGs with 8519 tubes of 0.02 m (0.75 in.) outer diameter. These tubes were modeled in TRAC as a single flow path. The heat-transfer area was adjusted 107. so the TRAC calculation would match the steady-state conditions supplied by BG6E. The SG model in this study consisted of 20 primary cells and 26 secondary cells. Seventeen cells modeled the primary side of the tubes. The outlet plenum was divided into two cells so that the flow split of the cold-legs was at the correct location.
The secondary side had three TEE components with a separate injection port for the main feedwater (MFW) and AFW flow. The downcomer converged rapidly at the third cell of the downcomer TEE. The correct recirculation flow was obtained using, correct geometrical data and adjusting the additive friction factor. The moisture separator model in TRAC did not function properly at the time of the study and so phase-separation had to be induced artificially. Only two transients (the double-ended MSLBs) met conditions when the moisture-separators and dryers did not separate the steam from the liquid. For these transients, normal liquid entrainment was calculated. In the other transients, a very large flow area was placed in the steam dome to prevent any liquid carryover into the steamlines following a break. Two liquid-level-measurement instruments were modeled on the secondary. The " narrow-range" level indicator used pressure taps in the steam dome and in cell 2 of the downcomer TEE. This level indicator was used for the reactor / turbine trips and for controlling the MFW flow. The "*..ide-range" level indicator was used for AFW initiation (AFAS). Because of numarical problems not readily identified (conjectured to be caused by the inter;hasic drag), neither the narrow-range by the interphasic drag nor wid:-range A p level-measurement accurately simulated the expected behavior. In transients calculated later, the liquid inventory on the secondary was used to predict AFAS. The method used is specified in each transient section. 5 ._ _Co_1_d Le6s,. The pump-suction leg is a 2-m (6.5-ft) U-shaped pipe leading to the RCPs. The rest of the piping is horizontal, with the HPI and charging flow injecting downstream of the RCPs. Each cold leg was modeled separately and represented the piping f rom the SG to the vessel. Single phase 13
r homologous curves for the head and torque of the RCPs were supplied by C-E. Coastdown data were also given. Two phase homologous head and torque curves were not anticipated to be needed for the 12 transients that were specified.
- 6. Charging Flow. During steady state, makeup / letdown (charging) flow is injected or withdrawn to maintain a specified level in the pressurizer. If the pressurizer level drops more than 0.23 m (9 in.) below its setpoint, the charging flow is injected at a constant flow rate of 3.2 kg/s (7.0 lb/s) into one cold leg of each loop. If a safety-injection actuation signal (SIAS) occurs during the transient, charging flow is increased to a constant rate of 4.1 kg/s (9.1 lb/s) in each loop. Normally the operator terminates flow once the pressurizer level has recovered. However, for the transients in this study, it was specified that the operator would f ail to do so. Thus, a charging flow of approximately 8.3 kg/s (18.3 lb/s) was injected throughout the transient.
- 7. HPI Flow. The HPI pumps at Calvert Cliffs have a low shutoff head of 8.7 MPa (1270 psig), which is below the normal operating pressure, this is advantageous from a PTS standpoint. This limits the rate at which the system can be repressurized. HPI is injected into all four cold legs based on delivery curves supplied by C-E. HPI was modeled as a mass flow vs pressure boundary condition with the fluid at a temperature of 286 K (55 F). The warmer fluid in the HPI lines inside (322 K (120 F)), and outside (302 K (85 F)) of containment was also taken into account. When HPI was initiated, the warmer liquid was pumped into the system before the colder liquid from the storage tank filled the HPI lines.
B. Secondary Sid_e_ It was necessary to include parts of the secondary side of Calvert Cliffs in the TRAC model. This included the steam lines up to the turbine-stop valves (TSVs) and turbine-bypass valves (TBVs) and about half of the feedwater train. The AFW injection line and tanks were modeled approximately. 1._ Feedwat_er Train. Figure III.B.1 shows the major components of the complete main feedwater/ condensate train of the Calvert Cliffs plant. The geometry of the feedwater system was determined f rom isometrics and piping and instrumentation drawings supplied by BG6E. The high pressure (HP) heaters were modeled as one heater as were the low pressure (LP) heaters. Over 1000 m (3280 f t) of pipe length was modeled with 170 fluid cells. Cell lengths were limited to less than 10 m (32.8 ft) to minimize numerical-diffusion effects. Each of the two main feedwater pumps was modeled separately. This allowed one 14 l
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pump to run in manual and the other pump in automatic. The pump curves for the main feedwater pumps were obtained from Science Applications, Inc. (SAI) and converted into TRAC form. Two phase flow through the pumps was not considered possible for the PTS transients in this study and therefore two phase homologous curves were not included in the model. The speed of tne one MFW pump operating in automatic was controlled within the TRAC model using the control-system model. The TRAC'feedwater/ condensate train model was programmed to simulate the following operating behavior of the integral feedwater/ condensate train. Under normal full power steady-state (FP) operation, condensate le pumped from the hotwells of the three main condensers. The condensate is pumped through a series of LP heaters and one set of HP hesters where extraction steam is used to heat the condensate prior to its entrance into the SGs. The extraction steam that condenses during the condensate-heating process in LP heaters 11,12, and 13 is subcooled in the drain coolers and returned to the condenser /hotwells. The extraction steam that condenses in LP heaters 14 and 15, and HP heater 16 is drained into a holding tank and subsequent Ly injected 1 directly back into the MFW train at a point between the last two LP heaters. Following a turbine trip from FP conditions, the bleeder trip valves in the steam extraction lines close, isolating each LP and HP heater. The drain system on each heater will continue to drain condensed extraction steam from the heater until a low liquid level is obtained, at which time the valve on the drain line will close to prevent the heater from completely draining. The drain pumps (which were injecting condensed extraction steam back into the main feedwater) will begin to "run back" and will eventually trip on low level in the drain tanks. Under these conditions, the temperature of the feedwater being supplied to the SGs will begin to decrease at a rate that is dependent on both the rate at which the feedwater is being swept out of the feedwater line and the total stored energy associated with the heat capacity of the pipe walls and the residual amount of condensed and uncondensed extraction steam remaining on the shell side of each heater. Simultaneous to the changes that occur in the HP and LP heater sections of the feedwater/ condensate train following a reactor / turbine trip, the main-feedwater regulating valves (MFRVs) will close and the main-feedwater bypass valves (MFBVs) will open to a fixed position corresponding to a 33% stem position. The one MFW pump that is operated in automatic mode will run back in an effort to maintain a 0.72 MPa (105 psid) pressure drop across the 16 l
feedwater-valve system via the automatic control system. The other MFW pump will continue to operate at its initial constant speed of ~485 rad /s (4631 rpm). This leads to a pressure drop across the feedwater-valve system which exceeds 0.72 MPa (105 psid), and subsequently causes the automatically-controlled feedwater pump to run back to its minimum speed of 314 rad /s (3000 rpm). The combination of one feedwater pump operating in manual at a constant speed of 485 rad /s (4631 rpm) and the other feedwater pump operating at a minimum speed of 314 rad /s (3000 rpm) results in a net feedwater flow of approximately 3.5% of rated feedwater flow (that is, ~30 kg/s (2.38 x 105 lb/h) per SG). Depending upon the initial condition specified for each transient, portions of the model shown in Fig. III.B.1 were deleted or altered prior to the initiation of the transient if they were superfluous. This improved the running time for the integral model. For transients initiated from hot-zero power steady-state condition (HZP), the entire model upstream of the MEW isolation valves (MFIVs) was replaced with a constant mass-flow boundary condition of
~5 kg/s (11 lb/s) per SG f or the time in which the MFIVs were open. This is justified for this initial condition because of the small changes that can occur via the automatic control system at this low power level.
For all of the transients analyzed from FP conditions (with the ~~neption of the runaway-MFW cases), the feedwater/ condensate train model upstrea; of the heater-tank-drain-line injection point was replaced with a constant-tem,,erature boundary condition coupled with a variab1'e pressure boundary condition. The variable pressure boundary condition was used to simulate the aggregate pressure response produced by the pumps upstream of this point during periods of time in which the flow through those pumps was changing. The constant-temperature boundary condition was justifiable provided that a steam-generator isolation signal (SGIS) occurred within two thousand seconds. The total fluid swept out of the feedwater/ condensate train in 1000 s (assuming a flow rate of approximately 4% of rated flow following a turbine trip) represented less than 40% of the total mass of fluid within the feedwater/ condensate piping from the discharge of LP heaters 14 to the inlet of the SGs. Hence, the temperature of the liquid entering the SGs during this interim was completely determined by the temperature distribution formed in the feedwater pipes during the initial steady state and, for the 1000 s interim, was unaffected by the temperature of the liquid specified at the boundary condition. 17 l
In the runaway-main-feedwater transients, a special boundary condition was derived from the entire feedwater/ condensate model shown in Fig. III.B.1. This special boundary condition is explained more fully in Sec. VII of this report.
- 2. Steamlines. Figure III.B.2 shows the TRAC model noding diagram of the steamlines. -The model did not include the steamlines that supply the MFW- and AFW pump turbines. Furthermore, some liberty was taken with the arrangement of the line to the TBV. The line to the TBV is actually between the Loop-A main-steam isolation valve (MSIV) and the lines to the high pressure (HP) turbines.
The relative position of the lines to the TBV and HP turbines is inconsequential because both lines are downstream of the MSIVs, and the TBVs and turbine stop valves (TSVs) are never open at the same time. Venturi-flow restrictors were located between cells 1 and 2 in components 52 and 62 about 10 m (32.8 ft) from the SGs. They were calibrated to deliver 170% of FP steam flow under choked-flow conditions with 5.7 MPa (850 psia) SG pressure. Five sets of valves are in the steamlines: the TSVs, MSIVs, SRVs, atmospheric dump valves (ADVs), and TBVs. The two TSVs in the TRAC model represented four actual valves and closed in 0.25 s following a turbine trip. The TSVs were calibrated to deliver FP steam flow with a SG pressure of 5.86 MPa (850 psia). The MSIVs closed in 3.5 s following SGIS and never reopened. The SRVs represented a bank of pressure-modulated valves. They began to open when the upstream pressure reached 6.89 MPa (1000 psia) and were wide open when the pressure reached 7.45 MPa (1080 psia). A flow area vs pressure table was specified to simulate the behavior of the actual bant of valves. Each SRV wac 6 calibrated to deliver 763 kg/s (6 x 10 lb,/h) of saturated steam when the valve was wide open and the upstream pressure was 7.45 MPa (1080 psia). The ADVs were trip-activated and controlled by the average reactor temperature. They opened in 3.0 s following a reactor / turbine trip when the average reactor temperature exceeded 552 K (535 F). The flow area varied linearly with the average reactor temperature between 552 K and 565 K (535 F and 557 F). The small hysteresis between the closing and reopening temperature was not modeled, and the stroke rate was limited to 33%/s. The TBVs were trip-activated and controlled by either the average reactor temperature or the steamline pressure upstream of the valves. The TBV represented four actual valves, and its stroke rate was also limited to 33%/s. During steady-state (FP or HZP) the TBV was regulated to limit the steamline pressure to 6.24 MPa 18
e .. 2 64 67 6 b b b 83 88 nicH PRESSURE s es
\
SG 22
/
es mi t i , , ,7, 71 [ _3]72-- : I h CONDENSER 2 ti a l 3 1 TsV4 1x W@ 75 @ e hjh 54 57 Kh
]
4 @ V@ L\ ADV T@ L\SRV 3 J -- I SG 12 Im seks MSIV a e 4 tina @ HIGII PRESSURE 4 / TURBINE Fig. III.B.2. TRAC noding diagram for the steamlines at Calvert Cliffs-1. G .
(905 psia); in practice, the TBV is fully closed during FP and ~4% open during HZP. However, following a reactor / turbine trip from FP, the valves were controlled by the maximum steamline pressure and average reactor temperature signals, and the flow area fraction was given by
"#* [(P-895) '
[T-535g 10 22 where P is the steamline pressure in psia and T is the average reactor temperature in F. The average reactor temperature was calculated by averaging the temperature in each loop; the temperature in each loop was calculated by averaging the temperature in the hot leg with that in one of the cold legs of that loop. C. Contr_ol System and Trips Table III.C.1 suanarizes the setpoints for the trips at Calvert Cliffs. In addition to mo % ing those portions of the control system that directly affected the components explicitly modeled for this study, the TRAC model included simulation of systems that were not modeled explicitly but could potentially alter the behavior of the system by trips, etc. For instance, instead of modeling the ArW condensate storage tank, a controller that integrates the AFW flow is used to terminate flow when the tank's contents would be exhausted. In this way, many large auxiliary systems are modeled with relatively little modeling or computational effort. The control system included both control-block operators and logic-controlled trips. Control-block operators are mathematical operators that generate an output value as a function of one or more input values. The input values can be signal variables and/or control-block outputs that measure the values of important process variables. The set-status of a trip was either "on" or "of f", depending upon the values of signal variables, control-block outputs or the set-status of other trips. The control system was divided into two parts the primary-system controllers and the secondary-system controllers -- and the controllers in each part were categorized ac:ording to their control objective. 20
TABLE III.C.1 SETPOINTS FOR TRIPS AND SIGNALS Trip or Signal Setpoints Primary-side:
- 1. Reactor trip a. PPRI < 14.5 MPa (2100 psia)
- b. SG level < -1.27 m (-50 in.)
(on narrow-range instruments)
- c. Asymmetric-SG pressure signal
- d. SGIS
- e. Turbine trip
- 2. SIAS on low thermal margin PPRI < 12.1 MPa (1740 psig)
- 3. HPI flow P PRI < 8.8 MPa (1270 psig)
- 4. Charging flow (3 pumps) SIAS
- 5. RCPs trip SIAS + 30 s (as specified by ORNL)
- 6. PORVs open PPRI > 16.3 MPa (2400 psia)
S condary-side:
- a. Reactor trip
- 1. Turbine trip
- b. SG-level > +1.27 m (+50 in.)
(on narrow-range instruments) P
- 2. SGIS SEC <>4.6 P
cont 0.127MPa MPa(653 psig) (4 psig) 1
- 3. MFIVs close SGIS
- 4. MFRVs close Turbine trip MFBVs open to 33% Turbine trip 5.
- 6. MFW pump trip a. SGIS
- b. High discharge pressure of 9.87 MPa (1450 psia)
- c. Low suction pressure resulting from low liquid inventory in hotwell (less than 1 kg (2.2 lb))
- d. Liquid in line to turbine pump (a factitious trip included to account for potential MFW-turbine damage) 21 i
TABLE III.C.1 (cont) Trips Setpoints
- 7. AFAS (method varied a. SG 1evel < -4.3 m (-170 in.)
in calculations) (on wide-range instruments) or
- b. SG liquid inventory < 45000 kg (99000 lb)
- 8. Asymmetric-SG pressure AP3g > 0.8 Maa (115 paid) signal
- 9. AFW flow AFAS - flow valved out to SG at lower pressure if asymmetric-SG pressure signal has been received
- 10. MSIVs close SGIS
- 11. ADVs modulate 552 K < T (5350 F<pht<565K PRI < 557 F)
- 12. TBVs modulate 552 K < T (5350 F <PRI T < 5650 K 6.17MPa<hRI<<557F)or 6.24 MPa (895 psia <S{CrSEc < 905 psia),
whichever normalized value is higher
- 13. SRVs open PSEC > 6.9 MPa (1000 psia)
- 14. TSVs close Turbine trip i
i
- 1. Primary-Side Controllers. This subsection describes the parts of the control system that regulate the power, pressure, flow, and volume of the i 3r' mary side.
- a. Power. During FP operation the reactor thermal power is a constant
! 2700 MW, and the RCP Lhermal power is 17.38 MW. The primary temperature adjusts itself to the value necessary to effect transfer of the power to the secondary system. A reactor trip will occur if at least one of the following conditions l is satisfied: 4 l l 22 l t
(a) the primary pressure is less than 14.5 MPa (2100 psia); (b) the narrow-range SG 1evel is less than -1.27 m (-50 in.); (c) the SG pressures differ by more than 0.8 MPa (115 psid); (d) SGIS occurs; or (e) the turbine trips. All these trips were modeled. Following a reactor trip, the turbine trips and the steam dump / bypass cystem regulates the ADVs and TBVs to control the average reactoe temperature. Above 565 K (5570F) the valves are wide open; below 552 K (535 0 F) they are fully closed, and between these limits, *Ne flow area is adjusted linearly with temperature.
- b. Pressure. The primary pressure normally is controlled by the pressurizer heater / sprayer system. Because thu sprayers would be activated only '
for a short time -in the simulated overpressure transients (their operating pressure range is small), they were not modeled. Thus, only the heaters were !c the TRAC model. During steady-state operation, the proportional heaters maintain the system pressure at 15.5 MPa (2250 psia) by delivering a maximum of 300 kW to compensate for heat losses from the system. These heat losses are approximately 150 kW, so that the net energy addition from the heaters ranges f rom +150 kW to -150 kW. The TRAC model of the proportional heaters represents a heat input of 150 kW at 15.3 MPa, decreasing linearly to -150 kW at 15.7 MPa. During a transient, the backup heaters were input to deliver 1200 kW to the pressurizer liquid when the primary pressure fell below 15.2 MPa (2200 psia). When the liquid level fell below 2.56 m (1.01 in.), the operation of both the proportional and backup heaters was prohibited. If the level cubsequently rose above 2.56 m, the proportional heaters, but not the backup heaters , reactivated.
- In the plant, two of the four banks of backup heaters come back on automatically if the level recovers. However, this information was received citer the calculations were initiated, and so we did not model this. The effect of these heaters is felt to be insignificant from a PTS standpoint because the cystem is already at a high pressure by the time the backup heaters would reactivate.
23
Excessive pressure relief was provided by the trip-controlled PORVs. These valves were input to be fully open 1.0 s after the pressure increased , above 16.5 MPa and fully closed 1.0 s after the pressure decreased below , 16.3 MPa.
- c. _ Flow. The RCPs were modeled to operate at constant speed until the operators tripped them 30 s after SIAS, as specified by ORNL. The SIAS trip occurred when the pressure fell below 12.1 MPa (1740 psig).
i d.. Volume.. The primary system volume is normally controlled by the j makeup / letdown system. In the event of a severe depressurization, however, SIAS overrides the makeup / letdown system, and the safety-injection system (SIS) begins injecting borated water into the system. , l The makeup / letdown flow was determined by a proportional pressurizer-level controller. The control setpoint was 5.5 m (215 in.) during FP and 3.7 m f j (144 in.) during HZP, and the controller gain was 28.34 kg/s-a (11.45 gpm/in.). The maximum make-up flow rate of 6.48 kg/s (103 gpm) was achieved when the level 4 fell .23 m (9 in.) below the setpoint, while the maximum letdown flow of 8.3 kg/s (132 gpm) was achieved when the level increased 0.23 m (9 in). above its setpoint. The makeup / letdown flow was split evenly between two diagonally-opposite cold legs. Following any SIAS signal, the charging flow was increased to 8.3 kg/s and was not controlled automatically by pressurizer level. l 2. _S_e_co_ndary-Side _C_ontrolle_r_a_._ This subsection describes the parts of the f control system that regulate the pressure, flow, and volume on the secondary side.
- a. P re s su r_e ._ During FP, the secondary pressure was determined by the j inlet pressure to the turbine, and it was not directly controlled. Following a i turbine trip, the pressure was controlled only if it exceeded 6.17 MPa (895 psia), in which case the flow area of the TBV was adjusted to maintain the l pressure between 6.17 MPa and 6.24 MPa (895 pria and 905 psia). Normally the action of the steam dump / bypass system vents enough s tear. to aaintain the i secondary pressure well below 6.17 MPa (895 psia) following a turbine trip.
f During HZP, the flow area of the TBV was adjusted to maintain the secondary i j pressure between 6.17 MPa and 6.24 MPa (895 psia and 905 psia). ' b Flow._ The MFW flow is regulated by an instantaneous level error,
! integrated level error, and instantaneous feed-steam mismatch. The TRAC control i system used these same signals with the addition of the integrated feed-steam !
mismatch to regulate the MFW flow. The reset time of both the level error and i ! 24 i
feed-steam mismatch integrators was 240 s (4 min), the gain on level control was 100%/m, and the gain on feed-steam mismatch was 0.2%/kg/s. Following a turbine trip, the MFRVs closed in 20.0 s, and the MFBVs opened to a stem position of 33% in 1.33 s. During HZP the MFW flow was held constant. During FP, the MFW pump speed was regulated to maintain the pressure drop across _ the MFRV to Loop-A SG at 0.72 MPa (105 psid). The integral controller had a minimum output of 314.16 rad /s (3000 RPM), a maximum output of 586.4 rad /s (5600 RPM), and an-option to hold the speed of Loop-B MFW pump constant. c ._ Volume. As discussed previously, during FP the SG level is normally controlled by regulating the MFW flow. In the event of loss of SG mass, a low SG-level indication by the wide-range instrument would initiate AFW delivery to prevent SG dryout. Because the temperature of the MFW entering the SGs decays when flow from the heater-drain tank ceases, it is important to know the inventory of the tank. Although the tank was not modeled explicitly, the steady-state inventory, the inlet, and the outlet flow were all known. Before a turbine trip, the inlet and outlet flow were assumed to balance; but af ter a turbine trip, the inlet flow became zero. Therefore, a heater drain-tank mass integrator began reducing the steady-state tank inventory by the known outlet flow following a turbine trip. When the residual tank inventory fell below a specified value, the outlet flow
.was tripped off to simulate the low-tank-level pump trip that would occur.
< In the event of a steamline break, steam that normally would remain in the system escapes, and the condenser /hotwell inventory would fall below a specified value, thus it was necessary to know the inventory of the condenser /hotwell. Although the tank was not modeled explicitly, the steady-state inventory and the inlet and outlet flows were all known. Therefore, a condenser /hotwell mass inventory calculator was constructed with control-block operators to indicate when depletion of the inventory would trip the MFW pumps. Although the AFW condensate storage tanks were not modeled explicitly, the initial inventory and AFW flow rate were known. Therefore, an AFW mass flow integrator was used to reduce the initial inventory until the residual inventory c was less than 1.0 kg, at which time the AFW flow was reduced to zero. l l [ 25 l 4
IV. STEADY-STATE CALCULATIONS Our calculations began from two separate sets of initial conditions: HZP and FP. A. HZP The TRAC-PF1 HZP steady-state calculation for Calvert Cliffs-1 yielded very stable primary-side conditions but oscillatory secondary-side conditions. The fundamental difficulty in determining the secondary-side conditions during HZP occurred because the vapor generation rate was very small, and appeared to destabilize the steady-state solution for the SG model. Table IV.1 presents a comparison between the actual plant conditions and the conditions generated by TRAC af ter 15 min (reactor time) of the steady-state calculation. The comparison is reasonable with the exception of secondary steam flow. A simple energy balance dictates that in the steady state, the correct [ value for the steam flow was ~10.0 kg/s (22.0 lb/s). The over prediction of TRAC suggests that the SG had not yet reached a complete equilibrium condition in the steady-state run. Nevertheless, the temperature profiles appeared reasonably close to a steady-state condition in which the cold feedwater heated to the saturation temperature by the time it entered the riser section. In the riser, the small vapor-generation rate yielded a very small void fraction until the liquid surface was reached. Although a steady state was not completely obtained, we believe that the TRAC HZP steady-state solution was close enough to the actual plant conditions to allow reasonable simulation of transients initiated from HZP. , B. FP f l Eight of the thirteen transients were initiated from FP steady-state I conditions. During FP the reactor operates at 2700 MW with an additional energy input of 17.38 MW from the RCPs. The temperature increase across the vessel is
~26.4 K (47.6 F) with an inlet temperature of 559.3 K (547.0 F). The pressure drop through the loop is ~0.54 MPa (78.7 psid). Makeup / letdown flow regulates the pressurizer level to 5.46 m (215 in.).
I Heat is transferred through two SGs to the secondary loop. The feedwater flow is regulated by the MFRVs to maintain a specified liquid level using a three-mode controller. The valve area is determined from the SG level and feed-steam flow mismatch as described in Sec. III.C. The MFW pump speed is adjusted to maintain a constant pressure drop of 0.72 MPa (105 psid) across the MFRVs. l 26
f TABLE IV.1 COMPARISON BETWEEN TRAC AND MEASURED PLANT DATA AT HOT-ZERO-POWER CONDITIONS Parameter Measured Plant Data TRAC Predictions Primary Side
- 1. Pressure 15.51 MPa 15.51 MPa (2250 psia) (2250 psia)
- 2. Fluid temperature 550.9 K 551.8 K (5320F) (5340F)
- 3. Power 100 hrs after shutdown 9.38 MW decay heat
+ pump power + 17.38 MW from the pump
- 4. Mass flow 19300 kg/s 19700 kg/s 6
(153. x 10 lb/h) (156. x 106 lb/h) 5.- Pressurizer 3.66 m 3.66 m (144 in.) (144 in.) Secondary Side
- 1. Pressure 62.05 MPa 61.74 MPa (900 psia) (896 psia)
- 2. MFW temperature 300 K 300 K (80 F) (80 F)
- 3. Steam flow 10.1 kg/s 11.0 kg/s (22.2 lb/s) (24.1 lb/s)
- 4. SG inventory 95000 kg 102000 kg l (210000 lb) (225000 lb) l
- 5. TBV flow area -
5% open The feedwater is heated to 495 K (431 F) 0 by one HP feedwater heater and five LP feedwater heaters. The liquid nass in each SG is ~62350 kg (137458 lb). The FP transients were initiated from different FP steady-state calculations. 1As the Calvert Cliffs model evolved during the calculation of the transients, it was necessary to rerun a steady-state calculation whenever the model was modified. Table IV.2 gives a comparison between the TRAC calculation and the measured plant conditions for the last steady-state calculation. j Results compare well as did the results for the previous calculations. 27 l I - - - _ . - , - _ .
TABLE IV.2 COMPARISON BETWEEN TRAC AND MEASURED PLANT DATA AT FULL-POWER CONDITIONS Parameter Measured Plant Data TRAC Predictions j Primary Side
- 1. Core Power 2694 MW 2700 MW
- 2. Vessel flow 25.3 m3 /s 24.9 m3 /s (401100 gpm) (395250 gpm)
- 3. AP gg 0.19 MPa 0.19 MPa (28.15 paid) (28.15 psid)
- 4. AP 0.54 MPa 0.55 MPa 100P (78.7 psid) (80.5 psid)
- 5. T 559.3 K 559.5 K co d (547.0 F) (547.7 F)
- 6. 26.4 K 26.0 K AT*****
(47.6 F) (47.0 F) Se_condary Side
- 1. Feedwater flow per SG 749 kg/s 737 kg/s r
(5.95 x 106 lb/h) (5.85 x 106 lb/h) i f 2. SG Dome Pressure Loop-A SG 5.90 MPa 5.9 MPa t (856 psia) (856 psia)
- Loop-B SG 5.86 MPa 5.9 MPa j (850 psia) (856 psia)
- 3. MFW Pump Discharge Pressure Loop-A SG 7.8 MPa 7.66 MPa l
(1131.3 psia) (1111.0 psia) l Loop-B SG 7.63 MPa 7.56 MPa (1106.7 psia) (1096.5 psia) j
- 4. MFW Temperature 494.8 K 496.2 K (431.0 F) (433.5 F)
I
- 5. MFRV flow area (% open) ~90 93 l 1
- 6. SG liquid mass 62350 kg 63000 kg
( l (137458 lb) (138600 lb) ! l l 28
V. MODEL VERIFICATION On December 9, 1982, Calvert Cliffs / Unit 1 experienced an unanticipated loss-of-load frc a FP. For model-verification purposes, this plant transient was simulated using the TRAC model of Calvert Cliffs. The TRAC predictions were compared to the plant data that were collected during this event (see Figs. V.1 through V.10). (As previously mentioned, several different models for Calvert Cliffs have been used throughout this study. A loss-of-load comparison was run for most of the versions. However, only a comparison for the current model will , be presented.) Because the plant data supplied to Los Alamos included 28 s of data prior to the initiation of the transient, the TRAC calculation was shifted 28 s to compare both the transient behavior and the initial conditions. Hence, in Figs. V.1 through V.10, the initiating event occurred at 28 s. Figure V.1 compares the total thermal power (fission plus decay heat) predicted by TRAC to the fission power measured by two of the plant's power channels. The TRAC prediction correlated with plant instrument NR005A very well. Plant instruments NR005A and 1N009,- however, did not agree at all. One of the two instruments obviously was in error. Using the " prompt-drop approximation," it can be demonstrated that IN009 was grossly in error. Assuming that the control rods inserted approximately 15 dollars of negative reactivity (Ak = 0.100 is the value supplied by Calvert Cliffs), the prompt-drop i approximation predicts that the reactor fission power immediately following the rod insertion is:
"_ = - -
= 6.25 % .
no 1 - (-15) This corresponds to the value predicted by TRAC and measured by instrument [ NR005A. Figure V.2 compares TRAC's prediction of the primary pressure to the pressure measured by two pressurizer pressure taps. The comparison is excellent i for the period of time in which data was collected. j Figure V.3 compares TRAC's prediction of the pressurizer level to the level measured by three level taps (it was assumed that instrument lL110x was different from instrueent LilCx). TRAC modeled the level measurement as a a P tap. The calibration of the AP tap was performed at rated temperature and i 29
12 0 ,
^
10 0 ---- : :- * - c TRAC CALCUt ATION a a PLANT DATA (tP.C9) ,
- PLANT DATA (NR005A) b 60 a -
b h a
& t 40 ,
a - A 20 a, ( * *
- a 3 ,,
*. T. m +, q ;-- e - - ,,'*'?
- A*^_2t ^a,e c
-20 0 10 2G 30 43 % 00 70 60 90 10 0 Time (s)
Fig. V.l. Loss-of-load comparison--reactor thermal power. m
-2300
-t 1 1 1 1 1 7 ,
,5 - -
Y - 2150 o TRAC CALCULATION u.5 - -- 210 0 l E 6 PLANT DATA (P0100X)
-2050 l
- PLANT DATA (P0100Y) u- s y
l s 2000 ! 13.5 -
\ -1950 l 1900 l 13 i
0 10 20 30 40 50 60 70 80 to 10 0 Time (s) Fig. V.2. Loss-of-load comparison- pressurizer pressure. l 30
4 J pressure. The pressure signal was lagged by a 2-s time constant to approximate the actual lag time associated with the plant instruments. However, it appears that the TRAC prediction for pressurizer level deviated somewhat. The discrepancy cannot be explained at this time. i Figure V.4 compares TRAC's prediction of the Loop-A hot-leg temperature to the temperature measured by instrument T0111x, a resistant-temperature device ! (RTD). Two sets of T M C results are plotted against the experimental data. The solid line (with no markers) represents the actual liquid temperature in the hot leg. The solid line (with markers) represents the hot-leg temperature lagged by an 8-s time constant. The " lagged" hot-leg temperature simulates the RTD instruments that feed the control system logic, which, in turn, modulate the ADVs on the secondary-side steam lines. The comparison of the plant data to the lagged TRAC prediction indicates that a time constant of 8 s is slightly too
.long. Nevertheless, the comparison is still good.
In a similar fashion, the TRAC prediction of the Loop-A cold-leg ! temperature was compared to the plant instrument T0111y (also an RTD) in Fig. V.5. Although the general shape of the cold-leg temperature response was , predicted, the magnitude of the initial temperature increase was underpredicted. This resulted from under predicting the heat-transfer degradation that occurred , on the secondary side of the steam generators following the rapid decrease in I steam flow rate produced by the closure of the TSVs. In other words, the transient response of the SGs was not correct. They overpredicted the amount of energy removal from the primary system during the first 5-10 s following a reactor trip. Hence, the liquid leaving the SGs was overcooled. Figure V.6 compares the TRAC prediction of the SG pressure to the pressure measured by instrument 1P1013. Instrument IP1013 appeared to be reading 0.68 MPa (100 psi) low during steady state. Calvert Cliffs personnel informed l us that the actual steady-state value was approximately 5.78 MPa (850 psi). ! Hence, to compare the change in SG pressure, the plant data were increased by a i j constant AP to match the steady-state value predicted by TRAC. As can be
- i. observed from Fig. V.6, the plant data indicated an initially larger increase in ,
l SG pressure following the turbine trip than did TRAC. However, the final measured pressure was lower than that predicted by TRAC. These pressure differences are probably related to the incorrect energy removal rates that produced the cold-leg temperature disparities previously discussed. i i 31
6 S.S - 4......._..... a 200 5-- a - l a PLANT DATA (LO110x) 4.5 -
=g . PLANT DATA (LO110Y)
~
n ** ji -
- t 4 .
- PLANT DATA (1110X) T v
di. a, s 3.3 ** - * .6 a* 3-~ a - "O n'N
.S -- " ' * = . . . . . , , , ,~50 ,,-
3
..sa LS . .
60 0 10 20 30 40 50 80 70 80 to 100 Trne (s) Fig. V.3. Loss-of-load comparison- pressurizer level. 590 600
- a 3 . 4,3 A'^
MS ' "a 'a a ~
- TRAC CALCULATION 590
$80
, A PLANT DATA (T0111x) n 6 S80 g
. TRAC CAL (LAGGLD BY a >> L r s75 - ..
570 i 570- .g
* ~ 588
.T m
**S~ \
.If
- .6 a
sso
\
1
,,, x~s:s . . . ,___ w.
i 550 - 0 10 23 33 43 $0 CO 70 80 9J 100 Trne (s) Fig. V.4. Loss-of-load comparison--Loop-A hot-leg temperature. 32
564 555 a e. *
- 562 , - TRAC CALCUL ATON
'^9
^ ft/NT D.*.1A (101111)
%0 a,
\.. 4 e U
a
- 1RAC CAL (LAGAD bT 8 5)
^
^
k 6 , , I 9 . d ,g 545 b J 556 g 543 A 554 N.'\.N %
'wN S35 552 , ,
O to 70 30 40 50 00 70 80 00 80 0 Trne (s) Fig. V.5. Loss-of-load corsparison--Loop-A cold-leg temperature. 6.6 , 6.S - g o TRAC CALCULATION
~
GA-" a PLANT DATA (F1013)
~
8 I
6.3-I i - y==- g 1 4.1 - - - -' E85 #* i l 6-- s70 ess s.s - S.8 0 10 20 30 40 50 60 70 80 s0 10 0 , Time (s) Fig. V.6. Loss-of-load comparison--SG-A pressure. i 33
Figure V.7 compares the TRAC prediction of the SG narrow-range level measurement to the level measurement obtained from instruments L1113a and l , IL1105. The level measurements in the SGs in the TRAC model were obtained using a AP measurement. However, for reasons not yet determined, the TRAC SG model did not respond in many respects to an actual SG response under similar conditions. For example, immediately following the turbine trip, the liquid velocity in the downcomer increased by a factor of four, whereas the actual increase in downcomer velocity should only be on the order of 10%. This surge resulted in additional form losses at the area contraction in the downcomer cnnulus, resulting in an erroneous AP measurement. This in turn produced an stroneous level measurement. Although the narrow-range level instrument is less censitive to form losses than the wide-range instrument, it also was affected by the large velocity increase. Hence, the disparity between the TRAC prediction cnd the plant data appears to be attributable to an imperfect SG model. Figure V.8 and V.9 compare the TRAC prediction of feedwater flow to SGs A snd B with the measured feedwater flow. According to Calvert Cliffs plant personnel, the flow meters used to make these measurements become unreliable as the flow decreases. This is very apparent in Fig. V.8. At 68 s the flow cbruptly decreases from 200 kg/s to 30 kg/s (1.58 x 106 to 2.38 x 105 lb/h) indicating that perhaps the instrument is in error. In Fig. V.9, the comparison is a much better, but still indicates a significant disparity. Figure V.10 compares the TRAC prediction of the MFW pump discharge pressure against the pressure measured by P4490. According to plant personnel, cne of the two MFW pumps is normally operated in manual control while the other pump is operated in automatic. The plant personnel claim that during a normal loss-of-load transient, neither feedwater pump trips on high discharge pressure (setpoint of 9.87 MPa (1450 psia)). This is consistent with the results cbtained by the TRAC model. However, during this particular transient, one of the two MFW pumps at the plant did trip. To make the TRAC simulation of this transient valid, the setpoint of the one feedwater pump operating in manual was crbitrarily lowered to 9.73 MPa (1430 psia), which caused it to trip at 10 s following the turbine-trip signal at 38 s. As can be observed in Fig. V.10, the comparison of the pressure history following the turbine trip is not good. The initial pressure condition of the TRAC model is 0.41 MPa (60 psia) lower than the value measured by P4490. The net pressure rise in the discharge header l cccording to TRAC was approximately 2.04 MPa (300 psia), whereas the net 34 l
r 0.5 . 0,,. 0 33....----. o TRAC CALCULATION
-0.5 - ---20 a PLANT DATA (L1113A)
- PLANT DATA (11105)
, ,. ,4O t
fC -LS-
* * ---80 b
- a 3 * *
--.30 2 a ,
..:..a
- aa' *
-2.5 a ---100
.,,,,,,,...+*******
-3 -. 12 0
-3.5 , , .
O e 20 30 40 So 60 70 40 oO 10 0 Tm (s) Fig. V.7. Loss-of-load cornparison-SG-B narrow-range level. e00 nSO ilooooooooooooo_
- ~-1500 000- c, -
, 1250 S00 - ,
o . TRAC CALCULATION
~
o o PLANT DATA (F1t11) o M , ,00 O, .
'o o -S00 o -
s00 go_ _
- 2M
( ooooooooooootoo2
-mo . . . - ,
0 m 20 30 43 S0 so 70 so oO iOO Time (s) Fig. V.8. Loss-of-load comparison--SG-A feedwater flow. 35
000
-1750
< ocococcooocco i i
700 o
-_g 600-o 1250 SCO- -
o - IRAC CALCULATION N 6 *' 3 o PLANT DATA (IFt121) 3 E 'b0 e 300- ' - E
.,00 200~ -
~
10 0 0- o c c o JtRRrt a c c o o n o o o o B o o o o o o o o o o 'o- 0 90 , , . . > 0 10 20 30 40 50 SO 70 80 90 10 0 Time (s) Fig. V.9. Loss-of-load co:nparison-SG-B feedwater flow. 9.S 1350 9- i o TRAC CALCULATION --1300 i 3.$ - j
/ *
- PLANT DATA (P4490)
I .. I s_ i y,$ . . l 250
- y. N; r?~ : :
l l 6.$ , O to 20 30 40 50 to 70 80 90 10 0 Time (s) Fig. V.10. Loss-of-load comparison--MFW purnp discharge pressure. I \ 36
pressure rise observeo during the loss-of-load transient was only 1.36 MPa (200 psia). Part of this discrepancy might be attributed to the discrepancy cbserved in the SG done pressure. Any disparity that occurs in the prediction cf the SG done pressure propagates back into the discharge pressure of the MFW pumps. In conclusion, this TRAC model correlates with some of the key parameters en the primary side reasonably well. However, the reliability of any TRAC model fcr predicting secondary-side parameters that interact with the control system (narrow- and wide-range level measurements, SG pressure, feedwater flow, etc.) is uncertain because the results are naturally subject to the large uncertainty inherent in dynamic two phase systems such as steam generators. Generally speaking, uncertainties in the basic correlations contained in the constitutive p:ckage are rarely accurate to more than i 10%; hence, rate processes governed by these correlations are similarly limited in accuracy. However, n2twithstanding the uncertainty in the rate processes the conservation equations cre highly accurate and faithfully track the transport of mass and energy in the icng-ters. In Sec. X of this report, an attempt is made to quantify these uncertainties and their effect upon the final results. This should aid the judicious application of the results of this study to assessing the PTS risk for Ccivert Cliffs. 1 i 37 l
VI. TRANSIENT ANALYSES METHODOLOGY From a thermal-hydraulic standpoint, only a few plant parameters are necessary to evaluate the potential of a severe overcooling transient leading to a PTS accident. The two most important parameters are the time-dependent liquid temperature in the vessel downconer annulus at the elevation of critical welds and the system pressure. The combination of a rapid and sustained cooldown of the liquid temperature in the vessel downconer annulus followed by system repressurization could result in a PTS accident. Of equal concern are overcooling transients in which the flow in one or more loops becomes very low or stagnant. If cold HPI/ charging flow is injected into a cold leg of a stagnant loop, the liquid in the cold leg stratifies, and, as a result, remains relatively unmixed with the hotter liquid in the stagnant loop. Subsequently, the cold HPI/ charging flow liquid flows directly into the vessel downcoser and possibly blankets the vessel downconer wall at the critical welds. If this situation persists for a prolonged period of time, a PTS condition may occur if the system repressurizes. The analyses of the transients presented in this study concentrated primarily on these PTS phenomena. A full explanation of the response of each plant component during every transient is not attempted. Rather, we discuss only the plant parameters that are necessary to assess the thermal-hydraulic consequences of the overcooling transients. In particular, the following parameters are presented in the text of this reports liquid temperature in the vessel downconer; primary system pressure; HPI/ charging flow; f cold- and hot-leg flows; and liquid temperature upstresa of HPI/ charging flow ports. l Additional plots of system parameters are included in the text if they directly aid in the understanding of that transient. Plots of system parameters that are not critical for evaluating the PTS risk have been relegated to the appendixes. A. _ Energy-Transfer Analyses To understand the time-dependent liquid-temperature history in the downconer annulus for each transtant, the net energy transfer into or out of the primary fluid is separated into its constituents._ For ease of discussion, the constituents are categorized as either non-convective or convective. Convective l 38
- _ . _ - - . - - . . = _ -.
N 1 caergy is added or removed from the primary fluid via mass transfers directly 4 into or out of the control volume (i.e., the primary system). Examples of l ccnvective energy transfers include the energy addition produced by the l injection of HPI/ charging flow, and energy rejection by mass leaving through a , primary-side break or an opened PORV. Non-convective energy is conducted into cr out of the control volume across the control volume boundaries. Examples of nsn-convective energy transfers include the energy addition produced by the 1 dscay heat from the reactor core, the energy added by the cooling of l ! primary-side structural material, and the energy removed or added by the SGs. Although the energy added by the operation of the RCPs is technically considered i 1
" work," it appears in the control volume as heat addition because of pressure '
icsses resulting from friction. Therefore, during this study, the pump " work" a has been combined with the non-convective energy terms. Although TRAC is a non-equilibrium code, the liquid mass is so much I greater than the vapor mass in the primary system that the bulk temperature of I the liquid can be estimated accurately by a simple energy balance without including energy changes that occur in the vapor component. Hence, separating i the net energy transfer into its constituents represents a direct means of identifying the major cooling mechanisms in each of the transients. Once the magnitudes of the individual cooling mechanisms are known, the fraction of the tcaperature decrease produced by each can be calculated using the following ! thermodynamic relationship. ) An energy balance for the control volume can be written as i
"[in eh-[oug sh + [g P g (VI-1) l where l
U = total internal energy of primary fluid, l l [in th = sum of energy convected into primary fluid per unit time, bout th = sum of energy convected out of primary fluid per unit time, and [gPg = sum of energy conducted into or out of primary fluid per unit l time. Ws have assumed that all changes in kinetic and potential energy are negligible. In addition, we assume that the primary fluid does no work, nor is work done on l 39 {
the primary fluid (other than by the pumps). Integrating the above expression yields U = U, + EC+ ENC , (VI-2) where Un = initial internal energy, E C = net energy convected into primary fluid, and ENC = net energy conducted (non-convective) into primary fluid. Mathematically, EC "" NC *** #"" 7 E C" in [ hdt-[out[ hdt , (VI-3) and ENC " i [ Pg dt . (VI-4) The specific internal energy is thus Muoo +EC+ ENC , u= (yg_3) M where M,= initial mass of primary fluid, u,= initial internal energy of primary fluid, M = time-dependent mass of primary fluid, and u = time-dependent specific internal energy of primary fluid. The time-dependent mass of the primary fluid is determined by M=M,+[g,[adt-[out[edt . (VI-6) l J 40 [ L_
Using a standard thermodynamic table (see Ref. 3), we can determine the time-dependent bulk temperature corresponding to the time-dependent specific internal energy. Assuming that the downcomer temperature changes are in direct proportion to changes in the bulk temperature, we can estima_t_e_ the temperature changes in the downcomer annulus produced by the various constituents. For exatsple, at any given point in time, the temperature change produced by c!nvective terms can be calculated by setting ENC =0 and recalculating u and its corresponding T. The energy transfers produced by the convective and non-convective sources cre discussed for each transient. In addition, we have plotted a time history cf each of the non-convective terms. (Currently TRAC does not readily allow the user to extract and plot the time-dependent history of the convective terms. BIcause of the time constraints of the project, a " hand" calculation of E had C to suffice.) To put the energy transfers into perspective, the following numbers chould be kept in mind when reading the transient descriptions. The bulk liquid temperature at HZP conditions is approximately 552 K i (5340F) and the initial liquid mass is approximately 224 x 103 kg (491 x 103 lb). The total energy that must be removed from the liquid to d: crease its temperature to 373 K (2120 F) is AU = M, (uS52 K ~ "373 K)
= 224 x 103 (kg) x (1227.4 (kJ/kg) - 418.9 (kJ/kg))
= 181 x 109 W-s .
I H:nce, the energy removal required to change the bulk temperature from 552 K to 373 K is 181 GW-s. At FP conditions, the bulk liquid temperature is approximately 573 K (572 F) and the initial mass is approximately 219 x 103 kg (482 x 103 lb). To d: crease the bulk liquid temperature to 373 K (212 F) requires an energy removal of AU = M, (uS73 K ~ "373 K ) 41
= 219-x 103 '(kg) x (1332 (kJ/kg) - 418.9 (kJ/kg))
= 200 GW-s.
r. 4 As will be shown in the transient analyses, the energy removal capability ' for one SG at HZP conditions is ~150 GW-s following a steam 11ne break. In
- comparison, the energy-removal capability for one SG at . FP conditions is
~110 GW-s following a steamline break. The cooling capabilities are different because the initial asss in the SG at FP conditions is ~35% less than at HZP
- conditions.
In contrast, to lower the primary fluid temperature by 200 K (360 F) by injecting-286-K (55 F) HPI/ charging flow into the primary system at the maximum } HPI/ charging flow rate of 90 kg/s (7.13 x 105 lb/h) would require ~4000 s. The total mass injected by the HPI and charging system necessary to produce the
-200-K cooldown would be ~1.6 times the initial mass of the primary system. This analysis assumed that at ~900 s, the primary system would become liquid full.
For times greater than 900 s, it was assumed that the cold HPI/ charging flow
- sixed perfectly with the liquid in the primary system and was subsequently rejected-through the PORVs at the mixed temperature. This results in an exponential cooldown with a period of ~2400 s (0.67 h) for times greater than 900 s.
B. Loop-Flow Stagnation Assuming this " perfect mixture" and using the above figures, we see that one broken SG can cool the primary system down to 386 K (236*F) starting from HZP conditions. However, if one loop of the primary system stagnates, i approximately one-third of the liquid mass becomes unavailable to six with the liquid subjected to che cooldown. Hence, cooling the downconer temperature by 200 K would require only ~120 GW-s rather than 181 GW-s. Because loop stagnation occurs in the majority of the transients analyzed for this study and I potencia.11y effects the downconer temperature in two different ways (by cold-les e t t e tifica tion and reduction of effective primary-system heat capacity), a description of the basic mechanism that leads to loop stagnation is presented. Following SIAS, the operator is presumed to trip all four RCPs within 33 s. Any tion that occurs in the primary loops following the trip of the RCPs ! !c produced by natural circulation. The major driving force inducing the l-i natural circulation flow is produced in the vessel. The liquid entering the l
- 42 L
core is heated by the decay heat dissipated from the fuel rods. The density difference that forms between the liquid in the downcomer annulus and the liquid in the core region produces a force, FDH, that tends to drive flow in the
" normal" direction of hot leg to cold leg (see Fig. VI.1). (Some liberties have been taken in the drawing of the primary system shown in Fig. VI.1. First, the two cold legs of each loop have been combined into a single cold leg to simplify the schematic. Second, the elevations at which the cold legs penetrated the vessel are shown to be lower than the hot-leg penetrations; in actuality, the elevations of the cold- and hot-leg penetrations are identical.)The magnitude of this force is, of course, dependent upon the magnitude of the decay heat.
During a 7200 s (2 h) transient initiated f rom FP, the decay heat will decrease from 200 MW at 10 s af ter a scram to 30 MW at 7200 s (2 h). From an initial condition corresponding to HZP, the decay heat can be as low as a few megawatts. Hence, because the magnitude of the driving force depends on of the operating history prior to the initiation of the accident, the loop flows can vary significantly depending upon the initial conditions. An additional natural-convection driving force is created in each of the two SGs. The direction of the force produced in each SG depends on the direction in which energy is being transferred within the SG. If the temperature of the primary-side fluid entering a SG is higher than the temperature of the fluid in its riser region, then energy will be transferred in the " normal" direction of primary-to-secondary. The primary-side fluid will cool down as it passes through the U-tubes of that SG. Hence, the primary-side fluid in the downside of the U-tubes will be cooler than the primary-side fluid in the upside of the U-tubes. This creates a density difference between the two vertical sections of the U-tubes, which, in turn, produces a driving force, FA or FB , that acts in series with the driving force produced in the vessel and enhances the natural circulation flow in that loop (see Fig. VI.2). On the other hand, if the temperature of the primary-side fluid entering a SG is lower than the temperature of the fluid in its riser region, then energy will be transferred in the " reverse" direction of secondary-to primary. The primary-side fluid will heat up as it passes through the U-tubes of that SG. Hence, the primary-side fluid in the downside of the U-tubes will be hotter than the primary-side fluid in the upside of the U-tubes. This creates a density difference between the two vertical sections of the U-tubes and a corresponding 43 l-
' t l ' '
o A A n
. i G
S
'd o e t
a ( e
) r 2 c
= (
e s c g r g e y o e L t f L d i r g t l se n o A o wnt i H .m C oea v - LDW i
\ r l d l
y l .b I H Vd e D l
. F l
. c o O gu .
X i d t ( Fo a r e l ph wy oa l
"yy t i r l c f e d _
) hse n _
2 gnt oy g B ( i ea ib HDW t e ." cl - L s ee g t e rs is o L d e H v - d l o "laor j C mt rc oa o Ne _ " r B c_ s o A t l! 11ili:ti!!;! ,j!f iI ij! i) ,l!
SG.B SG. A (QA QB) in " normal" direction of primary-to-secondary i / High Low Low : c High Density / Density Density Density Water : g / 2 Water Water Water QB QA Resultant Resultant Force Force o
~ 't A-Hot Leg Hot Leg j (
AB g 5A ; , FDH Cold Legs (2) Cold Legs (2)
= sA sn :
( X 0 ll $ Fig. VI.2. Gravity head force produced in vessel and gravity head force 3 produced in SCs acting in series for situation in which heat is transferred in " normal" direction of primary-to-secondary. l
driving force that opposes the driving force produced in the vessel. The loop flow for this situation will be correspondingly lower (see Fig. VI.3). For the two-loop arrangement of Calvert Cliffs, six combinations of SG heat-transfer modes can exist. These are: both SGs in "zero" heat-transfer mode; both SGs in " normal" heat-transfer mode; both SGs ih " reverse" heat-transfer mode; one SG in " normal," one SG in "zero" heat-transfer mode; one SG in " normal," one SG in " reverse" heat-transfer mode; and one SG in " reverse," one SG in "zero" heat-transfer mode. Each of these combinations leads to a different solution for the flows that will occur in each of the two loops. The first three combinations cause symmetric flow in the loops; the magnitude of the flow depends on which direction energy is being transferred in the SGs. The remaining three combinations cause asymmetric flow to be produced in the loops and, under certain conditions, can lead to loop stagnation in one loop. (As will be discussed later, the third combination cannot lead to loop stagnation without cther situations occurring.) Both SGs can enter into a "zero" heat-transfer mode if, for example, a oteamline break occurs and the MSIVs fail to close. In this situation, both SGs would dry out. The energy transferred between the primary and secondary sides et this point would be essentially zero. The loop flows for this situation would be equal and governed strictly by the gravity head produced in the vessel (see Fig. VI.4). For the sake of comparison, the loop flows produced by this situation shall be referred to as the base case in the following discussion. Using the previous example of a non-isolable steamline break, the initial portion of the transient would correspond to a situation in which both SGs are operating in a " normal" heat transfer mode (see Fig. VI.5). The secondary-side fluid temperature would decrease symmetrically in both SGs, extracting energy , from the primary-side fluid during the blowdown. The loop flows during this cymmetric cooldown would be determined by the sum of the gravity head produced in the vessel and the gravity head produced in each SG. The loop flows would be i higher than those of the base case during this portion of the transient because cf the added gravity head created in the SG. When both SGs are in the " reverse" heat transfer mode, the gravity head produced in each is in opposition to the gravity head produced in the vessel (see Fig. VI.6). The loop flows will be lower than the base-case loop flows. However, neither loop can stagnate as a direct result of the opposing gravity 46
l ! SG.B SG. A (q,, 9 3) in " reverse" mecuan of i secondary-to-primary; & is flowing in
\ / " normal" direction since FDH ( A' B
! q B
/ \/
4 j A i
- Low High High Low Density Density Density Density
- Water
- = Water Water r : Water Resultant Resultant Force Force
_ 't _
l' F E "E B ) k A 1 - SB , , EA : FDH Cold Legs (2) Cold Legs (2) sa : = EA i v om e l 1 i ( Fig. VI.3.
! Gravity head force produced in vessel and gravity head force j , in SCs acting in opposite directions for situation in which 4
- heat is transferred in " reverse" direction of secondary-to-primary.
6 A G S 9 "A 0 F
=
A A e i n e r ( a
) A 2 6 s
; ( G S .
s - g g e h t F DH e L L bo~ d y t l o A o hb c H 5 C i n h e l wvi nr l id H 4 ns D a Ii oi B n Vt w ao 6 w . gtF ii ul l F s . A r e 6 fodon l s s r
)
2 : ne of g B ( i e b t s n L s g c e ar t e r t o L i - H d t-d a 2 l o we oh
} C l
_ n
& fpr"o B _ oe oz
_. G S A9 " B 0 F
=
B j L" A q 8
- ,1i!!!1 !l ;Ii!!l i
e uI A y A G S s
/
\
9 A (v n i A e( F r ( ) p a+ 2 ( s sH G D s SF g ~ g h e t y e L ob L b d n t l o A o he H E C cv i i hr l wd ns i i l H D S. nw oo Ii l B n m Vt a F 6
=
a . u gt i i l F sdeo 6 l ro m f r e sf
)
- ns g B 2
( on i a e b t r L s ct g t e e-rt o L i a H de d h l o w j C o n l "l s f a
< pr m, B _ oo)
/
G S
\
\
B E y L" ong F A q
)
g
o 1 i lt SG.B SG. A
! \ !
l
% /
QB 4A
~
l "A B t i l 1 l 1l 1l l - y- y - Hot Leg i B} Hot Leg ( A C bR 5A = j Cold Legs (2) DH Cold Legs (2) i coun y } l { Fig. VI.6. 1 Loop flow directions for situation in which both SGs are in i
" reverse" heat-transfer mode and FDH > (F A , FB ). Flow is j
driven by ~FDH - (FA , F B)
- I
. .. - .. _- . . - . _ _ _ - . - - =.
I i i head in the SGs. This is the consequence of the energy transfer via the SGs being produced by a temperature difference _ rather than by a predesignated heat . . flux. In the limit, if the loop flows were to decrease to zero as the result of
- . reverse heat transfer, the fluid in the U-tubes would eventually reach thermal squilibrium with the secondary-side fluid temperature, and the gravity head
- would dissappear. The gravity head in the vessel, however, being produced by a
{ predesignated heat flux (the time-dependent decay heat), could disappear only if l the decay heat in the reactor core disappeared. Hence, in the limit, as the flow decreases to zero, the gravity head in the vessel would begin to increase 1 if a finite amount of decay heat were present in the core. With the loss of the gravity head in the SGs and the increase in the gravity head in the vessel, loop ! flows would be induced. Note, however, that the equilibrium loop flows under l these conditions might be quite small, and for all practical purposes,
- considered stagnant. (Under LOCA conditions, enough primary fluid may be lost so as to allow the liquid level in the vessel to drop below that of the hot l legs. The hot legs will begin to drain and the top of the U-tubes will l
cubsequently void so that the natural-circulation loop will be lost. The
" reverse" heat transfer in each SG plays a major role in the integral process of this type of transient, but is not the di_ rect cause of the loop stagnation.)
! Under asymmetric SG conditions (that is, asymmetric in the sense that the I snergy transfers to the two SGs have different magnitudes and/or directions), l the loop flows also become asymmetric. For the case of one SG operating in the
" normal" heat-transfer mode and the other dissipating essentially zero heat, the
, loop flows will equilibrate at unequal values (see Fig. VI.7). The flow in the loop in which the SG is dissipating zero heat will correspond to the flow produced by the driving force of the vessel. The flow in the loop in which the SG is dissipating heat in the " normal" direction will be enhanced because of the cdditional gravity head produced in the U-tubes of that SG. This situation cccurs, for example, in a steamline break in which the broken SG is completely dry while the MSIVs have successfully isolated the intact SG and cold AFW is being injected into the intact SG. The cold AFW cools the secondary-side fluid { . temperature in the riser region below the primary-side fluid temperature, thus allowing " normal" heat transfer to occur in that SG. During the initial portion of the above example, the fluid temperature in I the broken SG is rapidly decreasing and extracting energy from the primary fluid via the " normal" heat-transfer mode. During this period, it is possible for the 51 1
- _ . _ . _ . _ , , _ - _ _ . _ _ _ _ _ _ . _ _ _ , . _ ____ _ _. _ J
U 2 i i i 1 l SG.B SG.A i k / \ / ! N
~
l Q" B 9A / b A B F" B i j o j F-Hot Leg Hot Leg A m j ( i ,j A ,B < > ",, ; j FDH Cold Legs (2) Cold Legs (2) ! sa : : 6 g am
- g I
i I. i j Fig. VI.7.
- Loop flow directions for situation in which one SG is in i " normal" heat-transfer mode and the other SG is in "zero" heat-transfer mode. Flow in loop A is driven by ~FDH + FA*
Flow in loop B is driven by ~F bH* t
r primary-fluid temperature to drop below the fluid temperature in the riser ! reigion of the intact SG causing it to enter into a " reverse" heat-transfer mode (cee Fig. VI.8). Depending upon the magnitude of the gravity head in the vessel and the magnitude of the gravity head produced in the broken-loop SG, the 4 opposing force produced in the intact loop (because of reverse heat transfer) may be sufficient to stagnate the intact loop. (Under the condition of no decay I h:st in the core, the flow may actually be reversed in the intact loop.) If the 1 cop does not stagnate, the flow will still be significantly lower in the intact icop than in the broken loop. As previously mentioned, if a loop stagnates, the
- fluid in the U-tubes will eventually come back into thermal equilibrium with the escondary-side fluid temperature. The gravity head in that SG will slowly dissappear. As the loop flow tries to start up again, the gravity head is partially re-established if reverse heat transfer persists in that SG. Hence, the loop flow is continually suppressed because of reverse heat transfer in that SG.
{ Eventually the broken-loop SG will dry out, and this will result in the occurrence of the sixth heat-transfer mode-one SG operating in the " reverse" hrat-transfer mode while the remaining SG is not dissipating any heat (see Fig. VI.9). The flow in the broken SG will correspond to the gravity head
- produced in the vessel. The flow in the intact SG will continue to remain essentially stagnated until the " normal" heat-transfer mode is re-established in j that generator. This is usually accomplished by the injection of cold AFW fluid into the intact SG causing the secondary-side fluid temperature to drop below l
i the primary-side fluid temperature. l Throughout this report, reference to loop stagnation will be made without re-explaining its causes. 1 l 53 l
t o A y - (v A
' F A .
n G A in "e i S k s
\ s re w
)
i v lo e F ( Gr
) A S" 2 5
( e nit n.
; s g o n e sa i ng g L h e = cG at L d i S l h s t o wres o A C H 5 nhi it l o B
. n e l
- 8. oi h p It t o H c Va o D
i i t
.t udLn
- A1l gi a n A m is F e i F l r o ow+
d fo m l H l r F D F t se ~ n nf a ) os .y n 2 ( ineb t ad g a g O s cro et m e n . t S e g r- v L - e i t ri rp t n L d aerefd ao o E d w h sns eo l l l 6 o NL o ai C ~ f "l r l j atA B m- . 5 prt p B ooao
/ '8
, oneo B L" hl
/
g B G S
\
Q . s'
- , ' : ); !i ! {i! i j ! '1 ;I 4 ii 1,llji: !,
A 0 0 A _ . = = G A A Q F n S i n "o i r
, s z ew
( 6 i o
) l 2 F
( Gn S i s . g e st g e n i n e L o a L d : Gn S g t l h a o o c t H 3 C ir s 6 he wh t s l n oi i l eB
. h H $ nt o p pD l
l 9. I t i d o o l d V a nl X . ua gi ei n ( t i F sd o l l r m wo . o l H t n f rF e DF a s fs ~ n g n n .y g ) a t e L 2 ( io ae b t t rd o n S rp o t 0 s g c-e t m e v
~ r ao H e i a ri e er eo n L 0 d
_ NL S hf d d ~ s o "ens l
} o C 5 B
l w sr ai f rt A B e-
! u
, pvt p B oeao oreo
. B L" hl G \
Q _ S - u" N !. ! : i :i* !1 ! ;iij.i.!l l 1jl i)ij2 jj')ji);
i VII. RUNAWAY FEEDWATER It has been postulated that if the MFW regulating valves (MFRVs) fail to close following a reactor / turbine trip, and the MFW pumps continue to pump at or above rated. flow into the SGs, the hct liquid residing in the SGs will be
, replaced with significantly colder liquid within a short period of time. The decrease in the secondary-side-heat-sink temperature will produce a corresponding decrease in the primary-liquid temperature and a potential PTS transient.
! An additional type of PTS transient associated with runaway feedwater has also been postulated from a different set of circumstances. Because the AFW , system, when activated, represents a large source of cold liquid that is injected directly into the SGs, failure of this system to be deactivated by operator action in accordance with operating procedures and/or mechanical failure would continue to produce a decrease in secondary-side-heat-sink
- temperature. Again, the consequences of a decrease in heat-sink temperature could lead to a PTS transient.
To assess the PTS consequences of these types of runaway-feedwater events in Calvert Cliffs-1, three transients were studied. In the first, runaway _feedwater was initiated by a reactor / turbine trip in which both MFRVs stuck open at their initial steady-state value (~ 90% stem position). This caused the ( liquid temperature in both SGs to decrease, and, subsequently, to cool the ! primary side in a symmetric fashion. The second transient was similar to the j first except that only one MFRV stuck open. This produced a lower secondary-side temperature in the SG receiving the high feedwater flow and led to an asymmetric SG condition that eventually stagnated one of the primary loops. The third transient was a runaway AFW event. In this event, it was assumed that the MFW pumps tripped, which led to a reactor / turbine trip shortly thereafter because _of low SG level (based upon narrow-range level). In addition,'it was assumed that the AFW system failed to activate on low SG level (based on wide-range level). At 1200 s, the operator was able to manually i activate the AFW system to a full flow of 400 kg/s (3.17 x 106 lb/h) per SG. !' Hence, at t.his point an overcooling transient was initiated.
- To run the first two transients, the TRAC feedwater-train model had to be t
altered sl:.ghtly. As stated in Sec. III.B.1, for the majority of the transients that were planned for this study, SGIS, which closed the MFRVs and isolated the 56
. - - . . _ - = - . - - - _ - - _ _ -
l' t feidwater train from the SGs, was assuned to cccur within the first 1000 s of 1 the transient. Under these conditions, it was necessary to model only half of the feedwater train to correctly predict the feedwater flow and feedwater tcaperature into the steam generators before SGIS. The boundary condition at the inlet to the partial feedwater-train model was set to a constant-temperature boundary condition corresponding to the temperature of the feedwater leaving low pressure (LP) heaters 14 A, B, and C (see Fig. III.B.1 for the location of I these heaters). For the two runaway-feedwater cases in which the MFRVs failed to close, the partial feedwater-train model with its constant temperature boundary 1 condition could not yield the correct temperature decay of the feedwater entering the SGs for the entire duration of the excess feedwater flow. The boundary condition for these two cases was changed to correspond to a semi-espirical correlation for the temperature decay at the exit of LP heaters 14 A, B, and C derived from a " stand-alone" model of the entire feedwater train. The validity of a semi-empirical correlation of this nature was demonstrated in Raf. 4. Figure VII.1 shows the derived semi-empirical correlation for temperature decay of feedwater exiting LP heaters 14. A limit to the total MFW that can be injected into the SGs under runaway MFW conditions does exist because of the low rate (75 kg/s (5.94 x 105 lb/h)) at which makeup liquid from the condensate storage tanks can be supplied to the l condenser /hotwell inventory to replenish the liquid drawn out at the excessive rates by the feedwater pumps. Once the inventory in the condenser /hotwells is dapleted, the feedwater pumps will trip on low-suction precsure. For the situation in which runaway MFW is supplied to both SGs, the feedwater flow will ba ~750 kg/s (5.94 x 106 lb/h) per SG during the runaway-feedwater portion of the transient. Starting with an initial condenser /hotwell inventory of 290000 kg, it is easily determined that the total duration of excess feedwater is only ~200 s. For the situation in which runaway MFW is supplied to one SG, the feedwater flow to that SG will increase to ~1000 kg/s (7.92 x 106 lb/h). At this rate, the condenser /hotwell inventory will be depleted in approximately f 300 s. Once the pumps trip, the liquid in the feedwater train and in the SGs I will stagnate. In reality, the feedwater pumps may actually trip earlier because of liquid filling the steam lines. If the liquid in the steam lines reaches the steam-extraction line for the MFW pump turbines, it may destroy the turbines and 57
4 s J ( u .a. - - 1 -- , - ~ 440-420-400-G
! 380-3
[ 360-E e 340-l 320 - 4 ^ 300-280 ' ' 0 Oh 0.'4 0.6 0.8 $ d t4 1.6 Js u u Fig. VII.1. 3 n 11 id exiting loepressure - .he es 4 and t 58
prsysnt furthar purping. Although an attsspt to model this effect was made, uncertainties in the location of the extraction line and the elevation at that i point probably would yield an unrealistic answer. The pumps were predicted always to trip on low suction pressure rather than on high liquid content in the extraction steamline. To run the third runaway-feedwater transient, the auxiliary feedwater flow j was increased f rom 320 kg/s to 400 kg/s (2.53 to 3.17 x 106 lb/h) per SG. (The AFW pumps actually can pump more than 400 kg/s; however, the plant operators are instructed not to exceed this value to prevent pump cavitation.)
, Subsections A, B, and C describe in more detail the transient sequences and parameters necessary to evaluate the PTS consequences from a thermal-hydraulic standpoint. Emphasis has been placed upon identifying the major energy transfers into and out of the primary fluid; these ultimately determined the downcomer temperature history. Details concerning the behavior of various plant components were relegated to the appendixes. Subsection D compares the three runaway-feedwater transients in an effort to identify similar and dissimilar behavior. In Section X the sensitivity of these results to major model uncertainties is presented.
A. Runaway MFW to Two SGs from FP This transient was initiated by a reactor / turbine trip from FP at 0 s with an assumed failure of both MFRVs to close. Table VII.A.1 tabulates the sequence of events that occurred during this transient. For the convenience of discussion, the downcomer temperature history has been divided into three phases (see Fig. VII.A.1). Figures VII.A.2 and VII.A.3 l summarize the non-convective energy transfers into and out of the primary fluid during these phases. As a reminder, a positive energy transfer indicates energy transferred into the primary fluid, whereas a negative energy transfer indicates j energy removed from the primary fluid. The first phase (0 - 283 s) shows a rapid decrease in downcomer temperature. The initial 10-K (18 F) temperature drop that occurred between 0 and 60 s is the normal temperature decrease that occurs when the reactor scrams. l The significant decrease in core thermal power caused the AT between the primary , cud secondary sides of the SGs to reduce to a much smaller value that still l i permitted dissipation of the decay heat. The energy removed from the primary fluid during this interim was ~22 GW-s per SG. At 60 s after the scram, the relatively cooler liquid that was in the feedwater pipes downstream of the HP i 59
. . _ . ._, _ . , _ . . - . _ _ - _ , - . _ - _-,._m-_-__,,,-- -_. _ -,,-.-_. ,. - , _,__ _ _.-. _ ,,..-,, ,--
TABLE VII.A.1 RUNAWAY MFW TO TWO SGs SEQUENCE OF EVENTSa Time (s) Event j i 0 Reactor / Turbine trip a) MFRVs failed open , b) ADVs and TBVs opened on " quick-open" logic ' 3 Pressurizer backup heaters trip on following low primary pressure 26 SIAS 33 Pressurizer heaters trip on following low-low level in pressurizer 52 High level (+30 in.) in SGs 56 RCPs tripped off by operator 30 s after SIAS 60 ADVs closed 120 SGs completely liquid full I 129 HPI began 218 MFW pumps trip off following loss of liquid inventory in condenser /hotwells 221 TBVs closed l t 490 HPI ended 630 Pressurizer level recovered; proportional heaters reactivated 4 2131 Proportional heaters trip off following high primary pressure 2310 PORVs opened following high primary pressure 4100 Primary-side temperature exceeded 552.6 K (535 F); 1 ADVs and TBVs reopened 7200 Calculation terminated i l "These transients assumed multiple operator / equipment failures. l 60
590 -
, -600 560--
3 -500 530-2 500- g 1 400 } 470 - { e a 440-e 1. Runaway main feedwater to both SGs h 4
~
- 2. Reheating following termination of excess 410 ~ feedwater 32
- 3. Quasi-equilibrium reached - slow boiling g in SGs _
3
-- -200 350- NOTE: These transients -
assumed multiple operator / equipment 320-f ailures. See TABLE H. . ,og 290 , , . . . . i 0 1000 2000 3000 4000 5000 60C0 7000 8000 Time (s) Fig. VII.A.l. Downcomer temperature during runaway main feedwater to two SGs. 10 0 - , , , , NOTE: These transients CORE 73_ ossumed multiple operator e pment f ailures. ee ABLE H. w_ S B 25- _ 0- A -
=
6 r ' i PUMP 4 NFT
/
~
_75- , SG A SG B
-10 0 - .
0 10 0 200 300 400 SCO 600 700 800 900 1000 Time (s) Fig. VII.A.2. Summary of non-convective energy transfers that occurred during runaway main feedwater to two SGs (0 - 1000 s). 61
heaters was swept into the riser region of the SGs, pushing the hotter liquid in the riser region into the steam-volume region above the tubes. The effective lower secondary-side temperature began to extract energy from the primary side at a rate of ~200 MW per SG. At 218 s, the MFW pumps tripped on low suction pressure because of depletion of liquid inventory in the condenser /hotwells. At this point, the liquid in the riser could no longer be replenished with cooler liquid. The riser region stagnated and quickly approached thermal equilibrium with the primary liquid temperature. The energy transferred to each SG decreased to ~15 MW. This is indicated by a decrease in the slope of the SG snergy curve at ~220 s (see Figs. VII.A.2 and VII. A.3). However, the thermal power produced by the decay heat was adding energy to the primary liquid at a rate of ~75 MW. ~As a result, the primary liquid began to heat again. The downcomer liquid temperature reached a minimum temperature of 477.5 K (399.80F) et 283 s. To aid the reader in understanding how the downcomer temperature reached a einimum value of 477.5 K (399.80F), an energy balance corresponding to the time of minimum downcomer temperature follows. The equations used in the following calculations were presented in Section VI and will not be repeated. A summary of the non-convective energy terms corresponding to the time of Einimum downcomer temperature (283 s) is tabulated in Table VII.A.2. The energy TABLE VII.A.2 ENERGY BALANCE FOR NON-CONVECTIVE TERMS AT TIME OF MINIMUM TEMPERATURE FOR RUNAWAY MAIN FEEDWATER TO TWO SGs I Source Energy (GW-s)a Decay heat +33.3 Primary-side slabs +7.3 RCPs +1.0 SG A -62.1 SG B -62.4 Net -82.9 l I l a + = energy addition to primary fluid.
- = energy removal from primary fluid.
62
convected into (positive) or out of (negative) the primary fluid during this same phase is tabulated in Table VII.A.3. Using Eq. (VI-5), the specific internal energy of the primary fluid is calculated to be: (219 x 103 )(1332 x 103 ) + 0.940 x 109 - 82.9 x 109 (219 x 103 + 7966) u = 924.1 S . kg Using the thermodynamic tables in Ref. 3, the bulk temperature corresponding to the above specific internal energy is 489.5 K (421.3 F). This corresponds to a bulk temperature drop of T bulk = (573.0 - 489.5) = 83.5 K (150.3 F). TABLE VII.A.3 ENERGY BALANCE FOR CONVECTIVE TERMS AT TIME OF MINIMUM TEMPERATURE FOR RUNAWAY MAIN FEEDWATER TO TWO SGs Source M (kg) h, (kJ/kg) Energy (GW-s)a Charging Flow
@ 299.8 K (80 F) +2221 111.9 +0.249 HPI Flow
@ 322.0 K (120 F) +564 204.5 +0.115
@ 302.6 K (85 F) +4270 123.5 +0.527 l @ 285.9 K (55 F) +911 53.5 +0.049 l
PORVs Total +7966 +0.940 l t ! * + = energy addition to primary fluid
- = energy removal from primary fluid.
l l
- 63 l
l
Of the 83.5-K temperature drop, 8.8 K (15.8'F) was produced by the convective cnergy transfers. TRAC predicted the downcomer temperature at 283 s to be 477.5 K (399.8 F), which is 12.0 K (21.6 F) below the bulk-fluid temperature of the primary system. This additional AT results from spatial variations of the liquid temperature within the primary system. In particular, the temperature difference between the hot and cold legs was ~25 K (44 F). If the bulk temperature is calculated as the arithmetic average of the cold-leg and hot-leg temperatures (477.5 K and 502.0 K respectively), a bulk temperature of 489.8 K (422.00 F) would be predicted. This is in excellent agreement with the previous analysis. Note that for this particular transient the decrease in downcomer temperature at the time of minimum downcomer temperature (ST = 559.0 - 477.5 K = 81.5 K) was almost equal to the decrease in the bulk-fluid temperature. This can occur only if the spatial variations that existed at the time of the initial steady state resembled the spatial variations tha' existed at the time of Einimum downcomer temperature. This was apparently the case for this transient because of the fortuitous situation that required a cold-leg-to-hot-leg temperature rise of ~25 K (~44 F) to drive the natural circulation flow following the trip of the reactor coolant pumps (RCPs). At the initial steady ctate of the plant (with the RCPs running), the cold-leg-to-hot-leg temperature rise was ~26 K (46.8 F). Hence, the spatial variations were comparable. Based upon this bulk-temperature analysis, it was determined that ~90% of the temperature decrease observed in the downcomer annulus was produced by non-convective energy transfers via the SGs, and the remaining ~10% produced by convective energy transfers associated with the injection of HPI and charging flow. To a first approximation, the spatial variations that existed at the time of minimum downcomer temperature were nearly equal to the spatial variations that existed at the initial steady state. The cooldown merely shifted both the l hot-leg and cold-leg temperatures downward by ~83.5 K (150.3 F). Phase 2 (283 - 4800 s) shows a relatively slow heatup of the primary fluid following the trip of the MFW pumps. As the primary temperature increased, cnergy was continually being transferred from the primary into the secondary. The stagnant liquid in the SGs began to heat up until it reached the saturation temperature corresponding to 6.2 MPa (900 psia), the pressure setpoint of the j TBV. The primary temperature leveled off at a small AT above the saturation l temperature of the liquid remaining in the SGs. A slow boiling process ' ben t 64
began (Phase 3). The small amount of steam being produced in the secondary side of the SGs was vented by both the ADVs and the TBVs. (The control on the ADVs cnd TBVs is designed to operate such that they open when the primary-side temperature exceeds 552.6 K (535 F).) The majority of the energy being added to ] the primary via the decay heat was removed through the SGs at a quasi-static l rate of ~11 MW. The balance of the decay-heat energy was removed by convective mass transfer (that is, cold fluid entering through the charging-flow system and hot fluid leaving through the PORVs). This accounts for approximately 8 MW of cooling capability. Figure VII.A.4 shows the primary-system pressure history for this transient. The primary system repressurized to the PORV setpoint because the operator failed to turn off the charging-flow pumps once the pressurizer refilled to its normal level. This caused the PORVs to cycle. The HPI and charging flow history produced by this pressure history is shown in Fig. VII.A.5. The HPI pumps stopped injecting fluid into the primary after the t
- primary pressure exceeded the HPI pump dead-head pressure of 8.8 MPa (1270 psig).
As shown in Fig. VII.A.2, both SGs removed equal amounts of energy from the primary fluid throughout the transient. Because of this symmetric cooldown , produced by overfeed to both SGs, neither primary loop stagnated (see Figs. VII.A.6 and VII.A.7). This allowed the primary fluid to remain uniform at similar spatial points within the primary loops as demonstrated by the liquid temperatures upstream of the HPI/ charging flow injection ports (see Fig. VII.A.8). I The uncertainties associated with the downcomer temperature will be [ discussed in Sec. X. l B. Runaway MFW to One SG from FP l This transient was initiated by a reactor / turbine trip from FP at 0 s with in assumed failure of one MFRV to close. Table VII.B.1 tabulates the sequence cf events that occurred during this transient. l i i For convenience of discussion, the downcomer temperature history has been divided into five phases (see Fig. VII.B.1). Figures VII.B.2 and VII.B.3 f summarize the non-convective energy transfers into (positive) and out of (negative) the primary fluid during these phases and will be referred to f throughout this discussion. 65
- - - - - ~ , _ - . - - ,- - - _ -._ _.. ---..-.-_ -. . _ . - _ , - .
350 , , , , , 300-CORE 250-o / 200 - / - 150 - NOTE: These transients y* assumed multiple OPerotor e ent 100-S f ailures. ee LE H. h So SLtB
] 0,
['- - I -. . P T ,,,,_ .. _ # - -* *** i - 1 50 -- ,
. . * " ' _ . .NET. . ( t PUMP
-10 0 - ~%
SG A _i3o _
-200 . . , , , .
0 1000 2000 3000 4000 6000 6000 7000 8000 Time (s) Fig. VII.A.3. Summary of non-convective energy transfers that occurred during runaway main feedwater to two SGs (0 - 7200 s). 18 . , , , , ,
- 2450 16 - [
-2100 12 - -1750
? )- ?
h5 'O ~ -
~
-i400 $
8 3 8 s- - E -
-1050
[ 6-
- 700 4-NOTE: These transients
~
ossumed multiple 2-Perotor e pmen t - 350 f ailures, ee ABLE H. 0 , , . . , . 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VII.A.4. Primary-system pressure during runaway main feedwater to two SGs. 66 t L
4 4 began (Phase 3). The small amount of steam being produced in the secondary side of the SGs was vented by both the ADVs and the TBVs. (The control on the ADVs cnd TBVs is designed to operate such that they open when the primary-side temperature exceeds 552.6 K (535 F).) The majority of the energy being added to the primary via the decay heat was removed through the SGs at a quasi-static ) rate of ~11 MW. The balance of the decay-heat energy was removed by convective mass transfer (that is, cold fluid entering through the charging-flow system and hot fluid leaving through the PORVs). This accounts for approximately 8 MW of cooling capability. Figure VII.A.4 shows the primary-system pressure history for this transient. The primary system repressurized to the PORV setpoint because the cperator failed to turn off the charging-flow pumps once the pressurizer refilled to its normal level. This caused the PORVs to cycle. The HPI and charging flow history produced by this pressure history is shown in I Fig. VII.A.S. The HPI pumps stopped injecting fluid into the primary after the ) primary pressure exceeded the HPI pump dead-head pressure of 8.8 MPa , (1270 psig). As shown in Fig. VII.A.2, both SGs removed equal amounts of energy from i the primary fluid throughout the transient. Because of this symmetric cooldown , ] produced by overfeed to both SGs, neither primary loop stagnated (see i Figs. VII.A.6 and VII.A.7). This allowed the primary fluid to remain uniform at cimilar spatial points within the primary loops as demonstrated by the liquid j temperatures upstream of the HPI/ charging flow injection ports (see
; Fig. VII.A.8).
The uncertainties associated with the downconer temperature will be
- discussed in Sec. X.
I ! B. Runaway NFW to One SG from FP This transient was initiated by a reactor / turbine trip from FP at 0 s with cn assumed failure of one MFRV to close. Table VII.B.1 tabulates the sequence of events that occurred during this transient. For convenience of discussion, the downcomer temperature history has been , divided into five phases (see Fig. VII.B.1). Figures VII.B.2 and VII.B.3 summarize the non-convective energy transfers into (positive) and out of
- (negative) the primary fluid during these phases and will be referred to throughout this discussion.
I 65 i i
. - . - - _.y.___r,_y,,._y_m c,.._._.y.,.-..,_.,.._.-__m_,y,,,.y, -, , y _y_7-_-.%,.y_,_~_m,-y, _ . , _ nv._
l 350 , , 100 CORE 250- /
/
200-150 - NOTE: These transients y, assumed multiple operator / equipment
,on _ ~ i 8 f ailures. See TABLE H.
h 50 SLAB y _... ..+ o, _ _ i .. . .-. .TT- 1 _ : _
'[ l 50 -
... e ,, .- I PUMP
. ner
-iOO - --
_ iso _ so A - SG B
-200 . , . . . .
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VII.A.3. Summary of non-convective energy transfers that occurred during runaway main feedwater to two SGs (0 - 7200 s). 18 . , , i .
-2450 16 - ! -
-2100 g4 .. -
12 - -1750 2 )- ? ( 5 0~
-- woo $
2 3 s- - 8 g -1050 g g. 6 --
- 700 j 4~ NOTE: These transients assumed multiple 1 2-perator / equipment -350 !
f ailures. See TABLE H. 0 , , . . . . 0 0 1000 2000 3000 4000 5000 6000 7000 5000 Time (s) Fig. VII.A.4. Primary-system pressure during runaway main feedwater to two SGs. 66
20 , , 1s-- -40 HPI FLOW Pte COLO EG is -. ------ --- cwAsome now Pte coto uG (LCCP5 A2 Ah3 81) ---3$ 14 -- -30 k' 12 - k NOTE: These transients assumed multiple
-25 e 3 ,
6 10 -
-f operator / equipment -
g a f ailures. See TABLE H. -20 ,
$ s- -
0 3 15 6-
-m 2-- -5 0, , , , , . , -0 0 1000 2000 3000 4000 5000 6000 7000 8000 rme (s)
Fig. VII.A.S. HPI/ charging flow history during runaway main feedwater to two SGs. 500 , , ,
. -1050 450 -
LDOP A1(00M8 25)
---------- LOOP A2 (CopP 45) 400-~ ^ ' 00ICU"' #
- LOOP B2 (CCwP 35) l
~
350-_ - 750 300-v *
- -600 R R 250-
~
Q [ N N
$ 200- \ $
Am m.;
~
50- _ -300 i 20- NOTE: These transien t s assumed multiple operator / equipment - 50 33 f f ailures. See TABLE H. 0- . . . . . . . -O O 1000 2000 3000 4000 5000 6000 7000 8000 ree (s) Fig. VII.A.6. Cold-leg flows during runaway main feedwater to two SGs. 67 l
10C0 i , , , , , 900 --2000 toor a (coup 20 Eco-. . -------- toor s (cowP 10 --1750 700 -_ 1500 6C0- - c, -1250 c, 1 3 g- sc0- - g g -
-1000 g S 400- -
h
, r.mmry, .
- 750
~
~5 200- l NOTE: These transients -
assumed multiple operator / equipment ~-250 10 0 -_ f ailures. See TABLE H. 0 , , . . , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VII.A.7. Hot-leg flows during runaway main feedwater to two SGs. 590 -
, , , , , ~600 560- -
530- tooP at
- -500
. . ... . . . . . too, ,2 LOOP 81 500 - Loop S2 -
12 F 30 e m-
-400 7y 440- -
)
-300 E
e 4 '0 - 380- - s
-200 333 NOTE: These transients _
ost,umed multiple operator / equipment 320-' f allures. See TABLE H. - 290 , . . . . C 1003 2000 3000 4000 5000 60C0 7000 8000 Time (s) Ff3. VII.A.8. Liquid tersperature upstream of HPI/ charging-flow ports during runaway main feedwater to two Sis. 68
$90 i , , , . , . 600 MO - -
530-- 1 --500 2 AW z 500- _ C-
~j - 400
$ 470 - 1. Runaway main feedwater to one SG ~
li 2. Reheating following termination of excess feedwater to SG A ]e 440- 3. Quasi-equilibrium ~ r . 4. AFW to both SCs -300 f e 410 -
- 5. Extrapolated -
,g Er 330 -
a
-200 350- ~
NOTE: These translents assumed multiple s20 - operator / equipment - f ailures. See TABLE H. -100 290 . . . . . . . 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VII.B.l. Downcomer temperature during runaway main feedwater to one SG. 10 0 . . . . , , . . NOTE: These transients assumed multiple operator equipmen t f ailures, se TABLE H. SLAB f 0 - 5>. ,% . i S B ! 6 .'.,, -'**',,,,,,,,,.........- l ',, ,,,,,,,,..... ! NET
- 10 0 - -
t j SG A ! - 15 0 . . . . . . . . . l 0 10 0 200 300 400 500 600 700 800 900 1000 Time (s) Fig. VII.B.2. Summary of non-convective energy transfers that occurred during runaway main feedwater to one SG (0 - 1000 s). 69
The first phase (0 - 363 s) shows a rapid decrease in the downcomer temperature. As with the transient discussed in the previous section, the initial 10-K temperature drop that occurred between 0 and 60 s was the normal temperature decrease that occurs when the reactor scrams. The energy removed by each SG during this interim was ~22 GW-s. Sixty seconds after the scram, the relatively cooler liquid that was in the feedwater pipes downstream of the HP heaters feeding SG A has been swept into the riser region of SG A. The effective lower secondary-side temperature in SG A began to extract energy at an average rate of ~260 MW. At 303 s, the MFW pumps tripped on low suction pressure because of depletion of the condenser /hotwell liquid inventory. (Unlike the runaway MFW to two SGs, failure of one MFRV to close produced a feedwater flow to the affected SG of ~1000 kg/s. This depletes the condenser /hotwell liquid inventory in ~300 s.) At this point, the liquid in the riser region of SG A was no longer being replenished with cooler liquid. The riser region stagnated and quickly approached thermal equilibrium with the primary liquid temperature. The energy transfer in SG A decreased to ~ 28 MW. This is indicated by a decrease in the slope of the SG A energy curve at 400 s (see Figs. VII.B.2 and VII.B.3). However, the thermal power produced by the decay heat was adding energy to the primary liquid at a rate of ~75 MW. As a result, the primary liquid began to heat again. The average downcomer liquid temperature went through a minimum temperature of 491.0 K (424 F) at 363 s. SG B received very little feedwater flow during Phase 1. The liquid temperature in SG B remained relatively constant at its initial temperature at the time of the reactor scram. As a result, SG B entered into a reverse heat-transfer mode (that is, the primary temperature dropped below the secondary temperature of SG B which caused energy to be transferred from the secondary l back into the primary) at 102 s. During the phase from 102 to 363 s (363 s corresponded to the time of minimum downcomer temperature), SG B added 3.2 GW-s of energy back into the primary fluid via reverse heat transfer. The phase of reverse heat transfer was marked by a positive slope in the energy curve for SG B. Hence, SG B removes a net energy of 22.5 GW-s from 0 to 363 s. The energy balance for the non-convective terms at the time of minimum downcomer temperature (363 s) is tabulated in Table VII.B.2. The energy convected into (positive) or out of (negative) the primary fluid during this same phase is tabulated in Table VII.B.3. 70
l TABLE VII.B.1 RUNAWAY MFW TO ONE SG SEQUENCE OF EVENTSa [ , Time (s) _ Event o O Reactor / Turbine trip a) MFRV to SG A failed to open b) ADVs and TBVs opened on " quick-open" logic 34 Pressurizer heaters tripped off following low-low level in pressurizer 90 ADVs closed on low primary temperature 123 SIAS 153 RCPs tripped off by operator 30 s after SIAS i 249 HPI began 303 MFW pumps tripped off following loss of liquid inventory in condenser /hotwells 325 TBVs closed 590 HPI ended 665 Pressurizer level recovered; proportional heaters reactivated 1550 TBVs reopened on high steamline pressure
'1700 Proportional heaters tripped off following high primary pressure 1850 PORVs opened following high primary pressure 2500 ADVs reopened on high primary-side temperature and TBVs began throttling on primary-side temperature rather than steamline pressure 4800 AFAS on low liquid inventory in SG B
( 4950 Primary-side temperature decreased below 552.6 K causing ADVs and TBVs to reclose t 5800 Calculation terminated "These transients assumed multiple operator / equipment failures. 71
TABLE VII.B.2 ENERGY BALA'. ICE FOR NON-CONVECTIVE TERMS AT TIME OF MINIMUM TEMPERATURE FOR RUNAWAY MAIN FEEDWATER TO ONE SG Source Energy (GW-s)a ; Decay heat . +37.6 Primary-side slabs +7.4 I RCPs +2.7 SG A -95.7 SG B -22.5 Net -70.5 a + = energy addition to primary fluid
- = energy removal from primary fluid.
Using Eq. (VI-5), the specific internal energy of the primary fluid is calculated to be:
, (219 x 10 3)(1332 x 103- ) + 1.05 x 109 - 70.5 x 109 (219 x 103 + 8510) ki u = 976.9 kg--
Using the thermodynamic tables in Ref. 3, the bulk temperature corresponding to this specific internal energy -is 501.0 K (442.2 F). This corresponds to a bulk temperature drop of < l i Tbulk = (573.0 - 501.0) = 72.0 K (129.6 F). Of the 72.0-K temperature drop, 8.5 K (15.3 F) was produced by the convective energy transfers. 72 1
.- . _ . _ ~ . _ _ _ .. . . _
TABLE VII.B.3 ENERGY BALANCE FOR CONVECTION TERMS AT TIME OF MINIMUM TEMPERATURE FOR RUNAWAY-MAIN-FEEDWATER TO ONE SG Source M (kg) hi ( W kg) Energy (GW-s)* Charging Flow 6 299.8 K (80 F) +4064 111.9 +0.455 HPI Flow 0 6 322.0 K (120 F) +564 204.5 +0.115 9 302.6 K (85 F) +3882 123.5 +0.479 PORVs 1 Total +8510 +1.049
* + = energy addition to primary fluid
- = energy removal from primary fluid.
TRAC predicted a downcomer temperature of 491.0 K (424.1 F) at this time. This is 10.0 K (18.0 F) below the bulk-fluid temperature of the primary system. Unlike the transient discussed in the previous section, however, the spatial variations in this transient are significantly different than the spatial variations that existed during the initial steady-state conditions. This is because of reverse heat transfer in the Loop-B SG creating a gravity head cpposing the normal flow direction and eventually leading to loop stagnation in l Loop B. Loop stagnation prevents mixing of the liquid in the stagnant loop with j liquid in the flowing loop. As a result, at the time of minimum temperature,
.the liquid temperature in the cold and hot legs of Loop B was relatively uniform et ~540 K (512 F), whereas the liquid temperatures in the cold and hot legs of Ltop A were 470.4 K (387.0 F) and 520.8 K (477.8 F), respectively.
[ Fcrtuitously, the full effect of the Loop-A cold-leg temperature of 470.4 K
.(387.0 F) entering the downcomer annulus was not felt. The density gradients that subsequently formed in the azimuthal direction around the downcomer annulus induced significantly higher azimuthal flow in the downcomer annulus. This allowed hot liquid on the stagnant-loop side of the vessel to mix with the 73 l
l
.._ _ _ __ a .._...,_.,___._._ _ __._-_._.,_ _ ._._._ _ _ _ ___. _ - _ _ _ _ _
s H colder liquid entering the vessel from the flowing loop. Hence, the average
- liquid temperature at the core elevation was 20.6 K (37.1 F) hotter than the flowing cold-leg (Loop A) temperature.
Based upon this analysis, it was determined that ~88% of the temperature decrease observed in the downconer annulus was produced by non-convective energy 1 4 l transfer to SG A, and the remaining ~12% produced by convective energy transfers associated with the injection of HPI and charging flow. Spatial variations that existed at the time of minimum downconer temperature were not similar to the spatial variations that existed during the initial steady state. Nevertheless, the downconer-temperature decrease (AT = 559.0 - 491.0 = 68.0 K (122 F)) was still. fairly close to the calculated bulk-temperature decrease of 72.0 K
- (129.6*F).
Phase 2 (363 - 3200 s) shows a relatively slow heatup of the primary fluid
- following the trip of the MFW pumps. This was similar to the heatup observed in the runaway MFW to two SGs discussed in the previous section except that the f ,
heatup that occurred in this transient has only one heat sink--SG A. The other [ SG cooled only slightly during the runaway-feedwater portion of the transient. As a result, the decay heat added to the primary fluid could be dissipated I through only one SG rather than two. Hence, the primary fluid heated up more rapidly for this case. After SG A was heated again to the saturation temperature corresponding to 6.2 MPa (900 psia), both SGs shared the heat load ! equally. The primary temperature leveled off at a small AT above the saturation temperature of the liquid remaining in the two SGs. A slow boiling process j began (Phase 3). As in the transient discussed in the previous section, the l primary fluid temperature during this phase exceeded 552.6 K (535 F). Both the ADVs and the TBVs reopened, which vented the steam being generated by the boiling process. Subsequently, about one-third of the decay heat was removed by each SG. The remaining one-third of the decay heat was removed by convective mass transfer associated with injecting cold charging flow into the primary system at a rate of 8.3 kg/s (6.59 x 104 lb/h) and rejecting, on an average, the l same mass flow rate through the PORVs with a much higher temperature. Because the initial mass inventory in SG B was depleted somewhat and was i not replenished during the runaway-feedwater portion of the transient, the slow boiling process that occurred in Phase 3 continued to boil the remaining liquid in SG B. At 4800 s, the level in SG B was finally low enough to activate AFW to both SGs. The continuous addition of cold 277.6 K (40 F) liquid to each of the 74 i
- > - - - e y.- .-T-.4.-r---,,,m7m_g. -M vw-M44--N'M-*"h*F-"'Pl'*TeP &Wy7a.e>m--e.g. -#yggums,.apg,c-m,-.m- m,g.y-y9yv,.-,,mm-.e-weres.e ,vgm+%g,p-w-.y - p qwi,g y p3 +3 p->yv+ M y, ym g -n.spe--ge g w
J . 'SGs resulted in a continuous reduction of the secondary-side-heat-sink
!' temperature. This, in turn, produced a decrease in the primary fluid l temperature (Phase 4). Once the primary-side temperature decreased below 4 552.6 K (535 F), both the ADVs and the TBVs reclosed.
i The calculation was terminated at 5800 s. However, it was anticipated that the primary-fluid temperature would continue to decrease at approximately the same rate observed in Period 4 for the interim from 5800 to 7200 s (Phase 5). Therefore, the temperature history shown in Fig. VII.B.1 for Phase 5 is an extrapolation, not a TRAC calculation. l Figure VII.B.4 shows the primary-system pressure history for this j transient. The primary system repressurized to the PORV setpoint because of cperator failure to turn off the charging-flow pumps once the pressurizer refilled to its normal level. This caused the PORVs to cycle. (For reasons ] . unknown, the TRCGRF file for the pressure of the primary system failed to store the~ pressure results properly during the time period of 2500 to 3200 s. Figure [ VII.B.4 indicates that PORVs stopped cycling during this interim. This, of course, is not the case. According to the TRCOUT files, the PORVs continued to cycle at the frequency previously established prior to entering this phase.) The i ! HPI and charging-flow history produced by this pressure history is shown in Fig. VII.B.5. The HPI pumps stopped injecting fluid into the primary system after the primary pressure exceeded the HPI pump dead-head pressure of 8.8 MPa (1270 psig). j Figures VII.B.6 and VII.B.7 show the hot- and cold-leg flows for this transient. Because of the asymmetric cooldown of the primary system during the l f runaway-feedwater portion of the transient, Loop B stagnated (or had very low flow) for ~500 s. As previously mentioned in this section, Loop B stagnated because a gravity head formed in the U-tubes of the SG B during the time that t l the SG entered into a reverse heat-transfer mode (for a complete explanation, f see Section VI). Flow began in Loop B after the primary liquid reheated to a temperature- above the secondary-side-liquid temperature of SG B. This re-established the normal heat-transfer mode - primary to secondary side. The gravity head produced under these conditions was acting in the same direction as ( [ the driving force produced in the vessel. Hence, a natural circulation flow of
~ 150 kg/s (1.19 x 10 6lb/h) was established in both loops during Phase 3.
i I ! 75
~y e , vr - ym y - y r-.,wy,-y.y-.m,w g -w-3 ,-----,wm%,y.,%,,, .,%w,-ew., _ , , , . _ , _ . _ ,
,,, m.e_.-ym.,-,.-.-p. mmv_--,
300 , , , , , , , , , , NOTE: These transients 22- ossumed multiple CORE operator / equipment 200- f ailures. See TABLE II. o - 15 0 - 10 0 - - 73 g. SLAB k _ g 0,<: : _ g ,,,...p -:-- : : 1-0 **** 50- ......e - SG B
- 10 0 - -
- 15 0 - -
-200- SG A
~
-250 , , , , , , , , , , .
0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000 5500 6000 Time (s) Fig. VII.B.3.
. Summary of a non-convective energy transfers that occurred during runaway main feedwater to one SG (0 - 5800 s).
18 , , , i
- - 2450 16 - p h h -
- 210 0
- u. -
u-- -- U50 d 10 - -
- 400 v o,
i e s & i s N 8- A { extrapolated g -1050 g s- - l
- - 70 0 4-NOTE: These transients ssumed multipl -350 2- operator / equ,eipment f ailures. See TABLE II.
0 , , . , . . , o j 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) l Fig. VII.B.4. Primary-system pressure during runaway main feedwater to one SG. l 76 i
20 . . . . . . 18-- --40 HPI Ft0W MR CCLD EG 16 -- ---------- CHARG;hG row MR COLD LEG (LOOPS A2 Ah3 8Q --35
~
~-30 NOTE: These transien t s 6
12 - assumed multiple operator / equipment
-25 3 o
f ailures. See TABLE H.
- o 10- -
-20 c
m g 2 8- - f . 33 6- -
- -10 4- .
2-
-~5 0 . . , . , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Trne (s)
Fig. VII.B.S. HPI/ charging-flow history during runaway main feedwater to one SG. 1000 . . , , . . .
--2000 900--
i Loop A (COMP 20 BGO-- it ------~ ~ -- LOOP 8 (Coup 1Q --1750 700 - --1500 _{ 6 600- [ NOTE: These transients ossumed rnultiple
-1250 3
o operator / equipment ~ d 500~ i f allures. See -TABLE H.
-1000 d
g m 3 400- j j
- - 750
'i i u i .qf4T
-: / -500 200-l *
/
1 l
,co _-l j _-250
;': ...l/
O
. . . . . . . 0
.0 1000 2000 3000 4000 5000 6000 7000 8000 Trne (s)
Fig. VII.B.6. Hot-leg flows during runaway main feedwater to one SG. 77 l
500 , , , , , , ,
-1050 450- -
LOOP At(COMP 25) [ ---------- LOOP A2 (COMP 45) 400-' q ""-~ - Loor s:(cove is) -900
- LOOP B2 (COWP 35) j
~
350-- NOTE: These transients -M0 qm ossumed multiple 6 300-operator / equipment f allures. See TABLE H.
.-600 5 m g 250- -
g R a j -450 E 200-- 2
--al.u... _
15 0 - ,# -
- -300 WO-
- 15 0 50-
,j 0
d' 0 0 1000 2000 3dOO 4$00 SdOO 6dOO 7dOO 8000 Trne (s) Fig. VII.B.7. Cold-leg flows during riinaway main feedwater to one SG. 590 -
, , , , , -600 560- -
~
f" 530- wop ai
.. .... .. . . toop ,2
. . .- .. . LOOP 8i 500- - LOOP B2 -
? F 7
y co-
-400 7 3
0 I' { m- - L
# - -300
-{ 410 -
h 3 ?r 380-
- -200
~
3"~ NOTE: These transients assumed multiple 320- operator / equipment -
-SC f ailures. See TABLE H.
2so , , . , , , 1 0 1000 2000 3000 4000 5000 6000 7000 8000 Trne (s) Fig. VII.B.8. Liquid temperature upstream of HPI/ charging-flow ports during runaway main feedwater to one SG. 78
.. - _ m .
4 During the phase in which Loop B stagnated, the temperature upstream of the HPI/ charging flow ports remained relatively elevated at a temperature of
~540 K (512 F). In the other loop, the temperature decreased to a minimum value of 470 K (386 F) just before the time of minimum downcomer temperature. Once flow began in Loop B, the liquid in both loops began to mix in the vessel region and the system liquid temperature became more uniform (Phase 3).
c The uncertainties associated with the downcomer temperature for this transient will be discussed in Sec. X. C. Runaway ~ AFW to Two SGs from FP
-This transient was initiated by an unanticipated trip of both MFW pumps from full power. It was assumed that the AFW system would fail to start following the auxiliary feedwater actuation signal (AFAS). After 1200 s (20 min), the operator was able to- manually activate AFW to both SGs at its prescribed maximum flow rate of 25 kg/s (400 gpm). Furthermore, it was assumed that the operator, would secure AFW to both SGs 180 s (3 min) after the narrow-range level indication in either SG reached the +50-in. high-level alarm.
Table VII.C.1 tabulates the sequence of events that occurred during this transient. I For convenience of discussion, the downcomer temperature history has been divided into five phases (see Fig. VII.C.1). Figures VII.C.2 and VII.C.3 summarize the non-convective energy transfers into (positive) and out of (negative) the primary fluid during these regions and will be referred- to throughout this discussion. The first phase (0 - 34.7 s) shows a slight temperature increase prior to the reactor / turbine trip at 34.7 s. This temperature increase was produced by
.the degradation of the heat-load capacity of each SG following the loss of MFW l
flow. Because the reactor power was programmed not to change during this interim, a net energy of 0.9 GW-s was transferred into the primary fluid,
- causing the primary temperature to increase a few degrees. The initial SG t
inventory of ~63000 kg per SG was reduced by 30% during this phase. (As described in Sec. III, the reactor is supposed to trip if the narrow-range level indication in either SG drops to -50 in. The mass reduction necessary to (
. produce this level drop corresponds to ~15% of the initial mass, not 30%.
i Hence, as suggested by the loss-of-load comparison made in Sec. V, it is 1 l suspected that the narrow-range level indication overpredicted the SG level
.during this transient. Based upon mass inventory, it is speculated that the i
l-79
TABLE VII. Col RUNAWAY AUXILIARY FEEDWATER TO TWO SGs FOLLOWING LOSS OF MAIN FEEDWATER SEQUENCE OF EVENTSa Time (s) Event . O Main-feedwater pumps tripped 35 Reactor / turbine tripped on low narrow-range SG 1evel and ADVs and TBVs opened on " quick-open" logic 35 AFAS -- AFW system failed to start
.76 Pressurizer heaters tripped off following low-low level in pressurizer 532 Pressurizer level recovered; proportional heaters reactivated 800 Both SGs dried out a) primary fluid began to heat up b) ADVs and TBVs transferred to full-open 864 SGIS on low SG pressure; MSIVs and MFRVs closed isolating TBV 919 High primary pressure -- PORVs began to cycle 1200 Operator was able to turn on the AFW system 1396 SIAS on low primary pressure 1426 RCPs tripped off by operator 30 s af ter SIAS 1469 ADVs closed on low average reactor temperature 3631 High primary pressure -- PORVs began to cycle to relieve unrestricted charging flow 6590 High narrow-range SG level (+50 in.)
l 6770 Operator terminated delivery of AFW to both SGs 180 s after high SG 1evel , j 7200 Calculation terminated aThese transients assumed multiple operator / equipment failures. 80
590 i
-500
-s
$60-) /
1 2 3 w
- 500 530- u.
SCs dried T~ %m 500-U T 470- 1. Loss of main feedvater before reactor scram 4o0 E-5
} 2. Loss of main feedwater after reactor scram y
- 3. SCs dried out followed by AFW initiation g 1 '43~ Q.
E 4. Maximum AFW delivered 4 . 5. AFW terminated -300 3 5 410 - h= 380-5
- --200 330' NOTE: These transients assumed multiple operator / equipment 320- f ailures. See TABLE H. -
. ,no 290 , , , . , . .
0 1000 2000 3000 4000 5000 ts000 7000 8000 Time (s) Fig. V11.C.l. Downcomer temperature during runaway AFW to two SGs. 200 , , , 15 0 - _, 10 0 - CORE I 50- FUMP t 9 i- d - (. K 0- ,
-~ --
SLAs p - - - - P
',..... ..... ..... .... ... . . ..... . . ..- + - -- -
O _50_ } NET
- - 10 0 -
NOTE: These transients sc A & a assumed multiple
-'50 ~ operotor / equipmenf l f ailures. See TABLE H.
l l
-200 , , , , . . . .
0 10 0 200 300 400 600 600 700 800 900 1000 l Time (s) Fig. VII.C.2. Summary of non-convective energy transfers that occurred l ., during runaway AFW to two SGs (0 - 1000 s). 81 i l ?
i reactor should have tripped at ~17 s, rather than at 34.7 s. The effect of this possible anomaly upon this transient will be discussed intermittently throughout this section if the effect is known.) As previously mentioned, the reactor / turbine tripped at 34.7 s because of low ~SG narrow-range level indication. The primary liquid temperature quickly dropped to a quasi-static equilibrium temperature a few degrees above the secondary-side-liquid temperature (Phase 2). The decay heat produced by the reactor during this phase was dissipated equally by both SGs at a rate of ~40 MW per SG; this' heat continued the boiling process in each SG. This continued to deplete the liquid inventory in each SG, and subsequently led to AFAS at 35.5 s. .(AFAS occurs whenever the wide-range level indication in either SG drops below
-170 in. The mass reduction necessary to produce this level decrease corresponds to ~30%. Because the mass reduction at 35.5 s corresponded to ~30%,
it was presumed that the wide-range level indication was correct. Whether the wide-range level indication was correct or not, however, is academic in this particular transient. The AFW system was assumed to fail and not deliver feedwater to either SG.) At ~800 s, the entire liquid inventory in each SG had been completely boiled away. (If the reactor had tripped at 17 s, the mass remaining in the SGs at the time of reactor / turbine trip would have required ~1100 s to boil away. This would have displaced the point in time at which the downcomer temperature began to increase. Rather than -at 800 s, the increase would have begun at 1100 s (see Phase 3 on Fig. VII.C.1). Nevertheless, the downcomer temperature would have remained -just as constant during an " extended" Phase 2 as in the case that was run.) The average primary temperature during Phase 2 was higher than 552.6 K (5350F) which caused both the ADVs and TBVs to be open. Together, they vented all the steam that was being produced. After the SG liquid inventory depleted, the heat-load capacity of each SG decreased to less than 1 MW. The decay heat produced by the reactor could no longer be dissipated from the primary fluid. The temperature began to rise sharply (Phase 3). This caused the ADVs and the TBVs to open fully, which caused each of the SGs to depressurize. As a result, SGIS occurred at 864 s. The MFRVs and MSIVs closed and isolated the SGs from the TBVs. As specified by ORNL, the operator was able to start the AFW system manually at 1200 s. The initial surge of cold AFW entering the SGs vaporized rapidly. This removed 15.4 GW-s of energy from the primary fluid over the next 82
300 s. The injection of cold charging flow over the same period of time resulted in a further decrease in the temperature of the primary fluid. The net result was a rapid temperature decrease of 22.5 K (40.50F) (see beginning of Phase 4). The average primary temperature dropped below 552.6 K (5350F), causing the control system to close the ADVs. This bottled up both SGs for the remainder of the transient. The continued addition of cold AFW to both SGs resulted in each SG removing energy from the primary fluid at an average rate of
~19 MW. This energy did not boil the AFW. Rather, the energy was added as censible heat to the liquid causing its temperature to increase. The increase in the secondary-side-liquid temperature, however, occurred for only a short period of ti.ne. The secondary-side-liquid temperature peaked at ~540 K (5120F) at ~1600 s. The rate at which energy was being added to the secondary-side liquid as sensible heat was offset at this time by the continued addition of cold AFW. The net result was an increasing liquid inventory in each SG with a modestly decreasing liquid temperature.
As observed in Fig. VII.C.3, the energy that was being added to the
- primary fluid via the decay heat, the slabs, and the pumps was balanced by the l energy extracted through the two SGs for all practical purposes. The net non-convective. energy transfer from the primary fluid during the interim from -
1600 to 5800 s was essentially zero. The primary fluid temperature, nevertheless, was decreasing slowly during this phase. This was because-of the convective energy transfer that was occurring with the injection of cold l charging. flow into the primary system and, during the second half of this transient because of the hot liquid leaving the primary system via the cycling PORVs. The energy convected from the primary via hot liquid leaving through the f PORVs and cold liquid entering through the charging flow system was ~10 MW. On an average, the primary fluid temperature decreased at a rate of ~32 K/h (580F/h) over Phase 4 because of convective cooling. (The temperature oscillations that occurred during the time interval of 5500 to 6200 s are believed to be produced by pressure anomalies. They are presumed not to be l real. The pressure anomalies will be discussed later in this section.) If the f operator had throttled the charging flow at the time of level recovery in the { pressurizer, the primary liquid temperature would have remained constant during ( Phase 4. I 83 l I - . - - _ - . - - - - , , - - - - - . - _ . . . - - . - . - . - - - - .- -. ,------ - -- - .
At 6590 s, the AFW system had refilled the SGs to the +50-in. level. Per
.the transient: specifications, the operator turned off the AFW system 180 s
~ (3 min) later. The energy that was being dissipated through the 'SGs began to heat up the liquid in the SGs. As the secondary-side temperature increased, the heat-transfer rate to the SGs decreased. The decay heat from the core finally exceeded. both the rate at which energy was being removed via the steam generators and the convective energy transfers associated with the charging flow. The primary fluid began to heat again (Phase 5) 100 s after the operator
- - turned off the AFW.
The energy balance for the non-convective terms at 7200 s is tabulated in Table VII.C.2. The energy convected into (positive) or out of (negative) the
- . primary fluid during this same phase is tabulated in Table VII.C.3.
Using Eq. (VI-5), the specific internal energy of the primary fluid is calculated to be: 3
, , (219 x.10 )(1332 x 103) - 3.862 x 109 - 34.4 x 109 (219 x 103 + 46615) u = 954.2 kJ ,
Eg l TABLE VII.C.2 l , ENERGY BALANCE FOR NON-CONVECTIVE TERMS AT 7200 s FOR RUNAWAY AUXILIARY FEEDWATER TO TWO SGs Source Energy (GW-s)a i i Decay heat +385.5 Primary-side slabs +20.2 j
+24.7 RCPs SG A -231.0 l
SG B -233.8 [ Net -14.4 f' t l l a + = energy addition to primary fluid i - = energy removal from primary fluid. 84
,Using the thermodynamic tables in Ref. 3, the bulk temperature c:rreshondingtothisspecific internal energy is 496.1 K (433.2 F). This c:rresponds to a bulk temperature drop of Thulk = (573.0 - 496.1) = 76.9 K (138.4 F).
l Of the 76.9-K temperature drop, 49.3 K (88.8 F) was produced by the convective snergy transfers. TRAC predicted a downcomer temperature of 498.0 K (436.7 F) at this time. This was 1.9 K (3.4 F) above the estimated bulk fluid temperature of the primary cystem. The spatial variations in the primary fluid temperature at this late time 'in the transient are not as large as those that existed early in the transient. The temperature difference between the hot- and cold-legs is only 14.4 K (25.9 F) at 7200 s. If the bulk temperature were calculated as the crithmetic averge of the cold-leg and hot-leg temperatures (492.2 K and 506.6 K, rsspectively), a bulk temperature of 499.4 K (439.2 F) would be predicted. This is good agreement with the foregoing analysis. TABLE VII.C.3 i ENERGY BALANCE FOR CONVECTION TERMS AT 7200 s FOR RUNAWAY AUXILIARY FEEDWATER TO TWO SGs I Source _ M_(kg) h, (kJ/kg) Energy (GW-s)a I Charging Flow
@ 299.8 K (80 F) +55074 111.9 +6.162 HPI Flow - - -
PORVs -8459 ~1185.0 b -10.024 Total +46615 -3.862 l a+ = energy addition to primary fluid
- = energy removal from primary fluid.
Estimate of average enthalpy of liquid entering pressurizer surge line while PORVs are cycling. 85
Note that-for this particular transient, the downconer temperature at 7200 s showed a decrease of only 61 K (109.8 F), whereas the bulk temperature was estimated to have decreased by 76.9 K (138.40 F). This resulted primarily from the reduction of the spatial variations. Rather than a hot-to-cold leg 0 temperature decrease of 26 K (46.8 F) which occurred at the initial steady-state condition, the hot-to-cold leg temperature difference at 7200 s was only 14.4 K (25.90 F). At the initial steady state, the downcomer temperature was 13 K (23.4 F) from the bulk temperature, but at 7200 s was only ~7 K (~13 F) from the bulk temperature. Hence, the downcomer temperature should decrease 6K less than the observed decrease in the bulk-fluid temperature. Based upon this bulk-temperature analysis, it was determined that ~36% of the temperature decrease observed in the downcomer annulus was produced by non-convective energy transfers via the SGs, and the remaining ~64% produced by convective energy transfers associated with the injection of HPI and charging flow. The spatial variations at the end of the transient (7200 s) were approximately one-half of the initial steady-state variations. Figure VII.C.4 shows the primary-system pressure history for this transient. The initial pressure drop observed during the time from 34.7 s to
~800 s was the result of the reactor / turbine trip at 34.7 s and the subsequent decrease in primary fluid temperature. Fluid contraction acconipanying the temperature decrease caused the pressurizer level to fall. This activated the
' ~ makeup / letdown systes, which began to inject makeup liquid at a rate of 7.30 kg/s (5.0 x 104 lb/h). This allowed the primary system pressure to regain nearly all its initial pressure value. At ~800 s, the SGs dried out and the primary-fluid temperature began to increase sharply. The decrease in pressurizer vapor volume resulting from increased fluid expansion accompanying the temperature increase caused the pressure to increase and, subsequently, reach the PORV setpoint. The PORVs began to cycle. The makeup / letdown system was removing liquid at a rate of 4 8.3 kg/s (6.6 x 10 lb/h) during this interim because of the high liquid level in the pressurizer. The PORVs continued to cycle until the operator was able to turn on the AFW to both SGs at 1200 s. The primary-fluid temperature began to decrease at this point, bringing the system pressure down with it. At ~1400 s, the primary-system pressure decreased to the SIAS. Thirty seconds later, the operator was assumed to have tripped the RCPs according to the transient i ! 86 l
F-500 ' 400 CORE _ 300- p p' 200
'/ NOTE: These transients assumed multiple -
4 y operator / equipment j , f allures. See TABLE 11. 10 0
& " sLAs f 9
- _k; ;
+ --
-10 0 - NF.T Sc g
'N'
-200 -
sc a [ w
-300 . . . . .
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VII.C.C. Summary of non-convective energy transfers that occurred during runaway AFW to two SGs (0 - 7200 s). is , ,
- 2450 16- ' ' $ h -
-7100 14 -
12 -- -1750 1 2 10 - - -1400
! 2 a
. s- .
5 - -1050 [ 6- ' - -100 4-NOTE: These transients ossumed multiple 33o operator / equipment 2 -- f ailures. See TABLE H. 0 , . . , , . 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VII.C.4. Primary-system pressure during runaway AFW to two SGs. 87
specifications. The- temperature decrease of the primary fluid slowed ! considerably following. the trip of the pumps. The pressure did not drop below the dead-head pressure of the HPI pumps during this interim. Consequently, no HPI flow was injected. However, the charging-flow system was permanently activated following SIAS. This caused the primary-system pressure to recover , for a second time. The system pressure reached the PORV setpoint at ~3600 s,
! whereupon the PORVs again started to cycle. I
! Because of unthrottled charging flow, the pressurizer completely filled with liquid at ~5500 s. The liquid began-to spill into the PORV/SRV header.
- TRAC calculated a change from a vertically stratified flow in the pressurizer region to a horizontally unstratified flow in the header which resulted in the
) condensation rate suddenly increasing by several orders of magnitude. This ! resulted in a large pressure decrease observed in Fig. VII.C.5 during the time i from 5500 s to 6200 s. The pressure decrease observed during this time interval is not believed to be real and, as previously mentioned, caused noticable changes in the downcomer temperature. These changes were the direct result of ) liquid being drawn into the pressurizer following the sudden decrease in primary pressure. This induced hot liquid located in the upper plenum above the hot-leg penetrations to be drawn into the hot legs of the loops, ultimately perturbing the loop flows already established. By perturbing the loop flows, the ratio of
, cold-leg flow to charging flow changed, resulting in the mixed fluid temperature t
entering the downcomer annulus to change. Driven by the two large pressure dips, the downconer temperature appeared to be oscillatory during this time period. However, these oscillations are also not believed to be real because i the pressure oscillations are not believed to be real. Following the I condensation of the majority of the vapor located in the PORV/SRV header, the pressure of primary system began to behave "normally" for the remainder of the l l transient. The HPI and charging-flow history produced by this pressure history [ are shown in Fig. VII.C.5. The HPI pumps did not inject any fluid into the primary system because the primary-system pressure never decreased below the ! HPI pump dead-head pressure. The charging flow for the first 1500 s of the l l , transient was being controlled by the pressurizer-level indications. Following i SIAS, .the charging pumps were permanently activated and the letdown valve completely closed. This increased the charging flow to 4.15 kg/s for the l j remainder of the transient. i i 88 I l l
r I i i i i As shown in Fig. VII.C.3, both SGs removed equal energy from the primary fluid throughout the transient. Because of this symmetric cooldown produced by uniform auxiliary flow to both SCs, neither loop stagnated (see Figs. VII.C.6 i cad VII.C.7). This allowed the primary fluid to remain uniform at similar t [ cpstial points within the primary loops as demonstrated by the liquid {- tccperatures upstream of the HPI/ charging-flow injection ports (see l
-Fig. VII.C.8).
l t . The uncertainties associated with the downcomer temperature will be discussed in Sec. X. D. Compari_ sons 1 ii Figure VII.D.1 compares the downconer temperature at the core elevation for the three runaway-feedwater transients studied. The two runaway-main-
- foedwater transients showed an immediate cooldown of the primary fluid.
1
, However, the cooldown was short-lived because of the limit to the cold feedwater I that can be swept into the SGs prior to the feedwater/ condensate pumps tripping cn low suction pressure. On the other hand, a runaway-auxiliary-feedwater f
j transient represented a long-term cooldown potential. The inventory in the i ccndensate storage tank from which AFW was drawn could supply cold liquid to the l SGs for potentially six hours. At a cooldown rate of 32 K/h (58*F/h), the f primary-liquid temperature would reach 381 K (226 F) before the inventory in the ccndensate storage tank was depleted. However, a cooldown rate of this magnitude is smaller than the cooldown rate prescribed for this plant during j .ncesal shutdown procedures. i From a comparison standpoint, the cooldown rates produced by the two runaway-main-feedwater transients were much more severe. Figures VII.D.2 and 4 VII.D.3 compare the energy-removal capabilities for the SGs (combined) for each l cf the three transients. (To make a valid comparison, the runaway-auxiliary- $ -fo:dwater data were modified to account for the reactor scram that occurred at l' 34.7 s. During the interim from 0 to 34.7 s, the energy removed by each SG was 46.8 GW-s. A total of 93.6 GW-s was subtracted from the sum of the energy [ rcmoved by the SGs during this interim, and the time base was shifted back by 34.7 s. Hence, the energy removal capabilities are being compared from the ti_me of_re_a_c_ tor scram, not the beginning of the transient.) As can be seen in Fig. VII.D.2, all three transients removed comparable caounts of energy from the primary fluid for the first 60 s following a reactor ceram. However, after the cold feedwater reached the riser region of the SGs, 89
o . . , , . . .
.- 20 7.5 - -
. wirtow pa com us .,3
---*--*--- CMaao#NB Ft0W PUt COLD EQ W A2 AM3 31) "
5- -
-m
^g 23-F i! -_3 p E !i 6 a 1: 3 2 0 -a g l
2.3 f . --: I U
--10
. . NOTE: These transients _ _,3
-7.5 assumed multiple -
operator / equipment
. f ailures. See TABLE H. - -20
-m . , , , , . .
0 1000 2000 3000 4000 SOOO 6000 7000 8000 Time (s) Fig. V11.C.S. HPI/ charging-flow history during runaway AFW to two SGs.
%0 i i .
- -1050 450-LOOP At(Coup 25)
..... . ...- LOOP A2 (Cou8 45)
* ""C ""5) 4 4c0-
- LOOP S2 (COMP SS) 350- ,
-m 300-O -600 O
3 r D Q 250- Q
" '50 h 2c0 15 0 -
'M y 300 10 0 - NOTE: These transients assumed multiple . . ,g operator / equipment 50 --
f allures. See TABLE H. ! 0 . . .
, , .o
! 0 1000 2000 3000 4000 6000 6000 7000 8000 Time (s) l Fig. V11.C.6. Cold-leg flows during runaway AFW to two SGs. 90
1000 . . . . . 900
-2000 LOOP A (C0W8 21) 800-- -- ------ war e (coe 's) -1750 700-, ,,339 600-3 -
1250 v R 500-
-1000 d, .
400-2 300- P - t -500
*"~
NOTE: These tronsients assumed multiple 250 10 0 -- operator / equipment f ailures. See TABLE H. 0 , , , . . . .0 0 1000 2000 J000 4000 b000 6000 7000 5000 Time (s) ) Fig. VII.C.7. Hot-leg flows during runaway AFW to two SGs. se0 -
, , , . . . . -600 sso-t -
g -
. ~500 800- .
g 8 war ni
-400 v
. .. .... . war ,1 - e 470 - ...-... m y ei g
- woe ez 440 -
-bGO h
}
4,0-3 se0-5 ,
- 200
*~ NOTE: These transients
~
assumed multiple 320 - operator / equipment ~ '00
~
f ailures. See TABLE H. soo , , , , . - em o 1000 2o0c 3000 4000 5000 so0o 7000 Time (s) Fig. VII.C.8. Liquid teinperature upstreain of HP1/chargind-flow Ports during runaway AFW to two SGs. 91
300 , , , i i '
.600
,i ,
-"~~~'
550- \' - _. -~-. t
.. .. ; p .. ****. ~~
...s,....
snn 500- *
, Y 't._ . ,,, .
3 -
-400 450-300 y 400' RUNAWAY uAmorttonAfte to TWo so *
......... . RUNAWAY.MalN.FTt0BAf!810 0Nt SG NUNAWAY. AURfuARY-FLLDWA'1R TO Two 30 .-200 .
350- - o 300 NOTE: These transients assumed multiple operator / equipment f ailures. See T ABLE II.
-o 250- . . . .
0 1000 2030 3000 4000 00'03 6600 7600 8000 Time (s) Fig. VII.D.I. Comparison of downcomer temperature for the three runaway-feedwater transients. 50 , , , . . . . . RUNAWAY.WAN.Fil3 TAI (4 70 TWO SC
..-....... RUNAWAY.Wan-rttow Aftt TO ONE SG
- - - PUNAWAY. AUIluaRY-FttDWA*tR 4 Two 50 0- -
I
, s. ,
-10 0 - 5
- 15 0 '
NOTE: These fransients ossumed multiple operator / equipment f ollures. See TABLE II.
-200 i 1 . 1 ; , .
0 90 0 200 300 400 SCO C00 700 600 900 1000 Time (s) Fig. VII.D.2. Comparison of the energy removed through the SGs (combined) for the three runaway-feedwater transients (0 - 1000 s). 92
m-200 , , , , , i hMANAT-MAN. FEE 0maite 10 two 5G
.......... RUNAWAY-MAM-filDE ATER 70 0NC SG
- RUNAWAYa AWIlUART*FECDWATR TO fa0 s0 g_ -
NOTE: These fransients 0-assumed multiple . operof or / equipment f ailures. See TABLE H. I -iOO - ,
.N.. .
s
.....['-.... ...
-200- N s... -
s
-300- '. ~
's
-400- , , , .
0 1000 2000 3000 4000 5000 60'00 70'00 8003 ilme (s) Fig. V11.D.3. Comparison of the energy removed through the SGs (combined) for the three runaway-feedwater transients (0 - 7200 s). 93
the rate at which energy was removed from the primary fluid depends on whether everfeed was supplied to both or to just one SG. Although the total energy removed by overfeed to one SG eventually equaled the total energy removed by ~ cverfeed to two, the rate was not as high. Thus, the decrease in the downcomer temperature was smaller for the overfeed-to-one-SG transient. Although loop flow stagnation did occur in the overfeed-to-one-SG transient, it did not occur until after the minimum downcomer liquid temperature had been reached. The initial portion of the runaway-auxiliary-feedwater corresponded to a itss-of-main-feedwater transient (by transient specifications). The rate at which energy was removed from the primary fluid during a LOFW transient l c:rresponded to the rate at which the SG liquid inventory was being boiled away. The secondary-side liquid temperature during this boiling process was constant ct the saturation temperature corresponding to the SG pressure, and the primary fluid temperature was constant at a slightly higher temperature. Hence, for this particular transient, an energy removal of 100 GW-s did not translate into o decrease in primary-fluid temperature. However, following the initiation of
; AFW, the primary-fluid temperature began to decrease slowly.
1 T t i f l 94 I
. ... . _ . ._ _ . . _ - - . - _ ~_ . _ _ _ - . _ . . _
l VIII. STEAMLINE BREAKS i -- l Transients initiated by a postulated steamline break would lead to overcooling on the primary side. The steamline breaks considered ranged from a double-ended guillotine break to a single stuck-open TBV. The general events following a steamline break were as follows. After a break or stuck-open valve , occurred in the steamline, secondary depressurization resulted. If the plant t was at FP, the reactor and turbine tripped (probably on liquid level in the SG) and the MFW flow ran back. Because the secondary liquid temperature decreased with the saturation temperature (which decreased in accordance to the depressurization history of the broken SG), the primary temperature was governed by the AT across the tubes in the SGs. The decresce in secondary pressure caused an SGIS, which initiated closure of the'MSIVs and MFIVs. If these valves operatcd correctly and isolated one SG
! from the break, asymmetric conditions were induced on the primary side. As described in the TRAC-Analysis-Methodology section (Sec. VI), this asymmetry could result in temporary flow stagnation in the " intact" loop (the loop with i the isolated SG). AFW was valved out to the " broken" SG so that it eventually boiled dry. AFW filled the intact SG and because of assumed operator inaction, I the intact SG overfilled. If neither or both SGs were isolated, symmetric conditions would exist on both the primary side and secondary side, and AFW . would be delivered to both SGs if a low liquid level in the SGs was reached.
To explore the range of stesaline break sizes and initial conditions, i seven transients were analyzed. The first set of these transients involved a 0.1-a2 (1.0-ft 2) break. Three cases for this break size were studied:
- 1. to initiate the transient while the plant was at HZP conditions;
- 2. to initiate from FP conditions; and
- 3. to initiate from HZP but to leave two diametrically opposite RCPs in
.l operation. The break location was specified such that it was non-isolable for the broken SG. This led to asymmetric SG conditions. This resulted in asymmetric primary conditions and led to flow stagnation in one loop. In the second transient, the higher decay heat created a driving force large enough to keep both loops in natural- circulation even though asymmetric SG conditions existed. The third f transient had no stagnation because an RCP in each loop was operating. J J 4 95 t
The next set of transients analyzed was the double-ended guillotine steamline breaks.- The two cases were initiated from HZP conditions. In one, it was assumed that AFW was not valved out to the broken SG. This made the broken SG a very cold heat sink and produced a corresponding low primary temperature. i i The delivery of AFW to the broken SG created asymmetric conditions producing flow stagnation in the intact loop. For the remaining double-ended break, the failure of both MSIVs to close after SGIS and also termination of AFW flow to l both SGs at 480 s (8 min) was specified. This was the one transient in which additional operator action (that is, terminating AFW) was assumed. Both SGs blev down identically resulting in strong primary natural circulation flows in both loops. Hence, loop stagnation conditions were never reached. The last set of calculations was for a small steamline break produced by a stuck-open TBV, initiated while the plant was at FP. The cooldown in the first calculation was not severe because both SGs were isolated from the break upon receipt of SGIS. For this reason, a second calculation was performed in which an additional failure of one MSIV was postulated. As in other calculations, asymmetric SG conditions existed but the decay heat immediately after FP shutdown was enough to drive natural circulation in both loops. The following sections give a more detailed analysis of each of the seven main steamline break (MSLB) calculations including a breakdown of the energy transfers that are important in understanding the resultant downconer temperature history. For completeness, the appendix contains additional plots for each transient. The uncertainties associated with these transients are estimated in Sec. X. A. 2 0 .1_-m MSL_Bs The transients described in this section were initiated by a 0.1-8 (1.0-ft2) break in the main steamline outside of containment, downstream of the flow restrictor, and upstream of the MSIVs.
- 1. From HZP_. The principal implications of HZP were:
- 1. the decay heat from the core was low (9 MW),
- 2. the stored energy in the system metal was lower than at FP,
- 3. the initial primary temperature was 552 K (534UF), which is 21 K (37.8 F) lower than the average fluid temperature at FP, and
- 4. the initial mass in the SGs was 50% greater than at FP, making its energy-removal capability 50% greater.
A sequence of events is given in Table VIII.A.1. 96
TABLE VIII.A.1 0.1-a2 MSLB FROM HZP SEQUENCE OF EVENTSa Time (s) Event 0 0.1-m2 (1.0-ft )2MSLB in Loop A; TBVs closed 9- All pressurizer heaters tripped following low pressurizer level 18 SGIS on low secondary pressure (MSIVs and MFIVs closed) 28 Asymmetric-SG pressure signal; AFW to SG A prevented 54 SIAS on low primary pressure 70' HPI flow began 75 Pressurizer emptied 84 RUPs tripped; natural circulation in Loop A began 90 AFW to SG B began (based on Ap level-measurement) 200 Flow stagnated in Loop B 300 Pressurizer began to refill 630 Pressurizer proportional-heaters automatically turned on because of level recovery 1000 HPI flow ended; charging flow continued 1230 High-level signal in SG B 1325 SG A dried out (natural circulation in Loop A decreased) 2800 Pressurizer heaters tripped because of high system pressure 3120 System repressurized to the PORV setpoint
- 3500 SG-B-secondary liquid level rose above the moisture-separator deck and natural circulation on this secondary side began i 4200 Natural circulation on Loop-B primary began (probably l should have begun at 3500 s but because of input errors was ef fectively delayed by ~600 s.)
i 7200 Calculation terminated aThese transients assumed multiple operator / equipment failures. 97 v --- m ~~-'--~ew-~ -w- ~-~- ---~- w---- '- -~ ~> --
Figures VIII.A.1 through VIII.A.3 show the downconer liquid temperature at the core midplane (axial level 5) and the non-convective energy transfers on two time scales. These non-convective energy transfers were between the primary and the SGs, the heat slabs, the core, and the RCPs. The convective energy-transfer mechanisms in this transient were inflow from the HPI and charging system and outflow through the PORVs. Figure VIII.A.1 indicates that the events of this transient may be divided into three phases. 4 Phase 1 (0-1300 s) was dominated by the blowdown of SG A. The blowdown was limited by choked flow through the 0.1-a2 (1.0-ft 2) break. The fluid [ exiting the break was 100% steam. As SG A depressurized, the saturated liquid l l flashed and the secondary temperature decreased according to the saturation curve. Power extraction slowed as the liquid inventory depleted because the rate of decrease in the secondary-side-heat-sink temperature slowed. Because f - AFW was valved out to SG A as a result of the asymmetric-SG pressure signal, its secondary eventually voided completely. This event marked the end of Phase 1 , (at 1300 s). Blowdown in SG B ended at 18 s when closure of the MSIVs isolated it from the break. After isolation, SG A continued to cool the primary, bringing its , temperature below that of SG B. Reverse heat transfer began and the energy removed by SG B during its short blowdown was restored to the primary system by 150 s. Even though AFW injection began at 90 s, it did not affect the temperature in the riser region during Phase 1. AFW merely collected in the SG downconer in Phase 1. The net contribution of SG B to the minimum downconer l liquid temperature was close to zero.
- . As described in the TRAC-Analyses-Methodology section (Sec. VI) the energy transfers into and out of the primary-system have been broken down into non-convective and convective terms. Table VIII.A.2 gives the total energy l transferred and the sources and sinks at 1300 s (time of SG-A dryout and the minimum downconer liquid temperature). As seen from the table, the contribution from the primary-side heat slabs was relatively small, only 20% of the total l energy removed by SG A. Input from the RCPs was minimal and because the plant J was at HZP, the energy input from the core was small. Table VIII.A.3 gives the i i
energy contribution of the HPI and charging flow at 1300 s. Approximately 47228 kg (104091 lb) of HPI and charging flow were injected by the time of the I minimum downconer liquid temperature. Although this injection increased the overall energy of the system, the specific internal energy decreased. Hence, the primary temperature decreased. 98
590 -
, , , , , 600
"~ ~
NOTE: These transients 330 4 assumed multiple -300 l\ operator / equipment 300 1 f ailures. See TABLE H. E F e 470 -
-400 Yy
}
$ I'e
- o. 440
-300
- 3 _
410 - 1 2 . g 380- G A dryout
- -200 350- 1. blowdown of SG A
- 2. heating from core and slabs; no energy removal from SG B (or A) _
_ 3. natural circulation and energy - iOO removal in Loop B 290- , , , . . . . 0 1000 2000 3000 4000 5000 6000 7000 8000 Tirne (s) Fig. VIII.A.1. Downcomer liquid temperature during 0.1-m2 MSLB from HZP. 50 , , , , , SLAB 25~ CORE 5 + o
. t t SG B PUMP NOTE: These transients ossumed multiple 9
~50~
I. i d , lures ce A NET - 10 0 _
~
-125 ~ SG A 33 o .
-95 . . . , , . ,
0 200 400 600 800 1000 1200 1400 1600 Time (s) l Fig. VIII.A.2. l Summary of non-convective energy transfers that occurred during the 0.1-m 2 MSLB from HZP (0 to 1600 s). 99
TABLE VIII.A.2-ENERGY BALANCE FOR NON-CONVECTIVE TERMS AT THE TIME (1300 s) 0F MINIMUM DOWNCOMER TEMPERATURE FOR THE 0.1-m2 MSLB FROM HZP Source Energy (GW-s)a Decay heat + 12.3 Primary-side heat slabs + 29.4
-RCPs + 1.5 SG A -153.5 SG B -
0.7 Net -111.0 l-a+ = energy addition to primary fluid.
- = energy removal f rom primary fluid.
TABLE VIII.A.3 ENERGY BALANCE FOR CONVECTIVE TERMS AT THE TIME (1300 s) 0F MINIMUM DOWNCOMER TEMPERATURE FOR 0.1-m2 MSLB FROM HZP Source M (kg) h _(kJ/kg) Energy (GW-s)s Charging flow + 10790 111.9 + 1.207 9 299.8 k (800F) HPIs 9 322 K (1200F) + 564 204.5 + 0.115 9 302.6 K (850F) + 4270 123.5 + 0.527 9 285.9 K (550F) + 31604 53.5 + 1.691 PORVs - - Net + 47228 + 3.540 , a+ = energy addition to primary fluid.
- = energy removal from primary fluid.
i ( r 100 i
The specific internal energy at the time of minimum downcomer temperature casociated with these energy transfers is given by Eq. (VI-5): 3 3 u = (224 x 10 ) (1227.4 x 10 ) + 3.540 x 109 - 111.0 x 109 (224 x 103 + 47228)
= 617.5 S .
kg The bulk fluid temperature corresponding to this specific internal energy (using tables in Ref. 3) is 420 K (297 F) 0 or a bulk temperature drop of 132 K (2370 F). However, the minimum downcomer liquid temperature predicted by TRAC was 400 K. The difference between the bulk temperature and the downcomer temperature was a result of three effects:
- 1. Loop-B flow stagnation during ~300-4200 s;
- 2. spatial variations throughout the vessel and loops; and
- 3. decreased upper-head circulation.
These phenomena were described in more detail in Sec. VII. After the RCPs tripped at 84 s, the flow in Loop B stagnated. This removed one-third of the primary system heat capacity, and resulted in a larger temperature decrease per unit of energy removed via EG A. Another " effective" mass loss resulted from the decreased circulation of the fluid in the vessel above the hot legs. This mass represents ~20% of the vessel mass and was ~60 K hotter than the average loop temperature at 1300 s. Of the 122-K bulk temperature drop, the convective transfers (HPI and charging flow) accounted for 31% and the non-convective
- transfers for 69%.
Phase 2 (1300-4200 s) of the downcomer liquid temperature history shown in Fig. VIII.A.1 was the period after SG-A dryout and before natural circulation I tas established in Loop B. The downcomer temperature went through a maximum of 435 K (3240F) at 4100 s. The non-convective energy transfers up to 4200 s are l cummarized in Table VIII.A.4. Energy was added via the core and the heat slabs. The RCPs and SG A had no energy input during this time period. From 1300 to 4200 s, a net energy gain of 30.9 GW-s to the primary system occurred. The convective energy transfers for Phase 2 are summarized in l Table VIII.A.5. The deadhead of the HPI pumps was reached at 1000 s, so HP1 flow was zero during Phase 2. Charging flow continued throughout the transient. I I 101
TABLE VIII.A.4 ENERGY BALANCE FOR NON-CONVECTIVE TERMS AT 4200 s FOR THE DOWNCOMER TEMPERATURE FOR THE 0.1-m2 MSLB FROM HZP Source Energy (GW-s)a Decay heat + 39.6 Primary-side heat slabs + 42.5 RCPs + 1.5 SG A -153.6 SG B - 10.1 Net - 80.1
"+ = energy addition to primary fluid.
- = energy removal from primary fluid.
TABLE VIII.A.5 ENERGY BALANCE FOR THE CONVECTIVE TERMS AT 4200 s FOR THE DOWNCOMER TEMPERATURE FOR THE 0.1-m2 MSLB FROM HZP Source M (kg)_ h(kJ/kg)_ Energy (GW-s)a , Charging flow at +34860 111.9 + 3.901 0 299.8 K (800F) HPIt 0 322.0 K (120*F) + 564 204.5 + 0.115 0 302.6 K (85 F) + 4270 123.5 + 0.527 0 285.9 K (55'F) +31604 53.5 + 1.691 1268.6 b 0 PORVs at ~561 K (550 F) - 89,64 -11.,372 62334 - 5.138 a+ = energy addition to primary fluid.
- = energy removal from primary fluid.
b estimate value. 102
The PORVs opened at 3120 s, relieving the fluid injected by the charging system but at a much higher temperature. The final specific internal energy at the end of Phase 2 (1300-4200 s) is dstermined from Eq. (VI-5): i
, , (224 x 10 3) (1227.4 x_ 103 ) - 5.138 x 109 - 80.1 x 109 (224 x 103 + 62334) i
= 662.5 S .
kg Using~ the thermodynamic tables in Ref. 3, the corresponding bulk temperature is 431 K (3160F). The convective transfers led to a temperature decrease of 14.5 K (26.10F), and the non-convective transfers led to a temperature increase of f 25.5 K (45.90F), with a net increase in the bulk temperature of 11 K (19.80F) during' the interim from 1300 to 4200 s. Again the bulk temperature varied from 1 the TRAC-calculated downcomer liquid temperat tre primarily because the flow was etagnant in Loop B. Energy transfer did not take place with the primary system j co a whole, as approximated in this analysis. Phase 3 (4100-7200 s) began with the onset of natural circulation in L op B. Because the operator was assumed to fail to throttle AFW, the liquid icvel in SG B rose above the moisture-separator deck and natural circulation was octablished within the SG (riser to SG downcomer). AFW mixed with the warmer f liquid in the riser, lowering the effective heat-sink temperature. Energy removal by SG B induced natural circulation on the primary side. As the primary fluid mixed in the system, the downcomer temperature approached the bulk temperature calculated at 4100 s. The calculation was ended at 7200 s with the primary temperature d: creasing slightly. SG B was slowly becoming a colder heat sink with the-centinued injection of AFW. The primary temperature was also decreasing as charging flow replaced the hotter fluid leaving through the PORVs. Figure VIII.A.4 gives the primary pressure history and Fig. VIII.A.5 gives the HPI and charging flow for the transient. The energy transfers described previously resulted in a rapid decrease in pressure to a minimum of 4.7 MPa by 150 s. Little voiding (maximum void fraction of 0.002) occurred in the upper h:ad. As the pressure decreased, HPI flow increased and began to make up for 103
20 . . . . . . CORE l 75- - 50- I - 1 ~ -- p - PUMP 25-
)
i
; ; I OH g 9 g -
NET
"" ~
- 10 0 - ',"3'....**..,.~~'y..........ossumed
' NOTE: These fronalents multipie operator equipmen t
- 12 5 -
f ailures. ee TABLE E. sc A
- 15 0 - = =
l -
-175 . , . . . . .
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.A.3. Suanary of non convective energy transfers that occurred during the 0.1-m2 MSLB from HZP (0 to 7200 s). ts , , . 2450 16 --
- 2170 u- ,
i 12 - ! -1750 9 a v 1D -
.. i40n O
- l. .. . 5
'- # ~ ' 5* NOTE: These transients E s- assumed multiple operator equ ipmen t -700 4 .- f ailures. ee TABLE H.
-350 2-0- 0 0 10 0' 0 2dOO 3dOO 4000 5do'n 6600 7000 8000 Tirne (s)
Fig. VIII.A.4. Priinary system pressure during 0.1-m2 MSLB frou HZP. 104 L
4 the primary liquid contraction that resulted from the overcooling by SG A. When the deadhead, 8.8 MPa (1270 psig), of the HPI pumps was reached at 1000 s, the HPI flow ceased. However, in accordance to transient specifications, the charging system was left on, causing continued repressurization. The PORVs cpened at 3120 s, and by 3250 s, the pressurizer was completely filled with
' liquid.
Figure VIII.A.6 shows the flow in each hot leg. After the RCPs were tripped and coasted down, Loop A went into natural circulation driven by heat in from the core and by the high energy-removal rates caused by the steamline break. In Loop B, however, energy was added to the primary fluid as the primary became cooler than the Loop-B secondary liquid. A gravity head that opposed the driving gravity head in the vessel was established in the tubes and thus caused the flow in the Loop-B hot legs to stagnate. This phenomenon was described in more detail in Sec. VI. When natural circulation began on the secondary side of SG B (4200 s), a newly created gravity head that formed in the the SG-B U-tubes ccted in unison with the gravity head in the vessel, and natural circulation was ts-established in Loop B. With no energy removal from SG A, the fluid flowed preferentially to Loop B after 4100 s. Errors in the gravity terms and in the friction terms in the input deck exaggerated the difference in the loop flows i during 0-7200 s. The imprecise gravity terms were equivalent to 0.025 m (1 in.) of water, and zero friction was calculated for the pump-suction legs and the SG tube region. The net effect was a loop flow of ~16 kg/s in the reverse direction (cold leg-to-hot leg) under isothermal loop conditions with no decay
- h
- st.
After the RCPs were tripped, a mass flow of ~50 kg/s was circulating i cround the cold legs in Loop B (see Fig. VIII.A.7) despite the stagnant hot-Icg B. These flows are not believed to be real and were aggravated by the strors in the gravity and friction terms. Physically, it is expected that the l flows in the Loop-B cold legs should have been steady at the HPI/ charging flow. [ Figure VIII.A.8 shows the liquid temperature upstream of the HPI/ charging parts. The temperature changes in Loop A were smooth because this loop was I always circulating. A wide swing in the temperature in Loop B occurred at 700 s when the previously-described circular flow began in this near-stagnant loop. Cold HPI and charging flow filled the cold legs in Loop B, while the fluid in the loop seals remained warm. Thus, wide variations in the liquid temperature f oxisted in Loop B. When the small circular flow began, HPI and charging liquid l \ 105 1
- . - - . _ . . - _ _ _ y .. ., ._._,....,....,._._,mr,-,-.___._,,_.__,,_--,,,...,,r
,-,_,r ,-.3._,,,,,-,_-.m.rm_,,y-,,
20 i , , 18 -- - 40 Mai rLow Ptn coto uG 16 -- ------ --- cMA#GmG FLQw PER (CLD gG (LQQP5 A2 Ah3 gQ .g 14 -- -
-30 R, 32 - -
3
-25 -
T '~ NOTE: These transients T e assumed multiple e n - 20 j *~ operator / equ.i pmen t a f ailures. See TABLE II.
~
- 15 6- -
~
- ' * " ' " ' " " " " " * * " " - ' - " - " ' - " " " - ~ ~ " - * "
4- - 2- -5 0- , . . . . 0 0 1000 2000 3000 4000 500G 6000 7000 5000 rma (s) Fig. VIII.A.S. HPI/ charging flow during 0. -r22 MSLB from HZP. 800 -; i , , , -1730 7"~ -: is00 { LocP A (cows 20 600- ! - ---~~ toor e (cowP t0 -
-; -1260
**~
i NOTE: These transients
! assumed multiple ~'
i operator / e ui ment O 400- i f ailures. See A LE H. -
! -/50 300- !
- II' Y
[ l - n _i / _300 e
- E 2C0 - !
wn/\).t/ i. f g_ -j l _-250 0 . !. [ .. ... . . . .~...
. . . . g ....y/..-l '-c%y
.0
- 10 0 -_ . 250
~200 , , , . . . .
0 1000 2000 J000 4000 0000 tiOOO 7000 6000 rme (s) Fig. VIII.A.6. Hot-leg flows during 0.1-m 2 MSLB from HZP. 106 J g - - - -
,e ~m ergg, a,,--ce----..,r. . , , , - , , - , -,,-,w,.-,,. m y-e n, , - r- --m-. w,r, .- - --n -- - - -
400 - i , i . 350--
~~ - LOOP A.(00w8 25) -~M
-- - L30P aJ (Coup 45) 300- LOOP b1(Coup n) -
- Lo0e e2 f.ccup ss)
^
23o_ NOTE: These transients _ ossumed multiple r zw--
) $P$iures a se AL - '*5*
o j . . ~ - 3 J $[~ f
.E , ,'
M d/ ' >
~"
- 15 0 E
50- '- -
.^ ~~ f
( ,r ,,.. j(\ %%% 0-- h -_o f, <v, N',)e N f' s
.50
-20 4 , , , , ,
0 1000 2000 3000 4000 5000 6000' 7300 8000 Time (s) Fig. VIII.A.7. Cold-leg flows during 0.1 m2 MSLB from HZP. 590 -
, i , . , , -600 560-530- I toop
-- 500 A. . . . . . . . . . wo, 3, ,2 4 *- _ . . .
. toop gg 500-
- ' LOOP B2
- g. <
( C
~4"
[ co. I i 440- ,,,... _ s a '~ %.S. -300 e 410 - ..:; p "- f 1 -
't ' 6 380
{J
-200 350- NOTE: These transients -
assumed multiple 32"- operator / equioment '
- f ailures. See TABLE H. 30 o 2$0 , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.A.8.
. Liquid temperature upstream of HPI/ charging flow ports during 0.1-m2 MSLB from HZP.
107
-_m.. , _ . . --
- _ . , . _ . , . _ _ . _ , _ _ ,-__..m- __w
was swept into areas occupied by the warmer liquid, and thus a large temperature cwing was indicated. Because we do not believe the circular flows to be physical, we also do not believe the dramatic temperature change at 700 s. The liquid upstream of the HPI/ charging ports would probably- have remained relatively warm, having received little backflow from the HPI and charging injection.
- 2. From FP_. A 0.1-m2 break was postulated again to occur in the main i
steamline, but this time the plant. was operating at FP conditions. Table VIII.A.6 lists the sequence of events for this transient.
, It was anticipated that this transient would be less severe than a comparable break froo HZP for two reas?ns:
- 1. decay hertt and stored energy in the system metal were substantially higher; and
- 2. the. initial mass in the SGs was two-thirds that at HZP.
The .downcomer liquid temperature at the core midplane and the non-convective energy transfers are presented in Figs. VIII.A.9 and VIII.A.10. Phase 1 (0-300 s), labeled on.the plot for the downcomer liquid temperature, was the period during SG-A blowdown. Because the system energy was higher and the SG mass lower, dryout occurred much earlier (at 300 s) than for the same transient from HZP (at 1300 s). Main feedwater was added to each SG for 15 s efter .the reactor / turbine trip (~5000 kg (11000 lb)), but this was balanced by steaa-flow'through the TBVs. Because loop flows were very low in Loop B from l
~250-750 s, the downcomer liquid temperature varied as much as 30 K (54 F) in the azimuthal direction. The non-convective energy transfers that led to a Einimum downcomer temperature of 468 K (383 F) are given in Table VIII.A.7. The total energy-removal capability of SG A was 98.1 GW-s. Blowdown of SG B removed 30.9 GW-s before SGIS at 44 s. After this, SG A cooled the primary below SG B, '
\
and the resulting reverse heat transfer, though small, severely slowed the flow in Loop B. l Table VIII.A.8 lists the convective terms at the time.of minimum downcomer temperature (300 s). Because the minimum pressure was not as low as in the break- from HZP, the HPI injection was not as high. Thus, HPI flow did not contribute as much to the cooldown. 108
l l TABLE VIII.A.6 l 0.1-m2 MSLB FROM FP a SEQUENCE OF EVENTS Time (s)_ Event 0 0.1-m2 (1.0-ft )2 steamline break on Loop A a) Turbine / reactor assumed to be tripped b) ADVs and TBVs opened on " quick-open" logic 20 Pressurizer heaters tripped following low pressurizer level 29 TBVs and ADVs closed on low primary temperature 32 SIAS on low primary pressure 44 SGIS: MFW pumps tripped; MSIVs/MFIVs began to close 62 RCPs tripped by the operators 52 Asymmetric-SG pressure signal; AFW to SG A prevented 58 AFAS because of low SG-A level; AFW initiated to SG B 70 HPI flow began 300 SG A dried out 495 HPI flow ended; charging flow continued l 595 Pressurizer proportional heaters turned on because pressurizer level recovered 1975 PORVs opened because of high primary pressure 2500 Calculation terminated - extrapolation based on previous calculation of this transient "These transients assumed multiple operator / equipment failures. i
- - 109 l
l l l
500 -
. . . . i i -600 560- --
L A A A A A 530- p e a s= m * --500 V 4 500- 3 - g k -400 I
,4 SC A dryout 440- -
E "5 1. SG-A blowdown E
- 2. Low flow in Loop B - 00 y 410 -
f 3. Natural circulation in loop B
~
]e l3 4. Extrapolation (based on previous
- g_ calculation) _
]
- 200 35 -
NOTE: These transients _ assumed multiple operator / equipment 320-f ailures. See TABLE H. ~ _ ,oo 200 . . . . . . . 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.A.9. Downcomer liquid temperature during 0.1-m2 MSLB from FP. 200 . , . . , . . . . NOTE: These transients g_ ossumed multiple _ operator / equipment f ailures. See TABLE I. CORE m_ \ - 4 7 50- SLAB 4 \ 0, < i _ i I I 1 5 I - I b / - PUMP _,,,m ,,.... ,.. t
\ ',,,....e-*- NET
\
,., ',,,... - /
SC B
-10 0 - = -
= = ..
/
SC A
-15 0 . , . . . . . . . .
0 250 500 750 1000 1250 1500- 1750 2000 2250 2500 2750 Time (s) Fig. VIII.A.10. Summary of non-convective energy transfers that occurred during the 0.1-m 2 MSLB from FP for 0 to 2500 s. 110 (. .__ _ _
TABLE VIII.A.7 ENERGY BALANCE FOR NON-CONVECTIVE TERMS AT THE TIME OF MINIMUM 2 DOWNCOMER TEMPERATURE (300 s) FOR THE 0.1-m MSLB FROM FP So_urce Energy (GW-s)a l l Decay heat +30.8 l Primary-side heat slabs + 9.8 RCPs + 1.1 SG A -98.1 SG B -30.9 Net -87.3 a + = energy addition to primary fluid.
- = energy removal from primary fluid.
TAELE VIII.A.8 ENERGY BALANCE FOR CONVECTIVE TERMS AT THE TIME OF MINIMUM 2 DOWNCOMER TEMPERATURE (300 s) FOR 0.1-m MSLB FROM FP \ Source M (kg)_ h (kJ/kg) Energy (GW-s)a l Charging flow + 2490 111.9 +0. 27 9
@ 299.8 K (800F)
HPI
- @ 322.0 K (1200F) + 564 204.5 +0.115
@ 302.6 K (850F) + 4270 123.5 +0.527
- @ 285.9 K (550F) + 4402 53.5 +0.236 PORVs
! Net +11726 +1.157 i a + = energy addition to primary fluid.
- = energy removal from primary fluid.
111 ( i
The resulting specific internal energy of the system was:
, , (219 x 10 3) (1332 x 10 3) + 1.157 x 109 - 87.3 x 109 (219 x 108 + 11726) l
= 891 kJ .
kg 4 l Using the thermodynamic tables in Ref. 3, the corresponding bulk temperature at the time of minimum downcomer temperature was 482 K (408 F). Hence, the bulk temperature decrease was 90 K. 0 The 14-K (25 F) difference between Tbulk and the minimum downcomer temperature as predicted by TRAC was caused by the same phenomena as in the break from HZP:
- 1. very low flow in Loop B between ~250-750 s;
- 2. spatial variations; and
- 3. decreased upper-head circulation.
Non-convective energy transfers accounted for 78% of the maximum cooldown, whereas convective, energy transfers contributed 22%. Phase 2 (~300-800 s) was a period of relatively rapid heating following SG-A dryout. Because Loop B was close to stagnation, less primary fluid was available to receive the energy deposition from the core, and so the specific snergy of the flowing fluid increased rapidly. As the primary temperature i increased, SG B became an effective heat sink. In Phase 3 (800-2500 s), the
- average core power was ~46 MW. SG B removed ~24 MW and the PORVs removed some energy af ter they opened at 1975 s, but the primary fluid continued to heat.
Phase 4 (2500-7200 s) was extrapolated from a previous calculation of the
.came transient. As the core power decreased, a balance was achieved with the anergy removal by SG B and flow through the PORVs. A quasi-equilibrium state sxisted in Phase 4 with the downcomer temperature at 530 K (495 F). The primary temperature would decrease slightly with time because t
the decay heat was decreasing; l. l 2. SG B was becoming slightly colder with continued AFW; and
- 3. charging flow at 300 K (80 F) was replacing hotter fluid that left through the PORVs.
112
-- .._ -- _.- ~ - - _ . - . - _ _ -
Figure VIII.A.11 shows the system pressure during the transient. The fluid in the upper head saturated, slowing the depressurization and keeping the minimum pressure higher than in the break from HZP. Voiding in the upper head I' was as high as 40% void fraction between 70 and 460 s (see Appendix E for ! additional plots). Figure VIII.A.12 gives the HPI and charging flow. HPI flow was possible only when the system pressure was below the shutoff head of the HPI pumps (8.8 MPa (1270 psig)). Because the minimum pressure was higher than during the break from HZP, the HPI fluid injected in this case was 40% of that injected during the break from HZP. The primary pressure was extrapolated to 7200 s, based on the system pressure remaining at the PORV setpoint. Figures VIII.A.13 and VIII.A.14 depict the flows in the hot and cold legs. Coastdown of the RCPs took ~200 s af ter they were tripped. At ~ 500 s , natural circulation increased in Loop A as it decreased in Loop B. This corresponds to the "one SG in reverse, one SG in zero" heat-transfer mode described in Sec. VI. As the reverse heat transfer stagnated Loop B, the At across the core increased because of the reduced flow. This increased the gravity head in the vessel, which in turn increased the flows in Loop A. After this brief period of very f- low flow in Loop B, the flow resumed because of the increased gravity head and was sustained by AFW injection. The flow was higher in Loop B because of the additional driving force produced in SG B from the AFW compared to a zero i driving force produced in SG A (because dryout had occurred at 300 s). Figure VIII.A.15 shows the liquid temperature upstream of the HPI and charging ports. Because the flow was always circulating in Loop A, the upstream temperature in this loop did not vary much from the downcomer liquid
- temperature. During low flow in Loop B, the upstream temperature was j considerably warmer than that in the downcomer because this fluid was not mixed with the HPI and charging flow and was not affected by the blowdown of SG A.
- 3. With Two Operating RCPs from HZP. This transient was identical to the one described in Sec. VIII.A.1 except that two diametrically-opposite RCPs j remained in operation throughout the transient. This is a pump-trip philosophy
, currently proposed by C-E. A sequence of events is presented in Table VIII.A.9. l The principal effect of leaving two RCPs in operation was that loop-flow stagnation did not occur in Loop B and SG B became a considerable heat source r during the initial part of the transient (0-500 s). j 113 s
18 , , , , , , ,
~
a A A 16 - an = = -
- 210 0 i j
NOTE: These transients ~* assumed multiple 7 operator / equipment ?a l A 10 - f ailures. See TABLE II. 0 e _-1400 e s- - g j extrapolated -1050 [ 6- -
- 700 4- -
2- -- 350 0 , , , , , , . 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.A.11. Primary-system pressure during 0.1-m2 MSLB from FP. 20 , , , , , , , 1s -- -40 HPi rLow pga coto Ltc 16-- ---------- cwancWG row PLa COLD LIG (LCOPS A2 Ah3 31) -- 3, l
~
14 -_ _3o n-l $~ 6 -
-25 3 o
j io_ q NOTE: These tr ansi en t s . a s-assumed multiple OPerotor / equipment
-20 (
j f ailures. See TABLE H. _ ,3 s- -
-10 2- --S 0 , , , , , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s)
Fig. VIII.A.12. HPI/ charging flow during 0.1-m2 MSLB from FP. 114
200 , , , , , , i goo. -. --2000 Loop A (coup 20
- ~
800-_i ---------- Loop a (coup 10 ~-1750 i
#~ ~-1500 g ,o _
j; ' NOTE: These transients _
, g 6 -i assumed multiple -1250 o j 3co_ j operator / equipment _
c g
- _i f ailures. See TABLE H. -i000 E
- o 5
40- 2
~ . .(1 - ,so
_l ' 300-ii :
-I j -
-500 200- ;
- i. p 250 10 0 -~ !!
k,! ! 0 , , , , , , , 0 0, 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.A.13. Hot-leg flows during 0.1-m 2 MSLB from FP. SCO , , , ,
- -1050 450-Loor At(coup 25)
---------- Loop A2 (Coup 45)
Loor a1(coup is) _-900 400-- - Loor s2 (coup ss) I - _7g 350-_ { 3co_ NOTE: These transients _ g 6 _; assumed multiple _,oo o j 250- OPerotor / equi ment . j
, f oilures. See TA LE H. .
-- 450 E 200-- *
!%.m,..-
~
DO- d _3co 100-l
- 15 0 50-t
, 0 0 , , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.A.14. Cold-leg flows during 0.1-m 2 MSL3 from FP.
> 115
__ _ - -_ m . _ _
' TABLE VIII.A.9 0.1 m2 MSLB WITH TWO OPERATING RCPs FROM HZP a
SEQUENCE OF EVENTS Time (s) Event 0 0.1-m2 (1.0-ft )2MSLB in steamline A; TBVs close j 20 All pressurizer heaters tripped on low pressurizer level 30 SGIS on low secondary pressure 45 Asymmetric-SG pressure signal; potential AFW to SG A prevented 80 SIAS was received because of low primary pressure i 110 Two of the fou'r RCPs were tripped by the operator 115 HPI flow began
-125 Pressurizer emptied 220 AFAS based on low liquid inventory in SG A; AFW flow to SG B began
. 350 Pressurizer began to' refill
-500 SG A dried out and reverse heat transfer in Loop B ended 580 Pressurizer proportional heaters turned on because of level recovery 930
'PPI flow ended because shutoff head was reached 1700 High liquid level in SG B 2350 Pressurizer proportional heaters tripped on high pressure 2400 PORV setpoint was reached because charging flow was j
unthrottled 5300 Calculation terminated; conditions stable "These transients assumed multiple operator / equipment failures. 116 l
i Figures VIII.A.16 through VIII.A.18 show the downcomer liquid temperature at the core midplane and the non-convective energy transfers on two time scales. Again the downcomer temperature curve is logically divided into four phases. Phase 1 (0-500 s) corresponded to SG-A blowdown and ended at the time of minimum j downconer temperature. Because two RCPs were still operating, energy-transfer l rates were much higher than when all four RCPs were tripped. SG A dried out at ! 500 s. The forced circulation allowed SG B to deposit considerable energy into the primary while it was being cooled by SG A. The minimum downcomer temperature was 440 K (333 F), which was 40 K (72 F) higher than with all four RCPs tripped. Table VIII.A.10 indicates the non-convective energy transfers at the time of minimum downcomer temperature (500 s). Comparing Table VIII.A.10 and Table VIII.A.2, the decay-heat and primary-side-heat-slab addition were less for this transient at the time of Einimum downcomer temperature because the time until SG dryout was shorter. Heat input from the RCPs was greater because the two operating RCPs deposit 8.7 MW into the fluid throughout the transient. The energy removal by SG A was approximately the same (small differences exist because of slightly different initial masses). A large difference in energy transfer existed for SG B because of the forced-convection flow. Table VIII.A.11 gives the energy contribution of the HPI and charging flow in Phase 1 (0-500 s). Because the time of minimum downcomer temperature was shorter than in the previous transient, the amount of HPI and charging flow was considerably less: 26495 kg compared to 47228 kg. This was a significant contribution to the 40 K (72 F)0 temperature difference in minimum downcomer liquid temperature between the two calculations. The change in the specific internal energy associated with these energy transfers was: l 3 3 u = (224 x 10 ) (1227 x 10 ) + 2.043 x 109 - 100.7 x 109 I (224 x 103 + 26495) 1 i l = 698 kJ-- . kg I l Using thermodynamic tables from Ref. 3, the corresponding bulk temperature at the time of minimum downcomer temperature was 440 K (333 F). Hence, the bulk 117 l l v - ,,, - -- ~ , - . - , , . - - - - - - - . - - , - , , - , , , - , , - , .,-----..n-,-.,c,. - , . . , . . - . . - . , ,---,,-n--- ,-
500 -
, , , , , , . -600 560-530- . Loop At -@0 f
.........- LOOP AZ l I
, .. . . . . Loop gg 500- - toor e2 -- I g c 1
-400 470 - -
l E. 440- -- o, E
-300 y b,
410 - - E NOTE: These transients 3 30- assumed multiple - 5 operator / equipment -200 350- f ailures. See TABLE 11. - 320- -
- 10 0 290 , , , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.A.15. Liquid temperature upstream of HPI/ charging flow ports during 0.1-m 2-MSLB from FP. 590 -
, , . , , i -600 5s0- NOTE: These transients -
assumed multiple 333 operator equipment _-500 i f allures. ee TABLE H. 500- - y . G A dryout -400 p M- , E 1 3 d40~ 4 ~ E i g
- 1. Blowdown of SG A; SG B was a heat source -300 y 410 - 2. Heating from core and RCPs as SG B became a heat sink *y
. 3. Quasi-equilbrium between decay heat and 5 380- 4 SC B/PORVs Extrapolated -
[
-200 350- -
320 - -
.300 290 , , , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.A.16. Downcomer liquid temperature during 0.1-m2 MSLB with two ggg operating RCPs from HZP. l v ,,r-- -- g -- ew n- ~,g
50 . , . . . . . . . 25-
# ~
\
\, SG B I
PUMP
\ CORE
\
NOTE: These tronslents Y assumed multiple operator equipmen t g v f ailures, se TABLE I. f '.,
. NW -
a ............. ...- - -
-10 0 - . . . . . ' ' " * ----e~~~~"~~
-G5- SC A
.g$o _ _
-175 . . . . . . . , i 0 10 0 200 300 400 500 600 700 800 900 1000 Time (s)
Fig. VIII.A.17. Summary of non-convective energy transfers that occurred during the 0.1-m2 MSLB with two operating RCPs from HZP 0-1000 s. 75 . . . . . . . . . . Co tE 50- - 25-PUMP OH SG B NOTE: These transients - , ^II assumed multiple l operator e pment b f ailures. ee BLE I. M 50
,,,,,,.....e--~'""*"'
l g _73_ ,
'., , . . . ~ , , . .
-10 0 - **" -
- 12 5 - SG A -
.m_
-f75 . . . . . . . . . .
0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000 5500 Time (s) Fig. VIII.A.18. Summary of non-convective energy transfers that occurred I19 during the 0.1-m 2 MSLB with two operating RCPs frora HZP for 0-5300 s.
TABLE VIII.A.10 ENERGY BALANCE FOR NON-CONVECTIVE TERMS AT TIME OF MINIMUM TEMPERATURE (500 s) FOR 0.1-m2 MSLB WITH TWO OPERATING RCPs FROM HZP Source Energy (GW-s)a Decay heat + 4.7 Primary-side heat slabs + 18.0 RCPs + 5.3 SG A - 146.1 SG B + 17.4 Net - 100.7
* + = energy addition to primary fluid.
- = energy removal from primary fluid.
TABLE VIII.A.11 ENERGY BALANCE FOR CONVECTIVE TERMS AT TIME OF MINIMUM TEMPERATURE 2 (500 s) FOR 0.1-m MSLB WITH TWO OPERATING RCPs FROM HZP Source M (kg)_ h (KJ/kg) Energy (GW-s)a Charging flow + 4150 111.9 +0.464 6 299.8 K (80 F) HPI: 6 322.0 K (120 F) + 564 204.5 +0.115 i ' @ 302.6 K (85 F) + 4270 123.5 +0.527
@ 285.9 K (55 F) +17511 53.5 +0.937 PORVs - - -
Net +26495 +2.043
"+ = energy addition to primary fluid.
- = energy removal from primary fluid.
l l i 120 I
I j t(aperature decrease was 112 K (2020F). This corresponds almost identically to the minimum downconer temperature predicted by TRAC. The two operating RCPs ' kspt the primary fluid well mixed and so the bulk temperature would be expected ) to be close to the downcomer temperature. Thus, convective energy-transfer mechanisms accounted for 21% of the maximum cooldown and non-convective mechanisms accounted for 79% of the cooldown. Phase 2 (500-1900 s) was a period of primary-fluid heating (from the core, I the two operating RCPs, and the primary-side heat slabs) before SG B became a j cignificant heat sink. Not much cooling was provided by HPI and charging flow. Phase 3 (1900-5300 s) began with a significant increase in the heat-transfer i rate across the tubes in SG B. This abrupt increase was a result of an 1 inadequacy in the TRAC code
- but perhaps was physical to some extent. As the escondary side of SG B filled with AFW above the moisture-separator deck, the i liquid began to spill over into the steam space in the SG downcomer above the feedwater ring. The spillover allowed hot liquid in the riser region to mix with cold liquid residing in the downcomer of the SG, in addition to forcing
[ cold liquid in the downcomer to flow into the riser region. As a result, the i 4 cacondary-side-heat-sink temperature dropped rapidly, and the energy-removal
! rcte increased. The primary-side fluid temperature followed the decrease of the j cacondary-side-heat-sink temperature.
After 2500 s, a quasi-equilibrium state was reached. The PORV opened, rsmoving approximately 5.5 MW. SG B removed the remainder of the energy input from the core, the heat slabs, and the RCPs, which amounted to ~15 MW.- The . { esiculation was terminated at 5300 s with the system in this quasi-equilibrium state. Phase 4 represents the extrapolation to 7200 s. The system was cooling clightly with time because SG B was becoming a colder heat sink with continued ) AFW, and charging flow was replacing the hotter primary fluid leaving the PORVs. i
- Figure VIII.A.19 shows the pressure history for the transient and Fig. VIII.A.20 indicates the HPI and charging flow, which was very similar to the previous transient. (Because two RCPs were operational during this transient, the pressurizer spray would be available above 15.48 MPa (2275 psia).
However, this was not modeled during this transient. Presumably the rcpressurization rate would not be significantly altered with the sprays operating as compared to the case of no sprays.) However, because SG A dried out I ,
*This error has been corrected in later versions of TRAC-PE1.
1 121 1
,v-v .,~--~,,r_-, . , - - - - , - , _ , , , - . , , , , , , ~ ,e.,.,-,n,_v-~.--w,,,-,-,__,m.,.-,n.,,,..-,,,,-, --,,nn,m, n,~~
1s . . , , , , , 16 - - g, - 210 0 4
}
12 -- -
-rno
^ .
q v 10 -
~-"" 1 e NOTE: These tronslents "
l ,_ ossumed multiple ! c _ operator equipment
~
_ ,o3o E I f allurcs. ee rABLE H.
- 6- -
-yon 4- _
! ~ 2- '380 0 , . . . . , , 0 O 1000 2000 3000 4000 5000 6000 7000 8000 Trne (s) Fig. VIII.A.19. Primary system pressure during 0.1-m2 MSLB with two operating RCPs from HZP. 20 . . . . . . 18 -- --40 HPI FLOW PER COLD LEG 16 -- ------ --- CHAR 0mG FtDW PER COLD LIG $40P5 A2 AND 80 --35
~
14 -_
~30
! 3 u- NOTE: These transients - 3 i 6 - assumed multiple - 25 o j io _ operator / equipment - j f ailures. See TABLE H. -20 ,, s- - j
-is 6- -
I
~'
l 4- { 2- -~0 ! 0 . . . .
. . , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Trne (s)
Fig. VIII.A.20.
- HPI/ charging flow during 0.1-m2 MSLB with two operating RCPs 122 from HZP.
-- =-. -. .. ._ - _-
l much earlier, less HPI fluid contributed to the cooldown. A minimum pressure of 4.0 MPa (580 psia) was reached at about 200 s. The repressurization rate slowed at 1000 s when the deadhead of the HPI pumps was reached. The dip in pressure beginning at 1900 s was a result of the primary-side-temperature decrease that occurred when AFW was swept into the riser from the downcomer in SG B because of SG-B overfill. The PORV setpoint was reached at 2500 s. An average flow equal j to the total charging flow exited through the PORVs. Figures VIII.A.21 and VIII.A.22 give the mass flow in the hot legs and the i cold legs. When two of the four RCPs were tripped by the operator the total flow dropped by one-half. The operating RCP in the loop was able to pump more fluid than during steady state because the loop frictional resistance was lower. Some fluid flowed backwards through the cold leg with the tripped RCP. Because two RCPs were operating, the fluid in the primary was well mixed, and the temperature upstream of the HPI/ charging injection ports was about the same as the downconer liquid temperature.
- 4. Comparisons. The 0.1-m2 (1.0-ft )2 main steamline break transients may be compared in terms of initial conditions and pump operation. The transients i
initiated from HZP resulted in a much greater cooldown (Fig. VIII.A.23) because
- 1. the initial mass in the SG at HZP was 50% greater than at FP, and
- 2. the decay heat at HZP was considerably less than from FP shutdown
] (Fig. VIII.A.24). In all three transients, the SGs boiled dry and so the energy removal by the break was proportional to the initial mass in the SG. The net energy removal is compared in Fig. VIII.A.25. The decay heat affected not only the minimum i downcomer temperature but also the rate of reheating after the broken SG boiled i dry. Dryout occurred much later in the HZP transients and reheating was much ! slower. Because less decay heat is available at HZP, transients initiating from HZP are innately more severe in terms of PTS. L To assess the effect of pump operation, it should be kept in mind that the l l initial conditions were HZP. For the case of no pump operation, the flow i stagnated in the intact loop because the heat input from the core was l insufficient to overcome the opposing gravity head created by heat input from i the intact SG. The difference in the minimum downcomer liquid temperature l between forced circulation and no circulation in the intact loop was dramatic. l Heat addition from SG B (Fig. VIII.A.26) was the main contributor to the 40 K (72 F) difference in the minimum downcomer temperature. The additional heat 123 4
,_ _-__~s_3-_ - - - , _ . . - - , _ , , - - . - . , _ _ ,___ -r.-#-, ,,,,y_ . , _ , _ , , , , . , .,,_.__-,.-___-.__w _ , , . , - _ _ _ _ , _ , . . - . . , _ - , , _ , . _ .,,.-m
8000 , , , , , , , , , ,
-15000
*0' '
6000- -
---------- w oP a
-12000
,f 4000
. -9000 d, -
-6000 o j 2000- NOTE: These transienis -
5 n - assumed multiple -3000 5 s operator / equipment j' 0 . f ailures. See TABLE H. - 0
--3000
-2000-
--6000
-4000- , , , , . , , , , ,
0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000 5500 une (s) Fig.VIII.A.21. Hot-leg flows during 0.1-m MSLB with two operating RCPs from HZP. 8000 , , , , , , , , i _ I - -15000 LOOP At 6000- ---------- WOP A1 - wCP B1
- wop e2 -12000 4
i 4000
- . -9000 i
R 3
- 2 -
NOTE: These tronsients -6000 o j 2000- assumed multiple -
?
- - operator / equipment 3000 5 5 f ailures. See FABLE H. j 0 -
-0 l --3000
-2000- -
- --6000
-4000 , , , , . . . . , ,
0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000 5500 Trne (s) Fig. VIII.A.22. Cold-leg flows during 0.1-m2 MSLB with two operatLng RCPs fr a HZP. 124
sco . . i * ' '- ' 1
-600
,,_ NOTE: These transients _
ossumed multiple
- i ..-- operotor / equipmenf
...-',... f ailures. See TABLE H.
- l. -
i 500- *
$ .? C
- -400 .,
E e j m_ sy-N- - - - _ g
- -300 400- -
>f
'i - -200 f 350- From HZP -
.......... From FP
- - from HZP with 2 operating RCPs -10 0 300-
~
250 . . . . . . . 0 1000 2000 3000 4000 5000 6000 7000 8000 Trne (s) Fig. VIII.A.23. Comparison of the downcomer liquid temperature for the 0.1-m 2 MSLB cases.
.360 . . . . . . ,
3"' NOTE: These transients
~
assumed multiple operotor / equipment ....
,_ f ailures. See TABLE H. ,,, ' .
I n goo. ,,.* - i t ,...-
- m. From HZP -
- ,- .......... From FP -
,00 - ,-
! s' So- ' p i ,.- 0-1
-50 , . . . . . .
0 1000 2000 3000 4000 5000 6000 7000 8000 Trne (s) Fig. VIII.A.24. Comparison of the total energy input from the core from HZP and from FP. 125
50 , , , . . . . NOTE: These transients
,_ assumed multiple _
o erator e men t , f il u r es. ee LE H. l T 50-l
\
%g
\/
\
,A ,_
_ 30 0 -
\/ .
From HZP
.........- From FP
~ '
- From HZP with 2 operating RCPs
-200 , . . . . . .
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.A.25. Comparison of the total energy transferred for the 0.1-m2 MSLB cases. 30 . . . . . . . 20- ,..- l
.,,'; From HZP 10 - l }
l '
., ---------- from HZP with 2 operat*cg RCPs l ' '. ,
? 0-l ',
I_ *, l _,o _ , l ' ! NOTE: These transients , l ossumed multiple ', l operator o pr.ent 's, j f ailures. ee BLE I. ', _40 .,, , -W . . . . . . . l 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.A.26. l Comparison of the energy transfer from SG B for the 0.1-m2 126 M B from H7.P cases. l
l input from two operating RCPs was a small factor. From a PTS standpoint, these calculations indicate that for a steamline break in one SG it is preferable to i leave the RCPs in operation. ! B. Double-Ended MSLBs ! These transients were initiated from HZP by the appearance of a double-ended break in the SG-A steamline between the ADV and the MSIV.
- 1. With Unisolated AFW to Broken SG. This transient was initiated by a double-ended break with an assumed failure of the AFW system to isolate the l
broken SG. Table VIII.B.1 tabulates the sequence of events that occurred during this transient. TABLE VIII.B.1 DOUBLE-ENDED MSLB WITH UNISOLATED AFW TO BROKEN SG FROM HZP a ! SEQUENCE OF EVENTS Time (s) Event 0 Double-ended break appeared in SG-A steamline 1 2 AFAS (MFIVs closed caused by model error); pressurizer heaters tripped on low pressurizer level 9 SGIS on low SG pressure; TBVs were isolated ' 19 Asymmetric-SG pressure signal failed to isolate AFW to-SG A i 42 SIAS on low pressure; maximum charging flow and ! HPI delivery began almost immediately 71 Operator tripped all RCPs I ~300 Flow stagnated in Loop B 604 Pressurizer proportional heaters reactivated with level recovery
~1000 HPI pumps dead-headed
~1400 High-level alarm in SG B l
l 2980 PORVs lifted on high pressure 3275 Calculation terminated i I aThese transients assumed multiple operator / equipment failures. l 127 t _ - v v _ ~s... . ~ ,.. y -w. _ - , _ , .-m--.._m _ - - -. _ _ ,. . . . . _ , , _ . _ , , - . , , - _ _ . , , , , _ _ _ _ _ . _ - - - ,..-_,m,,.-,.%-. . . . . .
t As can be seen in Fig. VIII.B.1, the downcomer temperature history has been divided into three phases. Figure VIII.B.2 summarizes the energy transfers into (positive) and out of (negative) the primary fluid. Figure VIII.B.3 shows the system pressure history. The first phase (0--800 s) is characterized by severe overcooling of the primary caused by the rapid blowdown of SG A to l l atmospheric pressure. Although the blowdown was limited by the flow restrictors
~
downstream of the SGs, the initial mass flow out of the SGs was 1500 kg/s (11.9 x 106 lb/h) as a result of significant moisture entrainment. Furthermore, the assumed failure of the asymmetric-SG pressure signal to effect isolation of AFW
- to SG A resulted in a secondary-side-heat-sink temperature of 373 K (2120F).
1, During this period, flow in Loop B stagnated following the RCPs trip because of reverse heat transfer in SG B following SGIS (see Figs. VIII.B.4 to VIII.B.5).
- Also, during this phase, the upper head of the vessel voided briefly (90-350 s) i because the primary fluid contraction initially exceeded the HPI/ charging refilling capacity allowing the primary-side pressure to drop below the saturation pressure of the fluid residing in the upper head (see Figs. VIII.B.6
-to VIII.B.8). A model input error caused closure of the MFIVs on AFAS at 2s instead of on SGIS at 9 s, but this error has no significant effect on the j
results. The second phase (~800-3275 s) is characterized by repressurization of the i primary caused by unrestricted operation of the charging pumps. During this , phase of the accident there is an approximate balance between decay heat, heat transfer from the structure to the fluid and heat rejection to SG A, as indicated by the relatively flat slope of the net energy curve in Fig. VIII.B.2. j_ However, because the HPI and charging flow added substantial mass to the primary l [~46000 kg (101000 lb) during 0-800 s and ~30000 kg (66000 lb) during 800-3275 s to an initial mass of 224000 kg (493000 lb)] but very little enthalpy, the average specific internal energy decreased slightly. By 3200 s, the downcomer temperature had leveled off at 380 K with most of the heat load being dissipated by the AFW added to the broken SG. The energy balance for the primary fluid at ) 3200 s is tabulated for the non-convective terms in Table VIII.B.2 and for the convective terms in Table VIII.B.3. l l r [ 128
590 -
, . . . . -600 560-
~'
530 i NOTE: These tr ans i en t s k assumed multiple , m 5*- operator / equipment - e l ? f ailures. See TABLE H. 40 tg
)2 470 -
I' 440- - W
- p -300 .e 3 410 - 5 -
8 5 380-
% W 2 3 -200 350_
- 1. SG A blowdown _
- 2. Continued AFW to SG A
- 3. Extrapolated 320 -
- -10 0 290 , , , .
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.B.l. Downcomer liquid temperature during double-ended MSLB with unisolated AFW to broken-loop SG from HZP. 10 0 - , , . SLAB CORE 50 ,
,/ -e .
04
/ f - - = = -
'~..7
~~
4
~ -
/
PUMP n NOTE: These transients SG B I assumed multiple _ 3 operator / equipment 6 f ailures. See TABLE I. g , NET e -100 -- O -........,_.....,,,,,,,,,_, , , , ,
-15 0 -
Sg A
-200- % ~ ' '% -
-250 . . . . .
0 500 1000 1500 2000 2500 3000 3500 Time (s) Fig. VIII.B.2. Summary of non-convective energy transfers that occurred during the double-ended MSLB with unisolated AFW to 129 broken-loop SG from HZP during 0-3275 s.
18 . . . .
-2450 16 -- "
A A
" g
- 210 0 g4 _ -
g_ -1750
~
p NOTE: These tr onslent s ? 2 ,o assumed multiple 8. T -
/ operator / equipment ~"00 75 a ! f allures. See TABLE H.
E a
& -1050
[ j r 6- - d d extrapolated
-MO 4-
-350 2-0 . ,
i . 0 0 1000 2000 3000 4000 5000 6000 7000 5000 Tin.e (s) Fig. VIII.B.3. Primary system pressure during double-ended MSLB with unisolated AFW to broken-loop SG from HZP. 800 -
. . . . . . . -rn0 700-- -1500 wop A (coup 20
.00
.. . .. ... .. wo, , go , ,o .
l
-1250 500- -
.J
-1000 400- -
. D
_75g 300- - I 2 200- _i NOTE: These transients -300 a ' i assumed multiple - E P operator / equipmenf -2s0 t f ailures. See FABLE H. - p. O. _ g.,g . -. -.... ,. . . . .. . . . 0
- 10 0 -- ---250
-m0 . . . . . . .
0 1000 2000 3000 4000 5000 6000 7000 6000 Time (s) Fig. VIII.B.4. Hot-leg flows during double-ended MSLB with unisolated AFW to broken-loop SG. 130
400 , , , , , , , LOOP A1(CowP 25)
...-- ---- LOOP A2 (CowP 45)
Loop s:(coup is) 300- _
- toop s2 (COMP 35)
-600 230- -
-- '8 200--
l f, d 50- ~ _ NOTE: These transients -300 d 3 ; assumed multiple 3 10 0 - operator equipmen t - 3 f ailures. ee TABLE H. .o g) 3 50- h o.- N -.- --O
~50~ l
~
_ ..so
-100 o 1000 20'00 3dOO 4000 SdOO 6doo 7000 8000 Time (s)
Fig. VIII.B.S. Cold-leg flows during double-ended MSLB with unisolated AFW to broken-loop SG. 1.2 , , . . . . . THETA CELL 1
.-. THETA CELL 1
~
THETA CEu. 3
- THETA CEu.4 o.8 - THETA CELL 5 o.s-l NOTE: These transients i
f' o.4- assumed multiple - l 1 operator equi ment f ailures. ee TA LE 11. - o.2 - t [ - 0.0 - l 42 . . . . i no l o 1000 2000 3000 4000 sooo sooo sooo TihE (s) Fig. VIII.B.6. Voiding in the inner ring of the upper level of the upper l head during the double-ended MSLB with unisolated AFW to 131 broken-loop SG from HZP.
12 , , , , , , , THETA CELL 1
- THETA CELL 2 THETA C E L 3
- THETA CELL 4 0.8 - -
THETA CELL 5
.h 0.5-h NOTE: These transients l
g' assumed multiple f 0+ operutor / equipment - f ailures. See TABLE H. 0.2 - - 0.0 - -
~01 , , , , , , ,
0 1000 2000 3'XN) 4000 5000 6000 7000 8000 TIE (s) Fig. VIII.B.7. Voiding in the inner ring of the lower level of the upper head during the double-ended MSLB with unisolated AFW to broken-loop SG from HZP. 12 , , , , , , , THETA CELL 1
.......... THETA CELL 2 to- -
THETA CELL 3
- THETA CELL 4 0.8 - -
THETA CELL 5 g 0.6 - - i e t 8 NOTE: These transients
,& 0+ assumed multiple -
operator / equipment f ailures. See TABLE H. ' 0.2 - - ( 0.0 - - l l
-02 , , , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 8000 TIME (s) Fig. VIII.B.8. Voiding in - the outer ring of the lower level of the upper , 132 head during the double-ended MSLB with unisolated AFW to I broken-loop SG from HZP.
TABLE VIII.B.2 1 ENERGY BALANCE FOR THE NON-CONVECTIVE TERMS AT THE TIME OF MINIMUM DOWNCOMER TEMPERATURE (3200 s) FOR DOUBLE-ENDED MSLB WITH UNISOLATED AFW TO BROKEN SG FROM HZP 1 Source Energy (GW-s)a Decay heat +30.2 Primary-side heat slabs +51.3 RCPs + 4.3 SG A -206.2 SG B -8.2 Net -128.6 a + = energy addition to primary fluid.
- = energy removal from primary fluid.
TABLE VIII.B.3 ENERGY BALANCE FOR THE CONVECTIVE TERMS AT THE TIME OF MINIMUM DOWNCOMER TEMPERATURE (3200 s) FOR DOUBLE-ENDED MSLB WITH UNIS0 LATED AFW TO BROKEN SG FROM HZP Source __M(kg) h(kJ/kg) Energy (GW-s)a Charging flow +26560 111.9 +2.97
@ 299.8 K (80*F)
HPI: 6 322.0 K (120 F) + 564 204.5 +0.115 9 302.6 K (85 F) + 4270 123.5 +0.527
@ 285.9 K (55 F) +44606 53.5 +2.386 b
PORVs (T*583 K (590 F)) - 655 1444.6 -0.946 Net 75345 +5.052 a + = energy addition to primary fluid.
- = energy removal from primary fluid.
estimated value. 133
l L. l l l L The change in specific internal energy that resulted from these energy [- transfers was: l 1 ! 3 3
" ~ (224 x 10 )(1227.4 x 10 ) + 5.052 x 109 - 128.6 x 109 l (224 x 103 + 75345) l l
= 505.7 kJ .
.kg l
Using the thermodynamic tables in Ref. 3, the corresponding bulk temperature at the time of minimum downcomer temperature was 393 K (248 F). Hence, the bulk temperature decrease was 159 K (286 F). If the spatial variations of the primary-side remained insignificant during the transient (as it is during HZP), the TRAC-calculated downconer temperature should also have dropped to 393 K. However, the high heat transfer to SG A induced a ~10 K (18 F) temperature drop through Loop A, and stagnation in Loop B trapped a large amount of hot fluid in that loop. As described in Sec. VI, both of these effects tended to depress the downconer temperature relative to the mixed average temperature; hence the difference was expected. The problem was terminated at 3275 s because the transient had stabilized with respect to downcomer temperature and pressure. PORV cycling between , 15.7 MPa and 16.5 MPa would limit the pressure because PORV capacity was more
- .than adequate to relieve the charging flow. Furthermore, the AFW to the broken SG was more than adequate to dissipate the entire heat load produced by the core decay heat, the -heat slabs, and the energy added by the intact SG. However, l
l 'because the heat load was large enough to heat the AFW to its saturation temperature at 0.1 MPa (1 ata) for the specified AFW mass flow rate, the secondary-side temperature is expected to remain constant at 373 K (212 F) for l the remainder of the transient. Hence, the primary-side temperature is also expected to remain constant throughout the remainder of the transient. Figures VIII.B.9 and VIII.B.10 present. flows and temperatures associated l with the HPI and charging flow. The spikes in the Loop-B cold-leg temperatures upstream of the HPI/ charging port in Fig. VIII.B.10 were caused by small fluctuations in the essentially stagnant loop flow that occasionally caused cold HPI/ charging water to flow upstream. (The input errors described in 2 Sec. VIII.A.1 (0.1-m MSLB from HZP) were also present in the model used to 134
--- . - _ . - _ . . - , - . _ - . . _ - - . - . ~ .- . . - - - .... - - , - - - - - -
- 20 , , , , , ,
13.-
.- 40 HPI FLOW Pp COLD EG 16 -. ...---.-.- cuasome now ru coto us OnoPs A2 Ah3 e0 --35 u-
'-30 NOTE: These transients 3 12 - assumed multiple -
3 E operator / equipment ~** j 30 f ailures. See TABLE 11. T
-20 e
m a- }
~
- 15 6-4_.......... .
2- --S 0 , , , . . . , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s)
' Fig. VIII.B.9.
HPI and charging flow during the double-ended MSLB with unisolated AFW to broken-loop SG from HZP. 590 -
, , , , , , , -600 580- -
ggp u .-500 530-( l\ . . .. . .. ... LOOP A2 ! f ,\.q L0oP $1 500- - tDOP e2 _ E 3 e [ ,0 @3 NOTE: These tr ansi en t s _
-400 I=
l i assumed multiple
! operator / equipment I I440-_
i ! ., f ailures. See TABLE H. ** 4 10 - . ,.. - . 3a0- 8
#~7 --
s , d
)' ir~ -200 i'
- 350 i l
(' , l 1 ! 320- .
~
-10 0 290 , , , , , ,
l 0 1000 2000 3000 4000 5000 6000 7000 8000 l Time (s) Fig. VIII.B.10. Liquid temperature upstream of HPI/ charging flow ports during double-ended MSLB with unisolated AFW to broken-loop SG from HZP. 135
analyze this transient. Hence, the occasional oscillations observed are not believed to be real.)
- 2. With Two Stuck-Open MSIVs from HZP. This transient is the same as the one previously described except that the MSIVs failed to close upon receipt of SGIS and blowdown of both SGs continued. Also, (as specified by ORNL) the operator terminated AFW flow at 480 s (8 min). The sequence of events is given in Table VIII.B.4.
TABLE VIII.B.4 DOUBLE-ENDED MSLB WITH TWO STUCK-OPEN MSIVS FROM HZP SEQUENCE OF EVENTSa Time Event 0 Double-ended MSLB on Loop A 11 SGIS; MSIVs failed to close 13 Pressurizer heaters tripped because of low primary pressure 32 SIAS 37 IIPI began 62 RCPs tripped 92 AFAS based on low liquid inventory; AFW to both SGs 480 Operator assumed to have terminated AFW to both SGs 655 Pressurizer proportional heaters turned on following pressurizer level recovery 1250 HPI ended l 2444 PORVs opened because of high primary pressure l 3300 Calculation terminated; quasi-equilibrium reached aThese transients assumed multiple operator / equipment failures. 136
l l The transient .may be divided into three phases as shown on a plot of the dtwncomer liquid temperature in Fig. VIII.B.11. Figures VIII.B.12 and VIII.B.13 ch:w the non-convective energy transfers for the transient. During Phase 1 (0-1000 s), a minimum temperature of 376 K was reached, which was a few degrees cbove the temperature of the. liquid remaining in each SG secondary after the l l bicwdown to 0.1 MPa (14.7 psia). As indicated in Table VIII.B.5, each SG l ! rzmoved ~97 GW-s of energy from the primary. This energy removal resulted from tha blowdown of both SGs and the addition of 7900 kg (17380 lb) of AFW to each SG.. Table VIII.B.6 gives information about the convective energy transfers. At l 1000 s, the decrease in specific internal energy of the system because of these snargy transfers was: 3
, ,(224 x 10 )(1227 x 10 )3 + 4.025 x 109 - 151.3 x 109 (224 x 103 + 59000)
= 451 S .
kg i Thus, from the thermodynamic tables in Ref. 3, the bulk temperature is 380.6 K (225.60 F) giving a temperature drop of 171.4 K (308.5 F). The TRAC-calculated l tssperature in the downcomer was slightly colder because of the spatial variations from HPI/ charging flow injection into the cold leg. The ! ncn-convective energy transfers accounted for 71% of the cooldown. After the AFW ended at 480 s, the primary temperature leveled off a few drgrees above the secondary-side temperature (Phase 2). The downcomer temperature increased slightly after the termination of HPI flow at 1000 s. In extrapolated Phase 3 (3300-7200 s), it is expected that the power from the primary will slowly boil the remaining liquid in each SG (~18000 kg (~40000 lb)). At 3300 s, the power from the heat slabs was ~ 7. 5 MW. Together with 9 MW from the core, a steaming rate of ~4 kg/s would be produced in each l- SG. With this rate as a maximum (heat input from the slabs would decrease in time), the SGs would dry out in another 4500 s (7800 s), which is past the end of the transient. Thus, the temperature is expected to remain at ~378 K (2210F) fcr the remainder of the transient. 137 i i
590 -
. . . . . . . -600 580- -
- 1. Blowdown of both SGs
- 2. Quasi-equilibrium with both SGs --500 530- 3. Extrapolated; equilbrium maintained ,
500-6 - E c NOTE: These transients -400 l I 3 co- assumed multiple - operator / equipment 1 44o_ f ailures. See TABLE H. - g AFW terminated 4 410 -- _ HPI flow ended . sao-- 1 . f t i i 1 - 2 -200 3 350- - 320- ~
-100 200 . . . . . . .
o 1000 2000 3000 4000 5000 sooo 7000 sooo Trne (s) Fig. VIII.B.11. Downcomer liquid temperature during double ended MSLB with two stuck-open MSIVs from HZP. 50 . . . . . . . . i S e 25- '
= ? # 5 o - -
NOTE: These transients CORE PUMP
; assumed multiple -
t operator / equipment y '., f ailures. See TABLE H. 3- '., SG A -
\
=
\ ;
-loo - '
. SG B
_ 23_ ',
..,'s.. NIT
~1So - , ~ ' ' - -- . . . . . . . . . . . . , . . . . . . . . . . . . . - - - - - - - - * *
- 75 . . . . . . . . .
o 10 0 200 Joo 400 500 600 700 800 900 1000 Time (s) Fig. VIII.B.12. Summary of non-convective energy transfers that occurred 138 during the double-ended MSLB with two stock-open MSIVs from HZP during 0 to 1000 s.
TABLE VIII.B.5 i ENERGY BALANCE FOR NON-CONVECTIVE TERMS AT TIME (1000 s) 0F MINIMUM DOWNCOMER TEMPERATURE FOR DOUBLE-ENDED MSLB WITH TWO STUCK-OPEN MSIVS FROM HZP I Source Energy _(GW-s)a l.
- j. Decay heat + 9.4 i Primary-side heat slabs + 33.1 RCPs + 1.1 SG A - 96.2 SG~B - 98.7 i
Net -151.3 a + = energy addition to primary fluid.
-= energy removal by primary fluid.
TABLE VIII.B.6 ENERGY BALANCE FOR CONVECTIVE TERMS AT TIME (1000 s) 0F MINIMUM DOWNCOMER TEMPERATURE FOR DOUBLE-ENDED MSLB WITH TWO STUCK-OPEN MSIVS FROM HZP ? Source M(kg). h(kJ/kg) Energy (GW-s)* Charging flow + 8300 111.9 +0.929 9 299.8 K (80*F) i HPI: 0 204.5 +0.115 9 322.0 K.(120 F) + 564 i 9 302.6 K (8500F) + 4270 123.5 +0.527 9 285.9 K (55 F) +45866 53.5 +2.454 PORVs - - Net +59000 +4.025 l
* + = energy addition to primary fluid.
- = energy removal by primary fluid.
I I l 7 139
~ , . . - . _ _ _ _ __ _ _ _ , _ _ _ . _ , . . . ___ _ _ _ _ . _ _ _ _ _ , , _ . _ .
_ - _ - _ _ ,,-- _ _. 1
Figurta VIII.B.14 and VIII.B.15 give the system pressure and HPI/ charging flows. The blowdown of both SGs caused the system to depressurize to 4.1 MPa. HPI flow reached a maximum of 60 kg/s to make up for the primary liquid contraction. He upper head voided during 50-900 s. Charging flow eventually i repressurized the primary system to the PORV setpoint where it would be expected to remain until 7200 s. Figure VIII.B.16 shows the mass flows in the cold legs. Because the secondary on both SGs behaved identically, conditions on the primary were symmetric. Natural circulation of ~200 kg/s per loop was induced by the gravity head created from heat addition by the core. H e liquid temperature upstream of ~ the 'HPI/ charging ports (Fig. VIII.B.17) was the same as the downcomer liquid temperature except for the dip that was caused by the HPI flow into the downcomer during 38-1250 s.
- 3. Comparisons . He two double-ended MSLB transients differed both in AFW delivery and the operation of the MSIVs. In one, AFW was assumed to be delivered erroneously to the broken SG. In the other, AFW was terminated at 480 s and the MSIVs failed to close, so that SG B was not isolated from- the break. A comparison of the downcomer temperature history and the net energy transferred for the two transients (see Figs. VIII.B.18 and VIII.B.19) shows that the cooldown produced by the double-ended break in which the MSIVs failed to close is about ~50% faster than the cooldown produced by a double-ended break with runaway AFW to the broken SG. Until 480 s, the difference in the curves resulted from the stuck-open MSIVs. Af ter 480 s, AFW was terminated in the case with the stuck-open MSIVs but continued in the other case. Nevertheless, the temperature leveled off at the atmospheric boiling temperature in both cases.
In the case with two stuck-open MSIVs the liquid in the SG secondary was boiling and venting to the atmosphere through the break. In the other case the primary l temperature was dominated by energy removal through the broken SG, which was also at the atmospheric boiling temperature. Two significant differences occurred between these two similar downcomer-temperature history transients. First, neither loop stagnated in the case of the stuck-open MSIVs. B is resulted because of the symmetric blowdown that occurred in each SG. For the other case, the broken SG was dissipating the vast majority of the heat load, whereas the intact SG remained constantly in a reverse-heat-transfer mode. His resulted in the stagnation of the intact loop. The second dissimilarity is the long-term cooling potential. he unisolated AFW 140
75 . , . . . . SLAB 50- - r RE 25- - o, , _ _ ; _ _ _ _ _ _ - y PUMP - 4 i i g NOTE: These transients assumed multiple
! SG A
[ ! CPerotor / equipment - i f ailures. See TABLE H. _too- k. - I
' ~
- n5 - ',
SG B Nu - _ iso- '.,,,_,,,,,,......-+----*----+---*----+---*-----
- 175 . . . . . .
0 500 1000 1500 2000 2500 3000 3500 Time (s) Fig. VIII.B.13. Summary of non-convective energy transfers that occurred during the 0.1-m 2 MSLB from HZP during double-ended MSLB with two stuck-open MSIVs from HZP during 0 to 7200 s. ts . . , , ,
-2Aso a .u2 am , i A is - ,,. . . py - - - -
- 210 0 g- -
12 - NOTE: These transients -- C50 assumed multiple { '- operator / equipment ? o f ailures. See TABLE H. -- wCO I 2 h s- - h
$ i extrapolated
-1050 g
( s-l -
- 700 4- -
- 350 2- -
0 , , . , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.B.14. Primary system pressure during double-ended MSLB with two g stuck-open MSIVs from HZP.
20 . . . . . . . 18 -' - -40 j HPl FLOW PER C04.0 GG ! 16 -- ......-.-- cHARCm3 FWW PER COLD MG 0 00PS A2 AhD D1) --3$
"~
~-3o
'~
NOTE: These transients
~
-25 $
g c '~ assumed multiple - 7 operator / equipment e j a-f ailures. See TABLE E .
~* a j
-15 s- -
s
. .................... - 10 4
2- --S 0 , , . . . . . 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.B.15. HPI/ charging flow during double-ended MSLB with two stuck-open MSIVs from HZP. 500 . . . . > > .
-1050 450 - -
1.00P At(CowP 25)
..... ..... Loop 42 (cop as) 400-- toop et (cow is) _-MO
- Loop s2 (CowP 36) 350-_ - 750 300- -
E NOTE: These transionts $^ j assumed multiple 5a T
~
i 250 operator / equipment ' e l 200
- f ailures. See TABLE E -- 450
- a 2
20- -
-300 l
10 0 - M -
- 15 0 50- -
0 , , . . , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.B.16. Cold-leg flows during double-ended MSLB with two stuck-open gg MSIVs from HZP. l
i 4 M , . . . . . SLAB _ 50- - l 25- - 0, -- - - - - - - - - - - y PUMP -
@ { NOTE: These transients
! SG A essumed multiple j operator equipmen t - 1 i f ailures. ee TABLE I.
\
_,oo . 1, -
' ~
_ g$ - '., SG B gi 3 - _ iso. ,,,_ ,,,, .....-+----*----*----'---+-~~+"""- i -m . . . . . .
! O 500 1000 1500 2000 2500 3000 3500 Time (s)
Fig. VIII.B.13. Summary of non convective energy transfers that occurred during the 0.1-m 2 MSLB from HZP during double-ended MSLB with two stuck-open MSIVs from HZP during 0 to 7200 s. is . . , , , .
- 2460 a .to ui , i A 16 - * * " ~
,,. ,r pq
- 210 0 j 14 - -
i g_ - NOTE: These transients -
-1750 assumed mult.iple operator equipment 7 o
l f ailures. ee TABLE H. -- w00 L $ s- - N
^
$ - extrapolated 6- -
-700 4 -
-350 2- -
0 . . . , . . 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.B.14. Primary system pressure during double-ended MSLB with two g stuck-open MSIVs from HZP. i
-- --- ,..w-r,----, - - - , - - - - - , - , ,,,-w,
20 , , , , , , , l is -- . -40 ; um vton m coto uG 16 - -
.........- CHAlfClhG FLOW PUI COLD uG Q.00P5 A2 Aho DQ --3$
i 14 -
.-30 l
'*~
NOTE: These ironsients
- -25 $,
c g
'~
assumed multiple , c operator / equipment ~
~'" d
-j a-f oilures. See TABLE H. .
j a
- 15 s- -
u
- 10 4 _
1 --S 0 , , , , , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.B.15. HPI/ charging flow during double-ended MSLB with two stuck-open MSIVs f roin HZP. 500 . , , . , , i
-1050 450- -
Loop at(coup 25)
.......... toor42(co. pas) 400-* LDOP st (COWp 2) _-900
- Loor 32 (cour as) s50- .
73a f3 - g NOTE: These ironsionts ossumed multiple
~' f
= .
** ~ ~ '
C operator / equipment d j ,,,_. f ailures. See TABLE H. __43,
=
y i 15 0 -
-300 10 0 -
- - 15 0 So_
l 0 , , , , , , , 0 O 1000 '2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.B.16. Cold-leg flows .during double-ended MSLB with. two stuck-open 142 MSIVs f roin HZP.
$90 -
. . . i i . -600 560-
' -~#
550- LocP As
.-. . . . . . . wap g2 LDoP 01 500 - Loor 62 -
e -400 , j d' ~ NOTE: These transients ) l 1 d' ~ assumed multiple ~ E o erator / equipment @
~3
, 0 f ilures. See TABLE H. 4 g 410 - 380-a
- -200 350-320-
- -100 290 . . . . . . .
0 1000 2000 3000 4030 5000 6000 7000 8000 T;me (s) Fig. VIII.B.17. Liquid temperature upstream of HPI/ charging flow ports during double-ended MSLB with two stuck-open MSIVs from HZP. 800 . . . . . . . 580-
-500 n00- I -
E nonHze -- E
-i
.0 i, ....... nom Hzr with stuck-open usw. -
i s t -300
. i t
400- ',
. .. ....................................+........ -200
"~ ~
NOTE: These transients assumed multiple _, operator / equipment m- - f ailures. See TABLE H. m O SCO 1000 1000 2000 2500 3000 3500 4000 A Time (s) Fig. VIII.B.18. Comparison of the downcomer liquid temperature for the double-ended MSLB cases. M3 1
- - . _ _ _ . _ . . _ . , _ . . . _ _ . _ _ _ . _ - . . . _ . _ _ _ . . _ _ . . . _ . . _ . _ , , _ . . , . . _ . _ _ = . . _ _ , , _ _ _ . . . . _ _ _ _ _ , _ . _ , _ . , - . _ .
50 , , . . . . . i NOTE: These transients assumed multiple 0- operator / equipment - f ailures. See TABLE H. rtom HZP 7 i g .......... From HZP with stuck-open MSNs i i
-i00- {
t
\
\\ ~ -
-a0- ' ,,,,,__.. ....-- ------
-200 , , , , , , ,
O SCO 1000 1500 2000 2500 3000 3000 4000 Time (s) Fig. VIII.B.19. Comparison of the total energy transferred for the double-ended MSLB cases. I F i I ( 144 i t
- - - . _ _ - - _ . _ _ . . . ~ - . . _ . - - . - _ _ _ _ . , _ . _ _ _ _ _ _ _ - - . _ _ . - _ - - , _ _ _ _ _ _ _ _ _ _ . - _ . . _ . _ _ _ . _ _ _ _ . - . - _ . . . _ _ _
. case represents a ~ long-term (~7 hours) cooling mechanism, whereas the primary-side temperature will begin to heat back up in the failed MSIV case once
. the SGs dry out at ~8000 s.
~ C. Small Steamline Breaks (Stuck-Open TBV)
The failure of one TBV to reseat after opening on a turbine trip is postulated in these transients. One full-open TBV (0.05 sE (0.51 ft2 )) is about half the size of the 0.1-m 2 (1.0-ft2 ) break described previously. )
- 1. From FP. Because the TBVs are. downstream of the MSIVs, a stuck-open TBV is isolable whereas the 0.1-m3 MSLB described previously was not. A sequence of events is given in Table VIII.C.1. The stuck-open valve l
. communicated with each SG identically and so the thermal-hydraulic events on both the secondary and primary sides are symmetric.
The temperature history in the downcomer (Fig. VIII.C.1) may be divided 4 into five phases. Phase 1 (0-510 s) was the time before the stuck-open TBV was isolated from the SGs as a result of the closure of the MSIVs following SGIS. , The initial ~50 s of the transient should have been identical to a loss-of-load transient. Three of the four TBVs reseated as the primary temperature decreased. When one failed, a relatively slow depressurization began in both SGs. The secondary pressure decreased until the setpoint for SGIS was reached. This marked the end of the cooldown caused by the stuck-open TBV. The non-convective energy transfers are given in Table VIII.C.2. Figures VIII.C.2 and VIII.C.3 summarize the non-convective energy transfers for 4 - 0-1000 s and 0-7200 s. SG A and SG B removed about the same amount of energy. Decay heat was high because the transient was initiated from FP (initiation from HZP would be less probable because the TBVs are barely open and the turbine is latched). Heat input from the primary-side heat slabs and the RCPs was f negligible. i
*Because of an error in the initial liquid temperatures in the pressurizer, the i- primary side depressurized much too rapidly. This calculation was to be redone,
- but because it was already predicted not to be of PTS concern for the specified
( initial conditions, an additional failure of one MSIV was specified. The i recalculation is reported in the next section. The period (0-570 s) before SGIS in that transient was identical to the specifications of this transient. The transient described in this section is included to give details of a 7200-s I transient with the failure of one TBV only. i i 145 i
.-r- -- _ . . , .,,,u. -,..rm_w , ,,,,..,.__,_mumes, .,_,.,,,,_.pym,-,w_,_-
TABLE VIII.C.1 ONE STUCK-OPEN TBV FROM FP SEQUENCE OF EVENTSa Time (s) Event 0 Turbine / reactor trip; ADVs and TBVs opened automatically on a turbine trip j 28 SIAS on low primary pressure (error in calculation) l 38 All pressurizer heaters tripped on low pressuriser level 50 All four TBVs should reseat but one failed' 58 Operator tripped all four RCPs 510 SGIS on low secondary pressure 835 Pressuriser proportional heaters turned on after level recovery 1050 ADVs opened on high primary temperature 1100 Pressuriser proportional heaters tripped because of high system pressure 1270 PORV setpoint was reached i 4200 AFAS based on low level; AFW to both SGs 4450 ADVs closed because of low primary temperature 5800 Calculation terminated with minimum downconer temperature at 7200 s estimated to be 510 K (459 F) l I i "These transients assumed multiple operator / equipment failure. The convective energy transfer was caused by charging flow only and caounted to 0.473 GW-s. HPI flow did not actuate because the system pressure l did not fall below the shutoff head of the HPI pumps. The total change in 4 cpecific internal energy at the time of minimum downconer temperature resulting from the stuck-open TBV (end of Phase 1) wast 146
I TABLE VIII.C.2 ENERGY BALANCE FOR NON-CONVECTIVE TERMS AT THE TIME OF MINIMUM DOWNCOMER TEMPERATURE (510 s) FOR THE STUCK-OPEN TBV FROM FP Source Energy _(GW-s)a Decay heat + 46.7 Primary-side heat slabs + 3.8 RCPs + 1.0 SG A - 44.9 SG B _ 46.3 Net - 39.7 8 + = energy addition to primary fluid. , - - = energy removal from primary fluid. 3 3 u = (219 x 10 ) (1332219 x 10 ) + 0.473 x 109 - 39.7 x 109 x 103 + 4146 l
= 1131 S .
kg From the thermodynamic tables in Ref. 3, this internal energy corresponds to a 0 bulk temperature of 533 K (5000F) and a temperature drop of 39 K (70.2 F). The bulk temperature closely matches the TRAC-calculated downconer temperature l because both loops were in natural circulation, which kept the fluid well mixed. Phase 2- (510-1050 s) was a time of primary fluid heating that ended with , cpening of the ADVs on high primary temperature. Boiling on the SG secondary continued to remove energy but at a slower rate as the secondary repressurized. j. l .The ADVs were open in Phase 3 (1050-4200 s), modulating to maintain the average 0 l primary temperature at 552 K (534 F). The TBVs also opened, but they had no offect because the MSIVs were closed. Boiling in the SGs continued and mass was depleted through the ADVs. AFAS tas obtained at 4200 s based on low level in both SGs.* Phase 4 (4200-5800's) l
#AFAS was based on a Ap measurement of -4.3 m (-170 in). This corresponded to a ;
liquid inventory of ~17000 kg. Based on a collapsed liquid measurement, AFAS would occur with 45000 kg remaining in the SGs. It is unknown which method is more correct, but AFAS probably was obtained later than it should have been. 147
. -.- . _ - _ , . - . _ - _ . - - . . , _ . ~ , , - . - . - . _ . - - . . - . . - . . - .
590 -
. . . . -600 560-530-- '
-500 5
500-E ~ F
$ 470-Ts .i2 i t e .
p 440- 1. Slow blowdown of both SGs - E
,E through TBV E
, -300 e
*I 410 - 2. Heating from core
.E' }
- 3. ADVs regulated primary temperature ,
380- ~
- 4. AFW to both SGs
-200
- 5. Extrapolated 350- -
NOTE: These transients assumed multiple 320- operator / equipment - f ailures. See TABLE I. ~S0 290 . . . . 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.C.l. Downcorner liquid temperature during stuck-open TBV frorn FP. 20 , , , 75-NOTE: These fronsients ossumed multiple CORE - failures se AL I. f 50- /**'
/
~ S B 8
,/ ;.CL o._
h ' l PUMP
\.....,,,
- NET
_f. . . 5o_ **===--m- - - - - -
)
SG A 1 SG B i _y$
-10 0 , . , . . . 5 . i 0 10 0 200 J00 400 %0 t,00 700 500 900 1000 Time (s)
Fig. VIII.C.2. Suasary of non-convective energy transfers that occurred during stuck-open TBV from FP for 0 to 1000 s. 148 i I l
began with AFW to both SGs. A cooldown ensued as the AFW mixed with the boiling liquid in the riser region. AFW flow affected the primary temperature in this I transient more than in others because 'it was initiated to both SGs (no c0ymmetric-SG pressure signal) and both SGs were low in inventory (perhaps too icw, see ' footnote). Also, both loops were in natural circulation on the primary; this allowed rapid feedback to the primary side. The cooldown is supected to continue at the same rate until 7200 s, reaching a minimum of ~510 K (4580 F). Phase 5 (5800-7200 s) is the extrapolated temperature history. The pressure history for this transient is given in Fig. VIII.C.4. The system pressure was never low enough for HPI flow. Energy removal, and consequently depressurization, ended at 510 s when the SGs were isolated. As rantioned earlier, the initial depressurization was too rapid because all the initial liquid in the pressurizer was not saturated. SIAS should not have been r ached at 28 s. In fact, had the presssure been correctly predicted, SIAS eight never have been obtained and the RCPs would not have been tripped. With the RCPs running throughout this transient, the coupling between the primary and escondary would have been much stronger. This would not only affect the
" timing" of the sequence of events in a significant fashion, but also change the timperature history of downcomer. Nevertheless, it is our opinion that the transient would still have been benign from a thermal-hydraulic standpoint bscause the overall cooling potential is terminated once SGIS closes the MSIVs end isolates both SGs from the break. Hence, although the results presented for this transient are not an accurate representation of a stuck-open TBV transient, from a thermal-hydraulic standpoint, the transient is unlikely to be of PTS concern if initiated from this initial decay heat level.
After the initial depressurization that was caused by the reactor / turbine trip (which would have brought the system to about 13.2 MPa), a slow depressurize. tion continued because of the slow blowdown of both SGs. Charging flow repressurized the system to the PORV setpoint after energy removal ceased at 510 s. The pressure is expected to remain at the PORV setpoint until 7200 s. f Figure VIII.C.5 is a plot showing the flow in the cold legs (which is ! similar to the flow in the hot legs). Following the trip and coastdown of the l RCPs, both loops were in natural circulation driven primarily by the heat input from the core. The mass flow increased slightly af ter 4200 s as AFW cooled the primary and created an additional gravity head in each of the SGs. Because both loops were flowing throughout the transient, the temperature upstream of the l 149 i
300 . . . . .. . . . . . CORE 250 - NOTE: These transients , assumed multiple operator / equipment , 200- f ailures. see TABLE I. # _ i
/'
15 0 ,/ [ . l f 10 0 - - So. SIAB
^ PUMP d 0- ; 8-= -
- 4. . .===+ - e-@ Y Z - -
" ' _ , . . . , - - * * , , , , , ,/
NET .
- 10 0 - -
_ SG A and SG B ,
-200 . . . . . . . . . . .
0 500 1000 1500 2000 2500 3000 J500 4000 4500 5000 5500 6000 Time (s) Fig. VIII.C.3. Summary of non-convective energy transfers that occurred during stuck-open TBV from FP for 0 to 7200 s. 18 , , , , , A A -2450 16 - g I f' ' y M " -
-110 0 i2 NOTE: These fransients .
1730 assumed multiple { e operator / equipment f allures. See TABLE H. - '4a 1 a b 8- -
-1050 g 6-
'. A A _ sun a extrapolated 799 4-
-350 0 . . .
.0 0 1000 2000 4000 4000 5000 6000 7000 8000 Tim. (s)
Fig. VIII.C.4. Primary system pressure during stuck-open TBV from FP. 150
I HPI/chrrging porto w23 virtually th2 same es thtt in tha downcom;r (Sae Fig. VIII.C.1).
- 2. With Stuck-Open MSIV f rom FP. This transient is presumably the same as the previous transient with the additional failure of the MSIV on one loop after SGIS. Thus, one SG blew down completely in this transient. A sequence of events is given in Table VIII.C.3.
TABLE VIII.C.3 ONE STUCK-OPEN TBV WITH ONE STUCK-OPEN MSIV FROM FP SEQUENCE OF EVENTSa Time (s) Event
-0 Turbine / reactor trip; TBVs and ADVs automatically opened on turbine trip 39 Pressurizer heaters tripped on low level in pressurizer 85 One TBV stuck open while closing on low primary temperature 135 ADVs close on low primary temperature 422 AFAS on low SG collapsed-liquid level: AFW to both SGs
- 470 SIAS on low primary pressure 500 RCPs tripped 570 SGIS on low secondary pressure; MSIV-A stuck open 639 Asymmetric-SG pressure signal
- AFW valved out to SG A 1750 SG A dried out f 2092 Pressurizer proportional heaters turned on because of level recovery 2400 Pressurizer proportional heaters tripped because of high system pressure L
l 2500 PORV setpoint reached t 2500 Calculation terminated I CThese transients assumed multiple operator / equipment failure. 151
1 I As shown in Fig. VIII.C.6, the downconer liquid temperature has been divided into four phases. Both SGs blew down through the stuck-open TBV during Phase 1 (0-570 s). The end of this phase was marked by the closure of one MSIV and the failure of the other MSIV after SGIS. Phase 2' (570-1750 s) was a period of asymmetric SG conditions. One MSIV closed isolating SG B from the stuck-open TBV, while SG A continued to blow down. AFW was delivered to both SGs until asymmetric SG pressure was detected l at 640 s. AFW was then delivered to only SG B. Some azimuthal differences in the downconer temperature existed because-higher heat-transfer rates caused the - primary fluid to flow preferentially to Loop A. The dryout of SG A marked the end of Phase 2. Figure VIII.C.7 gives the energy. transfers for 0-2500 s. The non-convective energy transfers during Phases 1 and 2, corresponding to the time of minimum downconer liquid temperature, are given in Table VIII.C.4. The heat input from the core dominated the heat input from the heat slabs and RCPs. The entire energy-removal capability of one SG was 125.3 GW-s, which included 220 s of AFW flow (~4400 kg) at 277 K (40 F). 0 SG B removed 56.6 GW-s before SGIS at
.570 s. The only convective energy transfer was from the charging flow; this amounted to +1.625 GW-s. The system pressure was never low enough for HPI flow (8.8 MPa (1270 psig)).
! TABLE VIII.C.4 ENERGY BALANCE FOR NON-CONVECTIVE TERMS AT THE TINE (1750 s) 0F MINIMUM DOWNCOMER TEMPERATURE FOR THE STUCK-OPEN TBV WITH A STUCK-0 PEN MSIV FROM FP Source Energy (GW-s)a Decay heat + 109.1 Primary-side heat slabs + 12.5 RCPs + 8.7 ' SG A - 125.3 s-3 SG B - 67.8 Net - 62.8 r 1 a + = energy addition to primary fluid.
- = energy removal from primary fluid.
152
500 , , 1050 450-Loop A1(coup 25)
- -- ----- LOOP A2 (COMP 45)
~
Loor si(cowP is) 900 400
- Loop e2 (coup 35)
~
35 -- NOTE: These transien t s - isc g assumed multiple o, 6 300- o erator f il u r es. e ont 3 o
. ee LE H. -s00 3
D g 250-g n
-450 j S 2c0--
15 0 -
- -300 10 0 --
- 15 0 50-0- . . . .
0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.C.5. Cold-leg flows during stuck open TBV from FP. 590 -
. . i i i -600 560- S _
G A dryout 530--~l g g g --500 500- 2 s
# 3 4 -
b F
$ 470 -
- 1. Blowdown of both SCs through TBV }* 1 440- 2. Continued blowdown of SG A -
- 3. Heating by core after SG A was lost as f I
"h" -
a heat sink -300 e j 410 - 4. Extrapolated; quasi-equilibrium reached - between core and SG B/PORVs 380- ! NOTE: These transients 2# 350- assumed multiple - l operator equipment
- 320 f ailures, se TABLE H. .
' - 10 0 290 i , . . . f O i 1000 2000 3000 4000 5000 6000 7000 8000 l~ Tirns (s) Fig. VIII.C.6. Downcomer liquid temperature during a stuck-open TBV with a l stuck-open MSIV from FP. i 153
; These non-convective and convective energy transfers translated into a specific internal energy on the primary of:
1
, , (219 x 10 )3 (1332 x 10 )3 + +.625 x 109 - 62.8 x 109 219 x 103 + 14525
= 987 S .
kg 1 0 Thus, the bulk temperature is 503 K (446 F) from the thermodynamic tables in Ref. 3. The bulk temperature drop was 69 K (1240F). The non-convective energy transfers accounted for 81% of the cooldown with charging flow contributing 19%. Phase 3 (1750-2500 s) was a period of primary heating after SG-A dryout. The PORVs had not yet opened so SG B was the only heat sink for the energy deposition from the core. Phase 4 (2500-7200 s) was extrapolated based on the 0.142 MSLB from FP (the original run for 0-7200 s). The heatup to a quasi-equilibrium state should be similar for both transients because the energy transfers were nimilar. In both transients, SG B and the PORVs were removing the decay heat, and the primary-side heat slabs, RCPs, and SG A no longer influenced the transient. A quasi-equilibrium state is expected to be reached at ~525 K (4860 F). Figure VIII.C.8 shows the pressure history. The first 50 s corresponded i
- to a normal loss-of-load transient. When one TBV failed to resent at 50 s, the
- pressure continued to drop with a sharp decrease after the RCPs were tripped at 500 s. The pressure leveled at 11.2 MPa as the cooldown slowed and the primary [
. liquid contraction ended. The PORV setpoint was reached just as the calculation l was terminated.- The system pressure is expected to have remained at the PORV setpoint until 7200 s. SIAS was received at 470 s, but the system pressure was
! never low enough for HPI. , Figure VIII.C.9 gives the flow in the cold legs (similar to the flow in , the hot legs). After the RCPs were tripped, the blowdown of SG A induced a i higher flow in that loop. Although reverse heat transfer did not take place and stagnate Loop B, the flow slowed considerably until SG A dried out at 1750 s. After this, flow went preferentially to Loop B. An ~20-K (360 F) temperature j difference between the two loops upstream of the HPI/ charging ports existed 154 i L
200 , , , , NOTE: These transients assumed multiple 15 0 - operator e ui ment f ailures, se LE I. cou /~ 10 0 -
/
50- SIAB - S h e 0, - _g 3 ; c-- -
.. : i
" PUMP
- ,- ..... , ,, Y ri - - , v,,......-*--'"*-
- 10 0 _ SG A .
3C , xJ =
-tSO , , , .
0 250 500 750 1000 1:50 1500 1750 2000 2250 2500 Time (s) Fig. VIII.C.7. Suanary of non-convective coergy transfers that occurred during a stuck-open TBV with a stuck-open MSIV from FP (0 to 2500 s). 18 , , .
-2460 16 -
g A A
-2100
- y. -
g_- -1750
?
' a 10 - ,
.-400 O 8- -
I N
-1050
[
,_ A A extrapolated _
- 700 4- NOTE: These transients -
assumed multiple 2~ operator / equipment - 350 f allures. See TABLE H. 0 , , , , , . - 0 0 1000 2000 J000 4000 S000 6000 7000 8000 Time (s) Fig. VIII.C.8. Primary system pressure during a stuck-open TBV with a stuck- 155 open MSIV from FP. L
[ before SG A dried out (see Fig. VIII.C.10). As SG B became the heat sink after 1750 s, Loop B became the colder loop.
- 3. Comparisons. These two transients should have been identical until SGIS when the failure of one MSIV in one transient allowed the continued
. blowdown of one SG. The reasons for the discrepancies in the calculations I before SGIS have been explained previously.
Figures VIII.C.11 and VIII.C.12 compare the downcomer liquid temperature i cnd total energy removal for the .two transients. The downconer liquid temperature was approximately the same until SGIS around 570 s. Cooldown continued when the MSIV stuck open until the affected SG dried out (at 1750 s). Aside from the stuck-open MSIV, the major difference in the calculations lay in the prediction of the time for AFAS. In the case with only a stuck-open TBV, a Ap measurement for liquid level was used and only 17000 kg (37400 lb) remained in each SG at AFAS. Thus, AFW was not available for 4200 s, which c11 owed the primary to heat up to the ADV setpoint (the ADVs modulate to maintain a specified primary temperature). The primary temperature began decreasing when AFW initiated. In the other calculation, AFAS was based on a liquid inventory of 45000 kg (99000 lb), which occurred at 420 s. AFW together with the stuck-open MSIV continued to cool the primary after SGIS, and so the
.downconer liquid temperature responded similarly to the 0.1-nd MSLB from FP case but more slowly and less severely because the break was smaller (see Fig. VIII.C.13).
] Notice that before SGIS, the energy removal was higher for the case with the stuck-open MSIV while the downcomer temperature was also higher. This was en effect of the RCPs. Because the RCPs tripped at 500 s in the second case
; compared to 58 s in the first case, more energy was removed by the SGs. The cperating pumps also kept the fluid well mixed, minimizing the spatial variations.
i 4 l Y l' 156 1
500 , i i i
-1050 450-LOOP A1(Coup 25)
-- - ----- LOOP A2 (COW 8 45) 4gg..~ - LOOP 01(Coup IS) -900
- LOOP 02 (CCWP SS) 350- _ _g 300-NOTE: These transients o T 250 -
assumed multiple
-600 e C 1
. operator / e pment j 200 I f allures. See BLE H. . 43o @
,A 2
\
-300 10 0 --
-150 50-0 , . , . . , . 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s)
Fig. VIII.C.9. Cold-leg flows during a stuck-open TBV with a stuck-open MSIV from FP. 590- , . i i -600 560 530-- N '- Loop A:
. - . . . . . . . gogp p g tour si 500 - LOOP s2 m C I$ 470-
-400 Ig O **
13 8 Q-440~ E h NOTE: These transion'ts E assumed multiple ~300 4 operofor men t 410 - f oliur es. eee uigLE A L g g 380-j -200 350-- 320-tuo 290- , . . . 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. VIII.C.10. Temperature upstream of the HPI/ charging port during a stuck-open TBV with a stuck-open MSIV from FP. 157 l
800 . . . . .
-600
_$og M* - 8 -
-4oo C
o NOTE: These transient s e mo- assumed multiple - operator / equipment f ailures. See TABLE H. ~* 400- - A 3
. From FP _m g 300- ~
---------- From FP with stuck-open MSN
- 10 0 300- -
250
~
~0 0 1000 2000 3000 4000 5000 0000 Trne (s)
Fig. VIII.C.ll. Comparison of the downcomer liquid temperature for the stuck-open TBV cases. l So . . . . . o- - i Y --
.. ,,,,,,,, , ,. ~~~~
~
l l mo. - ! Rom FP l j .......... Rom FP 4 stuck-open MSN I
-mo- NOTE: These transients
~
assumed multiple operator e pmen t f ailures. ee BLE H. '
-200 . . . . .
o 1000 2000 3000 4000 S000 0000 Time (s) Fig.VIII.C.12 Comparison of the total energy transferred for the stuck-open TBV cases. 158
600 . > ' ' ' ' '
-600 550- .
1, . A A A .$oo l ,,..*****' M " " 500- ,j' -
$e '
\/
g g extrapolated -400 C L 450- .
)
- 8. NOTE: These transients @
E assumed multiple - 3M - o-(5 400 operator / equipment f ailures. See TABLE H. g5 v
~3 5
350-
-200 $
Stuck-open TBV and MSIV from FP 300- .
--- ------ 0.1-m' MSLB from FP 250- , ,
'O O 1000 2000 3000 4000 5000 6000 7000 8000 Time (s)
Fig. VIII.C.13. Comparison of the downcomer liquid temperature for the stuck-open TBV and MSIV and the 0.1 m 2 MSLB (both from FP). 159
IX. SMALL-BREAK LOCAs A small-break loss-of-coolant accident SBLOCA may pose a PTS threat if the break is large enough to generate significant overcooling yet small enough to sustain relatively high pressures. In the absence of SIS flow, the
.depressurization caused by a loss-of-coolant accident will cause the primary system to follow the saturation curve--a condition that is not likely to induce !
PTS. The break must be large enough to depressurize the system to the SIAS I setpoint, if it is-to generate PTS. However, if the break is too large, the rate of depressurization will be sufficient to maintain a pressure-temperature relationship close to the saturation curve despite the effect of the cold SIS water. Because the HPI flow rate is strictly a function of system pressure, reasoning suggests that the threat of PTS will be increased by any mechanism that localizes and concentrates the effect of the HPI water in the vicinity of the critical vessel welds. One such mechanism is loop stagnation. Loop stagnation not only localizes the HPI effect along the downcomer wall by promoting stratification in the cold legs, it also inhibits reverse heat transfer from the hot SGs that would mitigate the effect of the HPI. Consequently, there is some concern that certain break sizes may generate conditions conducive to loop stagnation yet limit depressurization sufficiently to cause PTS. This phenomenon has been discussed in greater detail in Sec. VI. To address this concern, two small-break LOCA transients were selected for investigation. The first was a small hot-leg-break LOCA with a break size (0.002 m2 (0.02 ft 2)) in the range suspected of causing loop stagnation. For that calculation, the FP model was modified to include a break in the hot leg of Loop A with a prescribed pressure boundary condition of 0.1 MPa (14.7 psia). The second transient was a small small-break LOCA caused by the failure of one of the two PORVs to close fully (.001 s3 (0.01 ft2 )). In addition, it was assumed that the SG-A ADV failed to close when it should have. These two j transients are described in the following sections. A. 0.002-m2 Hot-Leg Break from FP This transient was initiated from FP by the appearance of a 0.002 af 1 (.02 ft 2) hole in the Loop-A hot leg. Table IX.A.1 tabulates the sequence of svents that occurred during this transient. As can be seen in Fig. IX.A.1, the downcomer-temperature history was divided into two phases. Figure IX.A.2 shows the system pressure history. 4 Figure IX.A.3 summarizes the mass transfer into (positive) and out of (negative) 1 160 s
.---+.., --
TABLE IX.A.1 0.002-m2 HOT-LEG BREAK FROM FP SEQUENCE OF EVENTSa i Time (s) Event 0 0.002-m2 break appeared in hot-leg A l 5 Pressurizer back up heaters activated on low pressure 16 Reactor tripped on low pressure (thermal margin / low pressure); turbine tripped simultaneously; ADVs and TBVs " quick-opened" on turbine trip 34 SIAS on low pressure; maximum charging flow began; pressurizer heaters tripped off on low level 64 Operator tripped all RCPs; HPI flow began 502 SGIS on high containment pressure; TBVs isolated 529 ADVs closed on c4 reactor temperature 664 ADVs reopened on high reactor temperature 968 ADVs reclosed on low reactor temperature
~6500 Flow stagnated in Loop A 6636 Calculation terminated l aThese transients assumed multiple operator / equipment failure.
the primary system, while Fig. IX.A.4 summarizes the energy transfers into (positive) and out of (negative) the primary fluid. The first phase was characterized by a rapid depressurization of the primary that was halted by flashing in the upper head of the vessel at 110 s. During this phase of the accident the energetics were dominated by overcooling f by the SGs following the reactor trip. Heat rejection to the SGs decreased rapidly with the loss of forced convection following the RCP trip, however, and i 161
~ . .
500 -
, , , , 600
- 1. Rapid primary depressurization 0- 2. Approximate balance between ~
break flow and SIS flow S30-- . -500 2 Upper head S00- Volded - E F i T=
~
470 - w _ s NV y 440- )%
-300 y
~
q
~
)7 NOTE: These tr ansi en t s 3,,_ assumed multiple 5
_ operator / equipment -200 33o_ f ailures. See TABLE H.1. ..
- 32o_ _
-10 0 290- , , , , , , ,
0 1000 2000 3000 4000 5000 6000- 7000 8000 Time (s) Fig. IX.A.1. Downcomer temperature during 0.002-m2 hot-leg break from HP. 1s , , , , 2450 16 - - NOTE: These tr ans i en t s 2i00
'd -
assumed multiple operator / equipmenf i2 - f ailures. See TABLE H. -
~"
\
?
{ 10 - - m S ' .i400 e . R 5
~
L .
-1050 j 6-
-M0 4-
.-350 0 , , , , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Tima (s)
Fig. IX.A.2. Primary system pressure during 0.002-m2 hot-leg break from FP. 162
400000 , , , ,
-750000 300000-
-500000 200000-
\
un (tow _-2500
- l
,ooooo_-
5 charsi.
; flow __o l 9 o, ---**~~~*-"'*-~~~-*~~-~~~*
6-
- e '**'"'_'-~~~*'--'"'**-'*~~~~*--~~*\ut j -100000- _
--250000 $
-200000- _
--500000 Br ak
-300000-NOTE: These transients __,3oooo assumed multiple operator / equipment
-40#00-
_ f ailures. S::9 TABLE H. --iO00000
-500000 , , , , .
6000 7000 0 1000 2000 3000 4000 5000 Tme (s) Fig. IX.A.3. 2 Mass addition to primary system during 0.002-m hot-leg break from FP. 4# . , , , , , Decay heat
\
200- - 100 - HP1 flow Charging flow
? \ \ '
0 _ _= _- _- _- _- _- _- 8 _ioo.
-------....----------.........,sja _
Net l
-200- are'ak flow
~
~3 -
NOTE: These tr ansi en t s ossumed multiple
-4# - operator / equipment -
f ailures. See TABLE H.
-500- , , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 Trne (s) ! Fig. IX.A.4. i Summary of energy transfers that occurred during 0.002-m2 hot-leg break from FP. l l 163 l l
by the end of this phase of the accident, energy removal by the SGs was almost 90% completed, as shown by Fig. IX.A.4. The second phase (~110-6636 s) was characterized by the emergence of an approximate balance between the mass discharge rate from the hot-leg break and the SIS-injection rate, by a gradual decrease in primary pressure and temperature, and by extensive voiding in the upper plenum (see Figs. IX.A.5 - IX.A.9). At 502 s, SGIS was predicted to occur based on the l following calculation. The calculations are in English units because the primary data from the final safety-analysis reports (FSAR) are in English units. Because _ rl'e containment has a fixed volume, its pressure will increase when saturatel reactor coolant is discharged into it from the hot-leg break. A 6 simple integral method proposed by Darby can be used to predict the containment pressure excursion using TRAC histories of the enthalpy and flow rate of the discharged fluid. The method assumes thermal equilibrium between the containment atmosphere and steel structures in containment, and it includes heat conduction into the concrete. Darby's equations are: u, + Hsf - Ake (T-T o ) /t/wac " "acy, (T-Tr )
+ (ug(T) - ug(T)) + mfug(T) + matcvst(T-Tr ) (IX-1) 1 V
"o = mayc , (To-Tr) + matcvst (To-Tr ) + 8(To ) (IX-2) v (To) i p(T)=p,(T)f+p(T) o o
g (IX-3) i l l where the symbols are defined as: l a= thermal diffusivity i A= surface area of concrete exposed to the containment atmosphere cy = specific heat capacity at constant volume H= average fluid enthalpy k= thermal conductivity a= mass or integrated mass flow l 164 l
1.2 , i i i 1.0 -- ,-
~
~
NOTE: These transients l ! 5 assumed multiple ( operator / equipment 5 0. - .
@ f ailures. See TABLE H.
04-s ---- THETA CELL 1 0.2 -
- - - - - - - THETA CELL 2 .
THETA CELL 3
- THETA CELL 4 0.0 --
---- THETA CELL 5
-0.2 . . . . . i 0 1000 2000 3000 4000 S009 6000 7000 80C0 TIME (s)
Fig. IX.A.5. Voiding in the inner ring of the top level of the upper head during 0.002-m2 hot-leg break from FP. i.2 - , , . , , . . 1.0 - 0.8 - z NOTE: These transients assumed multiple S o 0.s - operator e pmen t - y f allures. ee BLE H. 0.4 - l s ! -- THETA CELL 1 0.2 _
- -- -- THETA CELL 2 .
THETA CELL 3
--- - THETA CELL 4
--- - THETA CELL 5 l
l
-02 , , , , .
0 1000 2000 3000 4000 5009 6000 7000 8000 l-TIME (s) Fig. IX.A.6. Voiding in the inner ring of the lower level of the upper t head during 0.002m 2hot-leg break from FP. l ! 165
L2 . . . . . . to- - 0.8 - - z NOTE: These transients O assumed multiple 0.6 ~
,. operator / equipment ~
-g
, ', j
,/ , f oilures. See TABLE H.
-h o4_ Y g
1 --
-- THETA CELL 1 0.2 - J. -------- - THETA CELL 2 _
l THETA CELL 3 0.0 -
--- THETA CELL 4 _
THETA CELL 5
-0.2 , , , . , ,
0 1000 2000 3000 4000 5000 6000 7000 8000 TIME (s) Fig. IX.A.7. Voiding in the outer ring of the lower level of the upper head during 0.002-m2 hot-leg break from FP. t2 , , , . , , ,
*~ NOTE: These transients ~
assumed multiple operator / equipment
- 0. _ f allures. See TABLE II. .
5_ 0.s - - Q E o-k 0.4 - -- 9 THETA CELL 1 0.2 _ ------- THETA CELL 2 - THETA CELL 3 I - THETA CELL 4 0.0 -
-- THETA CELL 5
-0.2 . , , . , ,
0 1000 2000 3000 4000 5000 6000 7000 8000 TIME (s) Fig. IX.A.8. Voiding in the upper level of the upper plenum during 0.002-m2 hot-leg break f rom FP. 166
., - . , . . . , _ -_,,m--..m,,--, , , - - - - , .
p= prsssure t= time T= temperature u= specific internal energy l v= specific volume V= containment volume L cubscripts a, air c, concrete g,1, saturated vapor, liquid o, initial condition r, reference st, steel 11 and af are obtained from TRAC. The following data was obtained from Chapter 14.20 of the FSAR6 A ~ 1.2 x 105 ft2 ke ~ 0.8 hr-ft OF "c
~ 0.03 ft 2/h V ~ 2.0 x 106 ft3 m,e ~ 2.0 x 106 lb m, ~ 1.4 x 105 lb
( c,y ~ 0.171 l lb0F c BTU vst ~ 0.12 lb0 F T,. = 1200F I relative humidity = 50% pa(To ) = 13.85 psia. When this data is used in Eq. (IX-2) with rT = 320 F, it yields i l 167
1 l
> . .
O S00 1000 1500 2000 2500 3000 3500 4000 4500 Time (s) Fig. IX.C.3. Mass addition to primary system during a stuck-open PORV from HZP. 50 . . . . . . . , core 25- Tab
/FI chargina_
0- --r : -WbI - pump
\
SGB 7 - net
.n. -
E .too. . l t -es- - break j _so . .
.ns- -
- -200- . . , , . ,
l 0 SOO 1000 200 2000 2500 3000 JS00 4000 4500 , T!r.le (s) I Fig. IX.C.4. I Summary of energy trane.fers dur4v,t ituck-open PORV from i HZP. 189
#00 . . . . , , ,
s00- -
. 2o0a t# P n (**"'IO 000- . -
-150
.......... Loo,egeo.,,,
700-- .
. son i
e00- . l uso s00- -
. ,ono k h -
E a00- . a 300-75o g l 500 2o0 _ ( 100-
~
. -250 0- .o
.t ........................ .
0 1000 2000 3000 4000 5000 6000 7000 8000 he (s) Fig. IX.C.5. Hot-leg flows during a stuck-open PORV from HZP. 20 . . 45n - -1000 Lonr
. .. . . .. .. LDO, A1A2(Consp 2'.)
(COUD d'-) 400 war mtenu ra
- Loop 92 (COMP JQ 800 l
300 9 . -r00 d 50 t e, es a z _ NO
- o Q, . -ano rt i
g so. .. . g 2 i2 - , NO *
.[.' .
I - E,U [' 200
- n. . . ,,.
sn i ', ' .,. .I I ,.
- I,. '..
0 'D , l Y ' l\ 0
-s0 .,
-200
_,00 - o soou 2000 s000 <uou u000 s000 7000 vuo0 he (s) Fig. IX.C.6. Cold-leg flows during a stuck-open PORV from HZP. 190 1
soo , , , , , , . -e00 560- - 7s.
" .'i .a. -
530-
- /: ji too, ,,
--s00
- 1, , ' .,. j ,, .......... Loor A2 g
500-l' i i i l*/ '.. i::/
/ '. .
's toor et
- toor s2 -
P
..t -400 44-o F
fg h i)! f , l 440- ' . ,g I fya 1 -300 4t0 - i q
\ -
., s,.~..
~-
380 s .
's 200 350- l, 320- --
. .W0 200 , , , , , ,
0 1000 2000 3000 4000 S000 6000 7000 6000 he (s) Fig. IX.C.7. Liquid temperature upstream of HPI/ charging flow ports during a stuck-open PORV from HZP. 20 , , , , , , , 1s -
- -40
**i rtow pro coin un 16- .
.......... CHA=GWG FLOW PUI COLD MG (LeoP5 A2 20 90 --35 1d -
. 3o s,
12 -
-1s kS, 8
- l Q 10- ~
g
- -20 ,
g a s- - j
.e 6- -
. .io 4, .
2- j - ~5 0-
' . , , . , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 he (s)
Fig. IX.C.8. HPI/ charging flow during a stuck-open PORV from HZP. 191
4 pressurizer became saturated and a bubble formed, which slowed the depressurization. The HPI flow, a function of pressure, increased and brought
.the pressure back to ~7.2 MPa (1058 psia). In the extrapolated region (4135-7200 s) the pressure is expected to level at ~6.8 MPa (1000 psia) as in the previous transient.
j In conclusion, this transient calculation demonstrated that flow l 1 stagnation in both loops is possible during an SBLOCA if the transient is , initiated from a low decay-heat level. D. Comparisons Figures IX.D.1-IX.D.3 compare the downconer liquid temperature and energy transfers for the SBLOCAs. The primary difference between the two SBLOCAs initialized from FP was the early loop stagnation caused by the stuck-open ADV in the second transient. Although the hot-leg break was approximately twice as large as the stuck-open PORV, the characteristics and timing of the primary-system depressurization were essentially the same in both cases. Of course, the quasi-steady pressure reached in the second transient was
- - significantly higher than that reached in the first because of the smaller flow through the PORV. However, except for this effect, the more extensive voiding in the hot-leg-break LOCA and the early loop stagnation in the stuck-open-PORV i LOCA, the accident signatures of the two cases were remarkably similar. The effect of the stuck-open ADV was twofold. First, secondary cooling was not
' terminated by SGIS because the ADVs are upstream of the MSIVs. Second, following SGIS, reverse heat transfer in the isolated SG together with enhanced heat transfer in the affected SC induced stagnation in the loop with the isolated SG. If the ADV had fuc 4..oned properly, it is expected that early loop
. stagnation would not have occurr. y,d voiding in the upper plenum of the vessel probably would have been more extens. re (because of reduced secondary cooling). The primary differences betw(en the two SBLOCA calculations previously described and the stuck-open PORV f t >m HZP were, of course, the initial conditions and the resulting phenomena. Because flow stagnation occurred in both loops in the SBLOCA initiated from HZP, the temperature in the downcomer decreased considerably faster because of the " direct" HPI injection. Also J because of the flow stagnation, heat input from the SGs was lower. The time for comparing the SG heat input in Fig. IX.D.3 is after 500 s because the initial difference is a result of the reactor trip from FP and the initial heat load removed by the SGs. 192
,. - .. .,..-.,_~....,.-.---_m--.,,~,m.. _- ,..m -- ,
575 , , , , , , ,
-560 Hot-leg break from FP 550-
'.. -520
--------- Stuck-open PORV/ADV from FP s
l- 525- . N..,, - Stuck-open PORV from HZP -
-480 i ',
- 8. 500- - I i
\ .,'
440 g O Is
' q[.,\ ,..
2
-400 '
2 c5 i - E. s i'I?
'. f, o-a E
n 9
~.. . . . , _
-0 ;
k, .-..' ..., -320 425- '...,, - 1 s l -280 400- ' - 240 375 , , , , , , , 0 1000 2000 3000 4000 5000 6000 7000 8000 Trne (s) Fig. IX.D.1. Comparison of the downcomer liquid temperature for two SBLOCA Cases. 10 0 , , , 0- ,
, snJcx oren Ponv + sitscr ortu Aov f rom FP '
---------- wor-go onta= from FP
\,,'..,'N " " ~ ' " " "
-10 0 - ------q,------~ sA+B
~
\
9 s, , SG p + B l 200- '. / - f v '. HOT-LEG x B AK
. B
-300- . (PORV)
.s.~.. '.
-400- '..,,'
NOTE: These transients '. assumed multiple J .. ""
-500- operator / equipment un f allures. See TABLE H.1.
- -600 0 1000 2000 3$00 4000 Lod 0 (,dOJ 7000 b.,0 Trne (s)
Fig. IX.D.2. Comparison of the energy transfers for two SBLOCA cases. 193
50 , , , , , SG,A+B 0- -
.50 '.2'.,....... .........,,,,,,,,
nt SG +B .
'............'.'.q....----""*""".....w--,,....l.....---
- 10 0 - .
-rA .
..) . .
9 - treak I N g -200- , x , an . D - 250- '. - - d ',
-300- *
',, break
-3# Stuck-open PORV from HZP -
~
0.02-a2hot-leg break from FP ',
-400- -
~400- -
-500- , , , , , .
0 1000 2000 3000 4000 5000 6000 7000 Time (s) Fig. IX.D.3. Comparison of the energy transfer for the hot-leg break from FP and the stuck-open PORV from HZP. r f I I 194
i l X. UNCERTAINTIES 4 ORNL has asked Los Alamos to estimate the uncertainties associated with the results presented in this thermal-hydraulic study. Oar interpretation of this request is as follows: given the nominal initial conditions of the plant, dafine an error band on three thermal-hydraulic parameters predicted by TRAC such that the actual plant parameters observed during a given transient I (assuming identical nominal initial conditions) would fall within that error ,
- band. These parameters are downconer-liquid temperature, primary-system pressure, and the downconer heat-transfer coefficient on the vessel wall.
To set realistic error bands on each of the transients presented in this otudy, we would have to run each transient numerous times to assess the effect upon the results of changing plant parameters, physical models within TRAC,
-ostpoints, etc., that have known uncertainties. This would yield a spectrum of answers from which the uncertainty of the TRAC results could be calculated. For example, at FP conditions, the mass in each SG is known to within i 10% (3s).
If a steamline break should occur when the mass in the broken SG was 10% higher than its average mass at that initial condition, the minimum downcomer 0 I tcaperature could be 20 to 50 K (36 to 90 F) colder when compared to the case of a break occurring when the mass of the broken SG was at its average value. Hince, because of these fluctuations in parameters that potentially affect the thermal-hydraulic response of the plant during any transient, the results
~
predicted by TRAC may show significant deviation from the actual plant response. In addition to the unavoidable uncertainties that arise because of i fluctuating initial conditions that occur naturally in the plant, the TRAC model can also introduce uncertainties as a result of uncertainties associated with the physical models (that is, heat-transfer correlations, choked-flow model, etc.) used within TRAC in predicting the thermal-hydraulic response. For example, during a steamline break or a small-break LOCA, a small change in the mass flow rate out the break can significantly alter the thermal-hydraulic response of the primary-side fluid temperature. However, it is known that the choked-flow model used in TRAC can predict the mass flow rate to within only 10%.- Again, determining the overall effect produced by the uncertainties associated with the physical models would require running the same transient numerous times to bracket the spectrum of answers that would result over the range of uncertainties of each of the physical models. 195
. _ _ _ , _ . _ ~ . _ . _ _ . _ _ _ , - . . _ _ . , _ _ _ _ _ . _
Because of the nonlinear behavior of the system, it is unrealistic to presume an analysis of this type (that is, the " numerous-run"' method) would result in an answer applicable to all remaining transients. It can be postulated that a small uncertainty associated with one of the physical models may produce a negligible change from the best-estimate answer for one transient, but the same change in another transient may cause the transient to take a different path, and the subsequent disparity between the TRAC prediction and the ! actual plant response may become significant. Therefore, the " numerous-run" i method would have to be applied to each transient in the study to assess its overall uncertainties. Such analyses would require time, manpower, and money, none of which was available in the time remaining to complete this study. In the absence of these resources, we estimated the uncertainty in the time-dependent downcomer temperature that arises from the uncertainty of some of the n.ajor parameters that may alter the thermal-hydraulic response of the system for each transient. A satisfactory method that estimates the uncertainty in the primary-side pressure has not yet been devised, and therefore can only be guessed at this point. The uncertainty in the downcomer-heat-transfer coefficient can be resolved directly using the information predicted by TRAC and will be discussed later in this section. The basis for this approximate uncertainty analysis for the downcomer temperature resides in the following assumptions:
- 1. over the range of the uncertainties, the transient progresses along the same path such that additional operator and/or equipment action (such as valves opening or closing, pumps tripping or failing to trip, loops stagnating, etc.) would not occur; j 2. each of the major energy transfers that occurred during a transient can l be treated as being independent of the remaining energy-transfer terms l for a short period of time, and therefore can be perturbed by a small amount without significantly altering the remaining energy transfers; I and j 3. the downcomer-temperature changes are in direct proportion to changes in the bulk temperature.
l Under these assumptions, if any of the energy transfers is subject to 1 I uncertainty, then those time-dependent energy transfers predicted by TRAC can be perturbed and a new internal energy can be re-calculated by adding the perturbed t energy transfers to the unperturbed energy transfers. Using the new 196 1
time-dependent internal energy and the original time-dependent internal energy, the AT produced by this perturbation can be determined by AT = U' - U Mc p where U' = new time-dependent internal energy caused by the perturbation, U = original time-dependent internal energy calculated by TRAC, M = time-dependent liquid mass of primary system, and cp = average value of specific heat of the primary liquid. Adding the AT to the time-dependent downcomer temperature predicted by TRAC yields an estimate of the uncertainty in the TRAC results arising from the uncertainties in some of the key parameters affecting the thermal-hydraulic response of that transient. It must be stressed that this method does not account for the coupling between the various energy t.ransfers. For example, consider the case of a
~
steamline break. If a p;-turbation is made to the broken SG energy transfer such that the primary-system-liquid temperature will be lower, then the energy transferred from the primary-system metal to the primary fluid should increase. This effect would tend to slow the temperature decrease of the primary fluid. l In addition, the primary-side pressure should decrease relative to the original pressure history that was predicted by TRAC if the primary-side liquid temperature decreased relative to the original temperature history. This would e induce more HPI flow, which would, in turn, produce an additional primary-side I temperature decrease. If the RCPs tripped during the original transient based on low primary-side pressure, then it is anticipated that they would trip carlier under the perturbed conditions. The energy transferred from the intact l SG back into the primary fluid would be terminated earlier because of the loss of. forced convection in the intact SG, and the intact loop would stagnate earlier. This would result in an earlier loss of a heat source to the primary fluid and subsequently lead to lower downcomer temperatures. ! Because of the inherent coupling between the energy transfers, each perturbation must affect the behavior of the remaining components of the primary cystem with some characteristic coupling time constant. If the time constants for coupling are long in coaparison to the time interval for the energy 197
m; 1 transfers that cause the original cooldown, then for that period, it is reasonable to assume that the energy-transfers are independent of each other. . LAs an example, if the energy removed via a broken SG were to be increased, the , l reverse heat transfer occurring in the intact SG would not be expected to be affected' significantly during the initial portion of the transient. However, during the latter portion of the. transient, the time-dependent energy load of the l intact SG is expected to adjust in accordance to the perturbed primary-side fluid temperature. -In conclusion, a small perturbation at the beginning of a transient. may produce a valid response, but towards the end of_the transient, the AT predicted via this method will become unrealistic because of the long-term coupling between the energy transfers into or out of the primary system. For this reason, the majority of the uncertainty analysis used in this study is confined to time intervals that do not extend far beyond the time of minimum downcomer temperature.- Note that the intent of this uncertainty analysis is to quantify the effect upon the downcomer temperature history produced by known uncertainties in the key parameters in the thermal-hydraulic solution. This approach, in our opinion, seems best to satisfy the request of ORNL and is consistent with our interpretation of that request. However, this analysis does not consider every possible effect and therefore should be viewed judiciously. A. Uncertainty Sources Considered Upon review of the transients analyzed in this study, there are several fluctuating plant parameters and uncertainties associated with physical models used within TRAC that can significantly alter the results for the time-dependent downcomer temperature. These are:
- 1. the initial mass in the SGs prior to the initiation of the transients,
- 2. the choked-flow model used in TRAC,
- 3. the amount of decay heat present in the reactor core prior to the initation of the transients,
- 4. the primary-side pressure history in terms of its effect upon the time of the RCPs trip and the total HPI liquid that is injected,
- 5. the feedwater temperature at the SG inlet, and
- 6. the condenser /hotwell liquid inventory for transients in which SGIS does not occur.
198
An error analysis was performed for each of the 13 transients using the uncertainties associated with these key parameters. Two sets of perturbations wire made for the majority of the transients. One set was made to produce a nigative AT so that a lower error band was formed. The other set was made to produce a positive AT so than an upper error band was formed. A discussion of the uncertainty analysis for each of the transients follows in the next esctions. The basis for the selection of the six key parameters previously discussed is as follows.
- 1. SG Mass. The nominal value of the total mass in each SG at FP ccaditions was transmitted to Los Alamos as 63000 kg i 10% (138000 lb i 10%).
The error was presumed by Los Alamos to represent 3a. In addition, we presumed that the error in the SG mass corresponded to the situation in which the feedwater train was operating in automatic, and the regulating valves on the fsedwater train were modulating to control SG level to the normal liquid level (2.95 m (116 in.) above the lower tap of the narrow-range level instrument). Under conditions in. which the operator controls SG 1evel manually (~5% of the time), the-level and corresponding mass may fluctuate somewhat more. With the level control of each SG in automatic at HZP conditions, the uncertainty associated with the mass of each SG is ~2-3% (3o). However, because of changes in the feedwater temperature relative to what was chosen in the TRAC model (80 0F) for this study, together with the possibility of the operator manually controlling level at HZP, we felt that it could be as high as 10% (3a) et this initial condition. Therefore, it is assumed that the mass of each SG at HZP conditions is 104000 kg i ~10% (229000 lb i 10%) (3a).
- 2. Choked-Flow Model. As previously mentioned, the choked-flow model currently used in TRAC has been demonstrated to be able to predict the mass flow rate out of breaks to within 10% (3o). In comparison to other commonly accepted choked-flow models (see Fig. X.1), the TRAC solution appears to yield a "ciddle-of-the-road" answer..
- 3. Decay Heat Following Shutdown. During the formulation period of this otudy, Los Alamos assumed that the decay heat for the FP cases was to be calculated assuming that the reactor had been operating for an infinite length df time prior to the initiation of the transient. The error (or bias) introduced by this assumption is on the order of a 5-6% excess decay heat everaged over the two-hour transient-analysis time. This was determined by 199
performing _ a decay-heat calculation using the current ANSI /ANS-5.1-1979 decay-heat correlation. The calculation was performed using three batches of fuel. It was assumed that the reactor was refueled once a year and that one-third of the core was replaced per year. Under these conditions, one-third of_the core resides in the reactor for three years, one-third for two years, and one-third for only one year. A 60-day refueling interval per year was included j in the analysis. A comparison of the decay-heat history predicted by TRAC for , this study (in error because of the assumption of an infinite operating history) as opposed to the decay-heat history predicted by the ANSI correlation for an infinite operating history and the three year operating history is presented in Fig. X.2. The total energy produced by the three cases is compared in Fig. X.3. According to ANSI /ANS-5.1-1979, the decay heat is accurate to within 6-7%
- (3o). Because the decay-heat energy addition to the primary fluid should have been 94% of what TRAC calculated, the 3o-lower error band would represent ~ 87%
of that calculated by TRAC. The upper band would be 100% of that calculated by
- TRAC. Therefore, for the majority of the uncertainty analysis performed in this
, study for cases initiated from FP, the decay-heat energy curve was perturbed by multiplying by 0.87 in the estimate of the lower error band and unchanged in the estimate of the upper error band. According to transient specifications, the HZP cases were to be initiated l 100 hours after shutdown. TRAC predicted a decay heat power of ~9.4 MW, whereas the ANSI correlation for the three year operating history case predicted
~7.7 MW. Running the HZP cases from a decay-heat power level of 9.4 MW, however, can be interpreted as merely initiating the transient earlier than 100 l hours after shutdown. To judge the sensitivity of the time after shutdown that the transient is initiated, the lower error band for each HZP case was formulated by reducing the decay heat by a factor of two, and the upper error band was formulated by doubling the decay heat. As shown in Fig. X.4, in the limit as the decay heat approaches zero, the impact on the downcomer temperature appears to be minimal. However, this is not truly the case. In the limit as the decay heat approaches zero, both loops can stagnate. This may cause the time-dependent downcomer-temperature history to follow a different path and l
invalidate the error analysis (see assumption 1 in the previous discussion of I error-analysis assumptions). This will be explained more fully in the conclusion section of this report (Sec. XI). 200
Moooo . . . .
- (se N d) 2eleudek -270o0 (doch) surneN 12oo00 Id'I) "" '
-(chn-deh) Herry Feuche .g (chn-det) Noody , , .
p , Ts,*MO K. P os e*71 MP* * * , . . . -'[,,// . ziooo O s E 'no"
;, // -
k
. . ./.. **/. * '-/'/'
\
- ' * * ^
i w i
, 0000o- g 3 -mooo
.'ff . . ~ . . '/_/'/
w ** 8 cocoo- .
/.- '-nooo y
./ / (seum) m e-pri
/ ~
-0000 4o000--
p/ 0000 20000 . . > > 6 8 10 12 14 16 Stagnation Pressure (MPa) . Fig. X.1. Comparison of TRAC's choked-flow model to other models. 10 0 , , , , . 4 6 i TRAC-PFI ' 90- 1
......... 79ANS-3 YR 80- ) - 7SANS-INRNITE -
)
70-T ',
%e 60-y ,
c2 \, so. ', -
,'. s 40-30- " " - -- ..........,,,,,,,
20 . . . > > > 0 1000 2000 3000 4000 SOOO 6000 7000 8000 Tkne (s) Fig. X.2. Comparison of decay heat power. 201
350 , , , , , , , 300-
, x'.. .... -
g,...
- g. . -
250- - 200-
- l. '
15 0 . di ,..-
' ~
10 0 - ..' -
, .......... /9ANS-3 YR 50 - 79ANS-INRNITE _
0 , , , , , , , 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. X.3. Comparison of total decay-heat energy. 575 , , , , , , , 40 550- -
~#
525 - -
- 450 8 50 ~
9.4 MW C
-- - -zero ,o3 $
475 -
~
- 0 450- -
N S g 425- .., - 400-v .,,,
--[ ......_.....--
- 270
-225 33_ .
- 18 0 350 , , , , , , ,
0 250 500 750 1000 1,'50 COO (75J **
. 0 'JO Tme (s)
Fig. X.4. Downcomer temperature history comparison for two different decay heat levels - 1) 9.4 MW, and 2) zero. 202
- 4. Pressure-History Effect. As previously stated, the change in the primary-side pressure cannot be predicted at this time. However, the consequences of changing the primary-side pressure were observed for one of the transients (0.1-m2 (1-ft 2) MSLB from FP). During one of the evolutionary changes of the TRAC model, the pressurizer was inadvertently filled with high'ly subcooled liquid. This resulted in an extremely rapid pressure decrease in comparison to the same case run at a later time using a corrected model in which the pressurizer was filled with saturated liquid (see Fig. X.5). For the case in which the pressurizer liquid temperature was corrected, the RCPs tripped 16.7 s later than in the original case. The additional pump power added to the primary fluid was ~0.3 GW-s of energy. Accordingly this should only increase the primary fluid temperature by ~0.2 K (0.36 F). % wever, as seen in Fig. X.6, ,
0 the minimum downcomer temperature was ~25 K (45 F) higher. This was the direct result of three effects associated with the RCPs runn' c,. The major effect was 7 the reverse-heat transfer that was allowed to contic > An the intact SG while j the RCPs . operated for the additional 16.7 s. essence, while the RCPs are l running, the secondary-side liquid in the intac, L will remain essentially in thermal equilibrium with the primary-side-liquid temperature. This increases , the effective total heat capacity of the primary-side liquid by as much as
- 20-30%, which produces a smaller primary-side-temperature decrease per unit of energy removed through the broken SG.
Closely related to this effect is the effect of loop stagnation that can occur if the RCPs are tripped. Loop stagnation will result in the loss of j approximately one-third of the heat capacity of the primary-side liquid. This, j in turn, will produce a larger primary-side temperature decrease per unit of snergy removed through the broken SG. Hence, the effective heat capacity of the primary-side fluid can change by a factor of ~2 for a change in the situation in which the RCPs are running (keeping the intact SG in thermal equilibrium with the primary fluid) to a situation in which the intact loop stagnates. As observed in the 0.1-m2 (1-ft 2) steamline break case that was rerun, the intact
- loop stagnated for approximately 600 s in the original (and erroneous) model.
(An additional input error was present in the original model that acted in ceries with the premature tripping of the RCPs. The error produced an artificial gravity head that opposed the gravity head created in the vessel which was sufficient to help induce and prolong loop stagnation.) When the model i 203 I 4
---,--.+,.,,,v,,e.n,..-.-n.,,-,,.,---m,, .,,.-m,,-. .,.-,~.,-------n .n...e--,n+--. ..--.,,--x- +-,,,-.m...n .,,w ,....-.--.-e-,. - - - , - . - .
is , , , , , , , ,
-ame
-ame
- a. .
..=
12 ..,se 7 ....., is. ..........,ee. ill J ,-wee a- =
' -nee
- a. ..
6- - m
- 4. .
~ ~
2- - 0.0 . . . . . . . . . e e ase see no asos esse ese use asse asse asse MW Fig. X.S. Comparison of pressure history for 0.1--m 2 steamline break for an n.rroneous TRAC model and a corrected TRAC model. 600 , , , , , , , , 575 cmGpeAL
' - - - -- RLRug -540
$ 525
~*,,...-
"~~~~O"-- ' '
b e Q ,.-....... 450 a 13-500- , . ./
~
E o h ' 405 h ! ,g 475 ** 1 *y .2 HOTE: These transients 3s0 ?y 450 assumed multiple 3 i operator equipmen t l f ailures. ee TABLE H. . sts 425
--270 400- -
375 . . , . , , . . , O 250 500 75J 1000 u50 1500 1750 2000 2250 2500 Time (s) Fig. X.6. Comparison of downcomer-liquid-temperature history for 0.1-m2 steamline break for an erroneous TRAC model and a corrected TRAC model.
4
? was corrected,-the intact SG loop did not stagnate, but flow did become very low I for a few seconds shortly after the RCPs tripped.
A third and relatively minor effect that helped minimize the temperature decrease observed.in the rerun of this particular transient was the reduced I amount of HPI liquid that was injected into the primary system accompanying the i relatively higher pressure history. This, however, accounted for only a few degrees change in the downconer temperature. The combination of all three effects (that is, higher pressure leading to the tripping of the RCPs 16.7 s later, no loop stagnation, and reduced HPI flow corresponding to the higher pressure history) led to a 25 K (45'F) higher downcomer temperature at the time of minimum downconer temperature. Assuming
, that the changes observed during this transient are representative of the i sensitivity of the TRAC results corresponding to relatively small changes in the pressure history, it can be concluded that if the pressure history is perturbed slightly, significant changes in the downcomer temperature can be observed.
However, as previously stated, we are unable to predict the changes in the i i pressure history that would occur if the primary-side liquid temperature. were perturbed. Hence, because the uncertainty-analysis method perturbs the primary-l l side liquid temperature but not the pressure, our confidence that the upper and
. lower error bands derived will bound 100% of all possible results obtained over the range of the known uncertainties is diminished. In spite of this
( limitation, the uncertainty analysis method is still beneficial in terms of establishing the order of magnitude of the change of the downcomer temperature !' that might occur as a result of known uncertainties. That is, for example, does I a 10% increase in the SG mass produce a 0.2 K, 2 K, 20 K, or 200 K decrease in the downcomer temperature during a steamline break?
- 5. Feedwater Temperature. Under steamline-break conditions, the initial mass in the broken-loop SG determines the amount of energy that can be removed from the primary system during the cooldown. However, under runaway-feedwater conditions, the temperature history of the feedwater entering the SG plus the duration of the excess feedwater determine the amount of energy that can be removed from the primary system. For example, if the temperature of the feedwater were initially lower than the nominal value used in the TRAC model, the temperature decay of the secondary-side heat-sink temperature would be more pronounced during the runaway-feedwater portion of the transient. This in turn 205 l
- - - + , , ,,--m-, rnmn ..,--.vn, ,. m n,an.wm-ny,,,,,e ,. -e-,.m.,m. ww,,.w,,n-- -,
would remove more energy from the primary system and produce a larger . decrease in downcomer temperature.
- 6. Condenser /Hotwell Liquid Inventory. In ' addition, because the )
runaway-feedwater portion of the transient is terminated by the tripping of the l l feedwater and condensate pumps, the point in time at which the pumps trip would 4 offect the energy-removal capability of a runaway-feedwater transient. Because the pumps trip on low suction pressure created by loss of condenser /hotwell liquid inventory, the initial mass in the condenser /hotwell is of prime importance. Although the fluctuations of the condenser /hotwell liquid inventory 2 were never formally quantified during the PTS-working group meetings, the combination of a fluctuating feedwater-inlet temperature and a fluctuating condenser /hotwell liquid inventory that occurs during normal plant operation could conceivably lead to a change in the energy removal on the order of 10% in comparison to the TRAC prediction for this transient. B. Runaway Feedwater Three runaway-feedwater transients were analyzed in this study. The , cnergy curves for each of these transients were presented in Sec. VII and will not be repeated. 1 The following uncertainty analyses were performed ast -tng that the energy , curves for the pumps and the slabs in each of the transients were unperturbed. Depending upon the individual transient, the energy curves for the SGs and the reactor decay heat were perturbed but for reasons that are different than those rationalized in the steamline-break analyses. For the runaway feedwater transients, it was assumed that the energy curve for the appropriate SG could be perturbed by i 10% as the result of uncertainties associated with tFr SG mass, feedwater temperature, and initial condenser /hotwell liquid inventory. As previously discussed, the uncertainty analysis is usually not valid for long periods of time into the transients. Hence, it must be presumed that the I
- uncertainty error bands derived during the initial portion of each transient
,. t:ill be representative of the error bands that would occur at later times in the ; transients. This assumption, however, is contingent upon the primary-side fluid temperature not converging on an equilibrium temperature controlled by the cecondary-side-heat-sink temperature and a pre-established decay-heat power. In this situation, the errors should be significantly smaller than those predicted by this method. { t 206 1
- 1. Runaway MFW to Two SGs from FP.
The energy curves for both SGs were multiplied by 1.1, and the decay-heat energy curve was multiplied by .87 in the assessment of the lower error band. The upper error band was obtained by multiplying the SGs energy curves by 0.9, and leaving the decay heat curve unaltered. Figure X.7 shows an estimate of the uncertainty of the downcomer temperature for this transient.
- 2. Runaway MFW to One SG from FP. Figure X.8 shows an estimate of the uncertainty of the downcomer temperature for this transient. The uncertainty analysis was based on the same type of argument presented in the previous transient discussion with one exception. Only the energy curve for the one SG receiving the. excess feedwater (SG A) was multiplied by 1.1 and 0.9 to obtain the lower and upper error band, respectively.
- 3. Runaway AFW to Two SGs from FP. During the loss-of-feedwater (LOFW) portion of the transient (the first 1200 s), the temperature on the secondary-side is fixed at the saturation temperature corresponding to the pressure of the SGs. This produces a constant temperature on the primary side prior to the SGs drying out. The uncertainty of the primary-side fluid temperature, therefore, corresponds to the uncertainty of the overall heat-transfer coefficient between the primary side and the secondary side in the SGs during this interim.
After the SGs dry out, the primary-side fluid temperature begins to increase at a rate dictated by the decay-heat power and the total mass in the primary system. At 1200 s, however, the AFW system is manually activated (according to transient specifications). As in the previous cases, it was assumed that the energy curves for both SGs were increased or decreased by 10% in the assessment of the error bands. I This is also applicable in this transient if it can be speculated that the mass flow rate of the AFW system set by the operator is within 10% of the value specified by ORNL. Under conditions of a different AFW mass flow rate, the time-dependent temperature decay produced in each SG would be different following the initiation of AFW. This would produce a different temperature i history for the primary-side fluid temperature. A crude approximation of the secondary-side heat-sink temperature following the initiation of AFW to a dry SG l can be obtained from 207
7.,... r. 600 , , , , , ,
)
500 l SM ~ NOTE: These h ansien t s ossumed m silple 560 l
, operator equipmen t 55c- - '. f c H ur es. ee TABLE L
's. *.s -520 b 525-
','., b
'.,, ' , , ,,,,,,,,,,,. 4gg g
'\ '
\ ., , . . .
500-- *
.+ -440 F
475 'g ..* 400
-a 3 360 450
-320 425 280 400 , , , , , , , , ,
0 90 0 200 300 400 f,00 000 730 000 S0J 000 Time (:) Fig. X.7. Downcomer-temperature uncertainty for runaway-main-feedwater to two SGs from FP. 600 , , , , , , , G00 575 560 SM k*~ *' b 8 j .
\, , , . . .
480
? .,. '
d 500
, 7 440 L:
',\ m /-
p c
*' ;U
,3 400 0
450 NOTE: These transients ossumed mulilple 425 operator / equipment 320 f o Hures. See TABLE I. 280 400 , , , , , , , 0 10 3 200 MG 400 %J C00 do 650 950 toda Time (s) Fig. X.8. Downcomer-temperature uncertainty for runaway-main-feedwater 208 to one SG from FP.
c,.- --
,3 .; a g[Qdt ac pt d
where T, = secondary-side heat-sink temperature,
~
r, = auxiliary feedwater inlet temperature, , Q = heat load to each SG, a = AFW asas flow rate, ; ep = specific heat capacity of secondary-side fluid, and - t = time after initiation of AFW. j. For this particular transient, the energy curves for both SGs were perturbed by 10% at times greater than 1200 s. The decay-heat energy curve was cleo~ perturbed in the usual manner but at times greater than 1200 s. Figure X.9 chows an estimate of. the uncertainty of the downconer temperature for this i transient. C. Steamline Breaks o Seven stesaline break transients were analyzed in this study. The energy curves for each of these transients were presented in Sec. VIII and will not be ; i repeated. The following uncertainty analyses were performed assuming that the energy curves for the pumps and the slabs in each of the transients were unperturbed. Depending upon the individual transie.nt, the energy curves for the SGs and the reactor decay heat were perturbed in accordance to the previously discussed uncertainties for those terms. As previously discussed, the uncertainty analysis is usually not valid for long periods of time into the transients. Hence, it must be presumed that the f uncertainty error bands derived during the initial portion of each transient ! 1 f will be representative of the error bands that would occur at later times in the transients. This presumption, however, is contingent upon the primary-side i fluid temperature not approaching an equilibrium temperature controlled by the secondary-side heat-sink temperature and the decay-heat power. In this situation, the errors should be significantly smaller than those predicted by this method. ~
- 1. 0.1-a Steauline Break _ftca HZP. Figure X.10 shows an estimate of the uncertainty of the downconer temperature for this transient. The lower error l band was produced by multiplying the decay-heat energy curve by 0.64 and the 209 1
600 , , . > i . 600 575 I.y G '
'.,, ~-'+ ........ 500 525 !
g '..'*..,,'*.. 's' '* , , . -,
,. C 500
-450 g 3
4 5 g ,,, -- . 400 , t j \v.(.. . ,* 2 y 450 ...- 350 .
.t~ .2
~
425 300 NOTE: These transients 400 assumed multiple 250 operator / equipment 375 f ailures. See T ABLE I. -
-200 350 , .
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. X.9. Downcomer-temperature uncertainty for runaway-auxiliary-feedwater to two SGs from FP. 53 , , , , , - 540 550-495 525 . NOTE: These fransients assumed multiple
. operator t _ 450 E 500 t o iiu, .s. eee .I E g
,a, 405 .$.
2 475 .
\. .
u0 l 450 ' ' .
>E
\ ' .. ,
', . 1
,3
.g.
_, 425
. . . . . . . , , , . .35. . . _.c .--
i s00
' .., W -[^ 270 i,,...* ..
.................,,.....+
225 3/5 wo 350 0 250 550 750 No 12A do S 2000 Time (s) Fig. X.10. Downcomer-temperature uncertainty for 0.1-m2 steamline break 210 from E P. L
broken SG energy curve (SG A) by 1.1. Multiplying the SG-A energy curve by 1.1 i corresponds to the assumption that the mass in the SG was 10% higher than its I mean value at the time of the transient and that the mass flux out the break was 10% higher than what TRAC predicted. The energy curves for the pumps, the slabs, and SG B were unaltered. Because Loop B stagnated early and remained stagnant for ~3000 s, the assumption that its energy history would remain , unchanged is well founded. l The upper error band was estimated by doubling the decay-heat energy curve and multiplying the energy history curve for SG A by 0.9. This corresponds to initiating the transient from a condition in which the SG-A mass is 10% lower than its average mass and that the mass flux out the break is 10% lower than j what TRAC predicted. , i Because of the coupling effect between the decay heat and the energy ! i j dissipated by each SG later in the transient, the error analysis is presumed to j become invalid at times greater than ~2000 s. The errors established during the i j first 2000 s, however, are presumed to be characteristic of the errors that would be present at later times in the transient. [
- 2. 0.1-s2 Steaaline Break from FP. Figure X.11 shows an estimate of the uncertainty of the downconer temperature for this transient. Because this l transient was initiated from FP, the decay-heat energy curve was multiplied by 0.87 in the assessment of the lower error band. The decay-heat energy curve was unaltered in the assessment of the upper error band.
i As in the previous steamline break uncertainty analysis, the energy curve for the broken SG was multiplied by 1.1 to obtain the lower error band and was l- multiplied by 0.9 to obtain the upper error band. Although the intact loop did not stagnate, the amount of reverse heat transfer occurring in the intact SG was . I j presumed not to change significantly during the initial portion of the l l transient, and therefore its corresponding energy history was unaltered. The cnergy histories for the pumps and slabs were also assumed not to change during the initial portion of this transient. l 3. 0.1-m2 Steam 1_ine Break with Two operatina RCPs from HZP_. Figure X.12 i t chows an estimate of the uncertainty of the downconer temperature for this l l transient. The uncertainty analysis was performed using the same multiplication ! factors discussed in Sec. X.B.1. i l l l 211 l r 1 i
e , , I , , s
#0 550 i
l
. . . . . - 495 i 525 . . . - !
\ ., ,
~
$ $00 ,
Q g \, .... e 1 475 5. \. , *' *,
' #UU
{ i .* 360 k
) 450 .,,*
. g 3' 3'
,{- 425 Iv0TE: These transients 3t5 't-
' assumed multiple l operolor equi f ailures. ee TA E I. 270 400 22s vs 350 , , , , , , ,
0 250 $00 750 1000 1250 1500 1750 2000 Trne (s) Fig. X.11. Downcomer-temperature uncertainty for 0.1-m 2 steamline break from FP. 600 , , , , , 600 f*0TE: These transients go assumed multiple opercier e pmen t 650 . I a llur es. ee isle I. 2 b
$25 4go g
1s .
!,00 440 1_ .. . . .
J . 1 ,,, i, . - 400 3'
._I\
. * .?
\ ~ . -q _ . _ _ _ _ _ _ . _ _ _ _ . 360 eso \
\ '
~
320 42S 280 400 , , , , , , , , O soo saao two toaa 7%0 3000 v.co 4000 ilmo (,) Fig. X.12. 2 Downcomer-temperature uncertainty for 0.1-m steamline break 212 with two operating RCPs from HZP.
- 4. Double-Ended MSLB with Unisolated AFW to Broken SG from HZP.
Figure X.13 shows an estimate of the uncertainty of the downcomer temperature fcr this transient. The uncertainty analysis was performed using the same multiplication factors discussed in Sec. X.B.1. However, the lower error band tas not allowed to drop below 373 K (212 F). This limitation arose as the result of the manner in which the transient was specified. Normally, during a cteamline break, AFW should be supplied to the intact SG, but in accordance to transient specification, the AFW was assumed to fail to isolate to the broken-loop SG. Because the AFW flow rate was relatively low in comparison to the heat load to the broken-loop SG, the AFW entering the broken-loop SG heated to its boiling temperature at the pressure of the broken SG (0.1 MPa (14.7 psia)). This resulted in a secondary-side heat-sink temperature of 373 K (2120 F) in the broken-loop SG following the initial blowdown of the broken SG. Hence, the uncertainty error band at the lower limit cannot be below 373 K (2120 F) for the decay heat level specified for this transient. Under circumstances in which the heat loads were reduced to below ~4 MW per SG, then the auxiliary feedwater would not boil in the broken SG. The cecondary-side heat-sink temperature would equilibrate at a temperature determined by the method described by the following expression 9 T, = T, + me p where T, = secondary-side heat-sink temperature, T, = auxiliary feedwater inlet temperature, Q= heat load to broken SG, A= auxiliary feedwater mass flow rate, and c = specific heat capacity of secondary-side fluid. p The primary-side fluid temperature could drop below 373 K (2120 F) during this cteamline break. The decay-heat level would determine the final equilibrium value in accordance with this equation. In contrast, the upper error band does not have the limitations that must be imposed upon the lower error band. For the heat load produced by the decay heat from the reactor core plus the reverse heat transfer from the intact SG, 213
1 l ~ the AT between the primary and secondary equilibrated at ~5-10 K. If the heat load were doubled, then the'AT between the primary and secondary would double. The equilibrium primary-side temperature for this situation would increase by 5-10 K. This is reflected in the upper error band shown in Fig. X.13.
- 5. Double-Ended MSLB with Stuck-Open MSIVs from HZP. Figure X.14 shows an estimate of the uncertainty of the downcomer temperature for this transient. l The uncertainty analysis was performed by assuming that the mass in both SGs was
.10% higher than its average value. In addition, it was assumed that the mass flux out of the steamline break was 10% above what TRAC predicted. Hence, the energy curves for both SG A and SG B were multiplied by 1.1 in the assessment of the lower error band. In addition, the decay-heat energy curve was multiplied by .64. To assess the upper error band, the energy curves for both SG A -and SG B were multiplied by 0.9 and the energy curve for the decay heat was multiplied by 2.0. As in the double-ended steamline break discussed in Sec. X.B.4, the lower error band for this transient was limited to temperatures above 373 K (212 0F) for the decay-heat level specified for this transient.
- 6. Stuck-Open TBV f rom FP. Figure X.15 shows an estimate of the uncertainty of the downconer temperature for this transient. Because this transient was initiated from FP, the decay-heat energy curve was multiplied by 0.87 in the assessment of the lower error band. The decay-heat energy curve was not changed in the assessment of the upper error band.
Because a stuck-open TBV results in both SGs blowing down symmetrically, the energy curves for both SGs were multiplied by 1.1 to obtain the lower error band and 0.9 to obtain the upper error band. Physically, this would correspond to the mass flux out the valve being 10% higher than the TRAC prediction. Although the more rapid blowdown of the SGs would have led to SGIS at an earlier time, this effect has not been considered in this analysis.
- 7. Stuck-Open TBV with Stuck-Open MSIV from FP. Figure X.16 shows an estimate of the uncertainty of the downconer temperature for this transient.
Because this transient was initiated from FP, the decay-heat energy curve was multiplied by 0.87 in the assessment of the lower error band. The decay-heat energy curve was not changed in the assessment of the upper error band. Because the MSIV on Loop-A steamline failed to close, the energy curve for SG A was multiplied by 1.1 to obtain the lower error band and 0.9 to obtain the upper error band. The energy curve for SG B was not changed because of the loop 214
575 , , , , , , ,
- -540 550-r -
-495 525-NOTE: These transients l . ,\ ' ossumed multiple ..
g 500- , - o erotor e uipment - c E
\. f ilur es. ee ABLE H. 7L 2 ***
' .\.
-405 3 B
2 \\
' E
'\.
\ '. -360 y
E y g* i,* 450-e . , m
* ' *5
,* .' - 315 o-
-l 425- ., ,
. '., %,,' 270 400- '.,- .....................................:
1 . 1,,- . ., 225 375- ' -. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . -
- -180 350- , , , , , , i 0 250 500 750 1000 1250 1500 1h0 2000 Trne (s)
Fig. X.13. Downcomer-temperature uncertainty for double-ended MSLB with unisolated AFW to broken SG from HZP. 600 , , , , , , ,
-600
~
NOTE: These transients 500 ossumed multiple operator e pmen t f ollures. ee BLE H. g g 3,, ..
, 400 450- , n.
O E
'y
\\* ' . ,
300 3 m l '~. 6 5 400- \, '.
.T i ..... . ..........................................
- -200 350- -
- -10 0 l 300- , , , , , , ,
0 500 1000 1500 2000 2500 3000 3500 4000 l l Tme(s) Fig. X.14. Downcomer-temperature uncertainty for double-ended MSLB with stuck-open MSIVs from HZP. 215 4
, . ~ ~ - - , + - . ~ , .
,m. . - - , ,--e .=.--.--n.,. ----,--,,-n- - , , , , . - -- .,,e- , -
600 , , , , , , , , , 620 NOTE: These transients ossumed multiple
'50- operator equipmen t go f ailures. ee TABLE I.
S 5s0 S q m. e is . . . . . - -54o j g gf
\
' ~
....~.....-
-s20 a f
y 54o .'.., p fi ..,... soo d
-4so
-4s0 soo , , , . . . .
o 10 0 200 ?,00 400 500 000 730 800 900 1000 Trne (s) Fig. X.15. Downcomer-temperature uncertainty for stuck-open TBV from FP. soo , , , , , , , , , 600 5a0 NOTE: These transients ossumed multiple 570 l operator o pmen t f ailures. es BLE I. oso - g ,s, 540 g m- .,'., . -
. se
.......- f s2o- '.
- ~480 4
.n l '. -
.., 4so soo .,
' .., 420 an '.,
390 460 , , , , , , , , , o 250 -500 7s0 1000 1250 1500 1750 2000 2250 2500 Time (s) Fig. X.16. Downcomer-temperature uncertainty for stuck-open TBV with stuck-open MSIV from'FP. 216
ctagnation that occurred in Loop B. It is anticipated that the amount of r; verse heat transfer observed in the base case would not change significantly under stagnation conditions for that loop.
-D. SBLOCAs l Three SBLOCAs were analyzed in this study. The energy curves for each of these transients were presented in Sec. IX and will not be repeated.
As in the previous two classes of transients, the uncertainty analyses were performed assuming that the energy curves for the pumps and the slabs in sach of the transients were unperturbed. For the combination steamline-break cad SBLOCA transient, the energy curves for the SGs and the reactor decay heat ware perturbed in accordance to the reasons rationalized in the steamline-break cnalyses. For the SBLOCA transient, only the energy curve for the decay heat was perturbed. Both SGs operated in a reverse-heat-transfer mode throughout this transient, and it is anticipated that their energy curves would not change cignificantly as a result of a small perturbation in the decay-heat energy curve. As previously discussed in Sec. X.A, the uncertainty analysis is usually not valid for long periods of time into the transients. However, for conditions in which both SGs operate in the reverse-heat-transfer mode, the analysis is sxpected to be valid for a much longer period of time contingent upon the total cnergy added back into the primary system being small in comparison to the Gnergy removed via the cooling mechanism. This is the situation that occurred during the LOCA analyzed in this study. Hence, the SBLOCA uncertainty analysis was performed for the. entire 2-h transient period. ! For the situation obtained during the combination SBLOCA and the l cteamline-break transient, the intact SG was 1.n a "zero" heat-transfer mode throughout the majority of the transient and therefore would not be affected by the broken SG. Hence, the coupling between the two SGs is almost non-existent in this particular case. Therefore, the uncertainty analysis for this transient should also be valid for a much longer period of time.
- 1. 0.002-m2 Hot-Leg Break from FP. Figure X.17 shows an estimate of the uncertainty of the downcomer temperature for this transient. Although the major l
l ccoling mechanism for this transient was the mass leaving the primary system via the hot-leg break, a perturbation that increases the energy term would imply that the break mass-flow rate would have been larger than the TRAC prediction. Although this is a plausible perturbation, increasing the break mass flow may 217 l
- _ . ~ . . _ _ . - _ . _ . , ._.___,___ _ . _ _ _ _ _ . .- ,__ _ _ _ . . - - . --
actually cause the liquid level in the upper plenum of vessel to fall below the hot-leg penetrations. This would cause both loops to stagnate; a result that was not predicted in the original TRAC result. The liquid level in the upper plenum of the vessel corresponding to the original TRAC results did, however, come within a few inches of falling below the hot-leg penetrations. If both loops should stagnate earlier under conditions of slightly larger break flow, then the answers obtained for the uncertainty analysis using an unstagnated result as the base case would be invalid. Hence, only the decay-heat energy for this transient was perturbed. This, hopefully, demonstrates the sensitivity of the results to one of the major energy terms that can potentially alter the downconer-temperature response. Therefore, the lower error band shown in Fig. X.17 was determined solely by multiplying the decay-heat-energy curve by 0.87, whereas the upper error band was taken to be the original TRAC result that was performed with too much decay heat.
- 2. Stuck-Open PORV with Stuck-Open ADV from FP. As in the previous transient, perturbations to the break flow (that is, the PORV flow) were not considered. However, as in the steamline-break transients, it was assumed that the energy curve for the broken SG could be changed by 10% to account for a higher initial SG mass and a 10% error in the choked-flow model applied to the stuck-open ADV.
Figure X.18 shows an estimate of the uncertainty of the downcomer temperature for this transient. The lower bound was obtained by multiplying the broken SG energy curve by 1.1 and the decay heat energy curve by 0.87. The upper bound was obtained by multiplying the broken-SG energy curve by 0.9 and leaving the decay-heat energy curve unchanged.
- 3. Stuck-Open PORV from HZP. Again perturbations in the break flow (PORV flow) were not considered and, because variations in the SG mass would have no f significant impact on the heat transfer, variations in the SG mass also were not considered. Figure X.19 shows an estimate of the uncertainty of the downcomer temperature for this transient. Because this transient was initiated from HZP conditions, the upper band was obtained by multiplying the decay heat by 2.0 and the lower band was obtained by multiplying the decay heat by 0.64.
l i 218 t
3* W p-c W 600 , , , , , , ,
-600 575 i r 550 550-500 525~
8 g 500 -450 g
,'. 3 e
3 475- ,'.,N' -400 *
,f
,y 450- . . ' ,,. --350 ,p
,j' ".,
425- ', 3
' 'g -300 NOTE: These transients 400 assumed multiple Q:a.,
operator / equipment '
-250 f ailures. See TABLE I.
375- .
-200 350 , , , , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 0000 Trne (s) Fig. X.17. Downcomer-temperature uncertainty for 0.002 m2 Hot-leg Break from FP. 600- , , , , , , ,
-600 575 L -550 550-
}.
'. N '-
'N -
500 525- . - 8 \,
,, Q f E soo ', ",.. -450 g i- *3 . a
- l. ]3 \., , ' ' ', , g f 475 - -
','.. '\'.,, "'...,.
. 400 8.
E
,1- p
,y 450-
,, v'.-.,,-....,, -
'350 g
,g- ,,
j 425- v',,,, -300 I 400-Y b .,, - NOTE: These transients -250 assumed multiple ..'..,'., 375- operotor / equipment i - f ailures. See TABLE I. -200 350- . . , , , . . O 1000 2000 3000 4000 5000 6000 7000 8000 Trne (s) Fig. X.18. Downcomer-temperature uncertainty for stuck-open PORV with stuck-open ADV from FP. - 219
575 , , , . . . , , 560 550- '-.
-520
' -. , l 525- -
480 I g 500 - t r'- -
-440 p' l
- . ~, v y - '
., r.. ' . t
]e 475
~
/ ,...,,:
, . -400 2 8
a ., . F Q . ' . ,,j '% '
.\s,/'.,,,,,,- , , 3,o @
450 / ., s -
.,., -s20 425- ., -
- N -2a0
\
400 ' -. . , - 240 375 , , . i . . . . 0 500 1000 1500 2000 2WO 3000 3500 4000 4500 Tryie (s) Fig. X.19. Downcomer-temperature uncertainty for stuck-open PORV from HZP. t l I i L 1 i I I 220 i
--.m., -y.- g -- , - p.-. ,._-,n - - - - . ,
E. Pressure Uncertainty As mentioned earlier in this section, a satisfactory method for estimating j the uncertainty associated with the primary-side pressure has not yet been devised. Therefore, these authors made a guess at the uncertainty to comply with the needs of ORNL. Keeping in mind that the guess has no technical basis, it was guessed that the uncertainty in the pressure history would be bounded by assuming the following: the pressure change cannot exceed the change in the saturation pressure corresponding to the change in the downcomer temperatures predicted by the uncertainty analysis. That is, if the uncertainty in the downcomer temperature for a particular transient at a particular point in time was determined to be 50 K (90 F) (the temperature at the upper error band minus the temperature at the lower error band), then the pressure uncertainty would correspond to the change in the saturation pressure going from the upper error band to the lower error band. This method would yield a pressure uncertainty of
~1 MPa (145 psia) as a 60 error for this AT.
F. Downcomer-Heat-Transfer Coefficient Uncertainty After running the thirteen transients presented in this report, it was discovered that TRAC was not calculating the downcomer-heat-transfer coefficient correctly. Fo r the two-dimensional flow field that occurs in the vessel downcomer (azimuthal and vertical flow) the magnitude of the velocity vector should have been used to evaluate the Nusselt number in each of the fluid cells in the downcomer annulus. However, because of an error, only the vertical component of the velocity was considered. In transients in which one loop stagnated and the other loop was flowing, significant azimuthal flows occurred in the downcomer annulus. In cells in which the velocity component in the azimuthal direction was large and the velocity component in the vertical direction was small, the Nusselt number was underestimated and a natural-
- circulation flow regime was predicted. Consequently , the heat-transfer I coefficients for those cells were under-estimated by a factor of ~2-3. In addition, if the Nusselt number for those cells (based upon the vertical
[ component of the velocity) were such that the code predicted the flow regime was forced convection, the exclusion of a small amount of azimuthal flow could potentially cause a laminar-forced-convection heat-transfer correlation to be erroneously chosen rather than a turbulent-forced-convection heat-transfer correlation. This could result in a factor of ~2 in the heat-transfer correlation. Although this error affects the energy transfer from the heat 221
c1:b] to thm primary fluid, tha ovsrall potsntial to chtngs thn tseparatura history is not there. For a 200-K temperature decrease, the stored energy in the heat slabs is on the order of only 50 GW-s. In view of the relatively long time constant required to produce this energy transfer, a change by a factor of five in the heat-transfer coefficient is presumed not to change the time constant enough to cause the heat slabs to represent an important energy source in terms of retarding downcomer-temperature changes. Hence, from a thermal-hydraulic standpoint, the error in TRAC is not importan*, However, from a fracture-mechanic standpoint, the value of the wall-heat-transfer coefficient may play an important role. Therefore, it is recommended that ORNL note the erroneous spread in the heat-transfer-coefficients predicted by TRAC, choose a mean value, and assign an uncertainty to that mean value corresponding to the maximum value minus the minimum value (60). P 222-
XI. CONCLUSIONS AND RECOMMENDATIONS The~ 13 transients analyzed in this report represent 3 classes of l overcooling transients: 1) steamline breaks; 2) runaway feedwater; and 3) small-break LOCAs. They were produced by a variety of initiators combined with
. assumed operator and equipment failures. Only two initial conditions were l considered: 1) full power with a decay-heat level corresponding to an operating j history of 100% power for an infinite length of time, and 2) hot-zero power with a decay-heat level corresponding to 100 h after shutdown. The analysis of these transients was an important and necessary step in identifying and understanding the important plant parameters that affect the thermal-hydraulic response of the system during postulated overcooling transients. As a result of this study, the fo'?.owing general conclusion is made.
Of the 13 transients, those that were initiated from high decay-heat levels were thermal-hydraulically benign and may not pose a pressurtzed-thermal-shock threat. However, with only a few , exceptions, any of the 13 transients if initiated from low decay-heat level with no operator intervention would be of potential PTS concern from a thermal-hydraulic standpoint. Simple routine operator intervention might have reduced the consequences of these simulated accidents. The operator failure assumptions (particularly failure to throttle charging and HPI flow to control the system pressure-temperature relationship) were the single most important
, contributors to the generation of severe pressure-temperature I
conditions in all cases. 4 The justification for this conclusjon is based upon the following. The primary-system fluid temperature (which ultimately dictates the downcomer temperature) is a function of the net rate at which energy is removed from or i added to the primary system. As demonstrated in this study, transients initiated from initial conditions in which high decay-heat levels were present were thermal-hydraulically benign. The energy addition from the decay heat counter-balanced the energy removed by rapid cooling mechanisms and precluded low primary-system fluid temperatures. Because-the major source of energy addition to the primary system is the decay heat from the reactor core with minor additions occurring via the primary-system structural { l material, work from the RCPs, and reverse heat transfer from one or two SGs, i: a significant reduction of the energy from the decay heat would lead to a ! t L more severe cooldown for the same cooling mechanism. 223
In general, initiating most of the transients in this study from a low decay-heat level would eventually produce downcomer temperatures well below 405 K (2700F), the NRC screening criterion for assessment of PTS risk. However, the rate at which the downcomer temperature decreases is highly dependent upon the cooling mechanism and may be less than the prescribed
)
cooldown rate that the plant operators follow during normal shutdown ! operation. For these situations, the overcooling transients may not be of , concern from a fracture-mechanics standpoint. The exceptions to our conclusion involve transients in which the j cooling mechanism is stopped early in the transient by operator or equipment action. For example, the stuck-open TBV transient presented in Sec. VIII.C.1 was thermal-hydraulically benign because SGIS occurred early which isolated the stuck-open TBV and ended the blowdown of both SGs. Another exception to the conclusion is the runaway-main-feedwater transients presented in Sec. VII. The total cooling capability of a runaway-main-feedwater transient is limited by the total liquid inventory in the condenser /hotwells. Hence, for runaway-main-feedwater transients initiated from low decay-heat levels, the minimum secondary-side-heat-sink temperature would be essentially unaltered thereby limiting the primary-side-temperature decrease. Because the decay heat is a function of the operating history of the reactor, the decay-heat level of the reactor is not a fixed value. For l example, the decay-heat level corresponding to FP operation can vary significantly. If the reactor has been operating for one year at 100% power level, a decay energy of ~300 GW-s would be dissipated during the first 2 h following a scram. On the other hand, if the reactor had been operating at FP for only 1h following a 60-day refueling outage, the decay energy of
~75 GW-s would be dissipated during the first 2h following a scram (see Figure XI.1). Hence, to claim that all 0.1-m 2 (1-ft )2 steamline breaks initiated from FP will yield similar results is incorrect. The operating history prior to the initiation of the transient must be specified to completely define the decay-heat level that will be present during that i
transient. Similarly the results obtained for the transients initiated from HZP correspond to a decay-heat level obtained 100 h after shutdown following i' infinite operation of the reactor at 100% power. However, the decay-heat level at HZP need not be the value specified by ORNL for the transients 224
cnclyzrd in thio study. Figure XI.2 shows tha dacsy-hnet 1sval that say be present in the reactor core at the HZP conditions following refueling outages of different lengths. Note that the probability (to be quantified by ORNL) for a particular initiator to occur at a particular decay-heat level is a function of the operating history of the plant and the probability that that initiator will occur. It is presumed that a normal operating plant will be at FP conditions more often than at HZP conditions. Coupled with the relatively short period of time of FP operation required to build in a significant decay heat (see Fig. XI.1), the thermal-hydraulic response of the system corresponding to low decay-heat levels may not be of concern from a risk-assessment standpoint. From a thermal-hydraulic standpoint, the decay-heat energy is a very important parameter that plays a major role in determining the overall thermal response of the primary system. Directly coupled to the decay-heat level are the equilibrium loop flows t'vt occur in the primary system following the tripping of the RCPs. For the steamline-break and runaway-feedwater cases, the decay heat was sufficient to induce loop flows in one or both loops. During the SBLOCAs (without an additional secondary break), both SGs operated in the reverse heat-transfer mode. Combined with a low decay-heat level, flow stagnation did occur in both loops (see Sec. IX.C). In this situation, little mixing of the HPI and - cold-leg fluid occurs and so the HPI flow (high during an SBLOCA) is virtually in direct contact with the vessel wall. Stagnation in both loops is not confined to any particular break size. In fact, producing a rapid primary-side cooldown while maintaining a relatively high primary-side pressure becomes a distinct possibility for break sizes that are smaller than the 0.002-m2 break. This, of course, is contingent upon a low decay-heat level. i SG-tube rupture transients may become of interest for situations in which the decay-heat levels are low. The thermal-hydraulic consequences of a SG-tube rupture are felt to be identical to the consequences that would be obtained for a hot-leg SBLOCA of equivalent size provided that choked flow in the ruptured tube is maintained throughout the transient. If choked flow is not maintained in the ruptured tube (because of the high secondary-side pressure of ~6.2 MPa (~900 psia)), then the mass flux out the ruptured tube will decrease relative to the mass flux out the equivalent hot-leg break and 225
350 , , , , , , , 9a 300- - 8a 2m 1m 2v 250- - 1e 7 20c- 1d h - 4 -
@ /
j ro- -
/
100- ,f - Ih so- - 0 , , , , , , , 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. XI.l. Decay-heat energy dissipated after shutdown for a typical three year operating history as a function of operating time into the current reload cycle. 20 , , , , , , , , , 17.5 - - 15 - - l u.5 - - m- - ! f 7.5 - - 5- - 2.5 - - , 0 to 2'o 80 /0 do 80 7'O s'O do 10 0 Time (days) Fig. XI.2. Decay-heat power after shutdown following a typical three-l year operating history (calculated at end of cycle). l 226 t f
l I l the cooling effect of the SG-tube rupture will not be as large as that produced in an equivalent hot-leg break. Nevertheless, a SG tube rupture may still produce rapid cooldown of the primary system at low decay-heat levels and may pose a PTS risk from a thermal-hydraulic standpoint. An additional conclusion related to loop stagnation and RCP operation I has also been inferred from this study. In situations in which the RCPs are tripped and one SG is operating in a normal heat-transfer mode while the other SG is operating in a reverse heat-transfer mode (for example, a steamline break), the flow in the intact loop will be very low or stagnant. This will effectively eliminate one-third of the primary-system heat capacity. Hence, any energy removed from the primary fluid via the broken loop will produce a larger temperature decrease in the downcomer annulus per unit of energy removed. Running the RCPs in this situation effectively doubles the heat capacity of the primary system because the RCPs couple the large amount of stored energy in the intact SG to the primary system. Hence, the temperature decrease per unit of energy removed via the broken SG is reduced by one-half in comparison to the situation in which the RCPs are tripped and the intact loop stagnates. In situations in which both SGs are operating in a normal heat-transfer mode (such as a double-ended steamline break in which the MSIVs fail to close or runaway-main-feedwater to two SGs), running the RCPs will increase the l i rate at which energy is being removed from the primary system. This will produce lower downcomer temperatures in comparison to the case in which the RCPs are tripped. Hence, it cannot be concluded that running the RCPs in all transients is beneficial to minimizing the temperature decrease produced by an overcooling mechanism. It must be determined for each transient taking into account the heat-transfer mode that will occur in the SGs. In conclusion, we have shown that the cooling capacity for the three classes of transients analyzed in this report is insufficient to produce low downcomer temperatures when initiated from high decay-heat levels and sufficient to produce low downconer temperatures when initiated from low decay-heat levels. Steamline breaks possess the largest potential to produce rapid cooldown. If initiated when the SG secondary side water mass is large, the subsequent primary-side temperature decrease will be greater than if the SG secondary side water mass is small. SBLOCAs possess larger potential for overall cooling of the primary system but the rate of cooldown will not be as 227
large as the rate produced by steamline breaks. Runaway-main-feedwater 4 transients can produce rapid, but short-lived, cooldown of the primary system. The liquid mass in the condenser /hotwells limits the cooling potential for this class of transients. The energy-analysis approach used in this report enabled us to identify the important heat-transfer mechanisms during an overcooling transient. We I were also able to perform better sensitivity analyses based on this method, and we recommend this approach for similar PTS analyses. e 228
l ACKNOWLEDGEMENTS l The authors would like to recognize the efforts of those who also l contributed to this project. We appreciate the cooperation that we received from those at the Baltimore Gas and Electric Co . and Combustion Engineering, Iac. Without their extensive time and effort, the quality of our work would l have been severely limited by lack of accurate information. The people who were t involved included Trevor Cook and Carl Yoder of BG&E and Dave Earles and Gerhart Menzel of C-E. We thank Doug Selby of ORNL who worked with us closely, both in the transient specifications and the use of our results. We also want to acknowledge the managers of the program: Lou Shotkin, Jose Reyes, and Carl Johnson of the NRC, Jim White of ORNL, Eric Titland and Steve Marsky of BG&E, and Dan Peck of C-E. At Los Alamos, the work could not have been accomplished without the prior efforts of those who developed the TRAC-PF1 computer code, particularly John Mahaffy, Dennis Liles, Manjit Sahota, and Susan Woodruff. Lois Sylvia and Sylvia Lee worked hard with us processing the graphics. Special thanks goes to the word processors for their helpfulness and patience: Cecilia Gonzales, Janet Stanford, Jean Martinez, and Wendy Rowley. I l 229
l I REFERENCES
- 1. Safety Development Group, Energy Division, " TRAC-PF1: An Advanced Best-Estimate Computer Program for Pressurized Water Reactor Analysis," Los Alamos National Laboratory report LA-9944-MS, NUREG/CR-3567 (February 1984). .
l
- 2. Gregory D. Spriggs, " Heat Slab Formulation for TRAC Input," Q-7-83-568, Los Alamos National Laboratory document to TRAC distribution (December 16, 1982, revised January 1, 1983).
- 3. William C. Reynolds and Henry C. Perkins, Engineering Thermodynamics McGraw-Hill, Inc., 2nd Edition (1977).
- 4. Gregory D. Spriggs, " Basis for Simplified TRAC Model of tihe Secondary Side of Calvert Cliffs for Use During the Pressurized-Thermal-Shock Studies,"
Q-7-83-193, Los Alamos National Laboratory letter report to Jim D. White, ORNL (April 21, 1983).
- 5. Baltimore Gas & Electric Co., Updated Final Safety Analysis Report of Calvert Cliffs-1 (July 1982).
- 6. J. L. Darby, letter to N. S. DeMuth dated November 11, 1982.
I l 230
JE NOTE TO THE APPENDIXES As explained in Sec. VI, an ef fort was made in the main report to detail the energy transfers in each transient and not necessarily to report every event cf each transient. The plots in these appendixes are provided for the motivated l raader who would like more information on the tratisients. j i i t l^ i 4 231
,e . - , - - - - . , , - - . - - - , . . - - , , - - , - - - - , - , , ,. ----
-,v,-. ~, - .
APPENDIX A RUNAWAY MFW TO TWO SGS FROM FP l t l l l l l l l 232 l t I l .
300000 . . . . . . . -660000
~630 3 NOTE: These ironsionts <
assumed multiple 280000-operator - equipmen t . . f ailures. ee TABLE I. -s00000 260000-. 37oooo
<a 2 o
- C j - s40000 j 240000-
- 510000 220 M - , -480000
. 43oooo 200000 . . . . . . .
0 1000 2000 3000 4000 5000 6000 7000 8000 Tme (s) Fig. A.1. Primary system mass--Runaway MFW to two SGs from FP. 1 . . . . . . . THUA 1
. .. . .. .. .. THUA 2 i o.3 - ~~------ THUA 3 .
- THUA 4 l THUA $
THUA 6 0.6 - - l 8 I NOTE: These transients assumed multiple
, o.4 _
o operator e ment p f ailures. ee LE H. 0.2 - 0- " -
-0.2 . , , . . . .
0 10:;0 2000 3000 4000 5000 6000 7000 8000 Time () l Fig. A.2.
- Voiding in the upper head--Runaway MFW to two SGs from FP.
l 233
to . . . . . . .
-30 8 .
25 1 l G-
. -20 l A.
o 6 y -
-15 y
.3 r. .
3 u j NOTE: These transients O ji 3: assumed multiple g operof or / equipment 2- f ailures. See TABLE H. - 5 0- - O
-~5
-2 . . . . . . .
0 1000 2000 3000 4000 5000 0000 7000 0000 Timo (s) Fig. A.3. Pressurizer liquid level--Runaway MFW to two SGs from FP. 5000 , , , , , , , TwcTA -800
.......... THETA 2 THETA 3 g
4000 -
- THETA 4 THETA 5
--700 a' 0-THcTA e
~
-600 2
, ::- - 8 ! j 3000-
-']
i ,
? -s00 8 l y l
k NOTE: These transients ? I assumed multiple J! l p 5 ' i l operator / equipment ' d# E i 2000- :.., . . .... f allures. See TABLE H. .
>2
\ fii.f i -300 % l - ' ~200 1000 I _- % l l f/ -100 l 1 0- . . . . . . . 0 0 900 2000 3000 4000 SC00 6000 7000 e000 Time (s) Fig. A.4. Downcomer heat-transfer coefficients at the core midplane-- Runaway MFW to two SGs (from FP). 234
s00 , , , , , , ,
$90- .
-_g j se0- -
. -575 l 570 - - g Loop a g ,. _
.. . . . . . .. . too, s
. -sso v 3'3 l W- -Il -
p
- l. -s2s g
@ S40- -
E
.T 4
,P sw- '"
-1
'E NOTE: These transients 3 520- ossumed multiple operator / equipment '-475 f ailures. See TABLE H.
50- - _4so 500- . 400 - , , , , , , .
- 42s 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s)
Fig. A.S. Hot-leg liquid temperature-Runaway MFW to two SGs from FP. 1 . ,
-ooz e 0.8 ]gO _
!i; 5 "g
- "1o1; g
~
g %!I ,;
~
_ .=_ , o $2a PE
- E,
( $ 0.4 - - l .E "5 8 ! 9 ma ; 0.2 - .
, , i; 0 '
L bal, 'Y
-01 , , , , , , ,
. 0 1000 2000 5000 4000 5000 6000 7000 8000 Time (s) Fig. A.( PORV flow area fraction-Runaway MFW to two SGs from FP. , l i 235 l
M i . . . . . 800-.
.-F50 LOOP A 700- .. ........ LDor s
--1500 600- -
1250 500- - NOTE: These transients '8000 d 400- ossumed multiple - operator / equipment b
~ 750
,_ f ailures. See TABLE H. _
g_
-500 10 0
-250 0-- --O
-10 0 . , ,
O S0 10 0 ISO 200 250 300 350 h (s) Fig. A.7. Main feedwater flow-Runaway MFW to two SGs from FP. 0.06 , . . . . . . 0.132277 0.04- -
- 0.079366 LDor A 0.02- . . . . . . . . . . too, ,
- u.02uss l } Q.00 -
g i s 3 -
--0.026465g l -0.02 - -
l NOTE: These transients l ossumed multiple operator / equipment ._o,o7, m _o,o4 f ailures. See TABLE I. _ l
-0.06 - , , , , , . . -0.132277 0 1000 2000 3000 4000 6000 6000 7000 8000 l Time (s)
( Fig. A.8. Auxiliary feedwater flow--Runaway MFW to two SGs from FP. l 236
250000 . . . . . . . 500000 LOOP A 200000-- . . .-.-.... too, e -
-400000 20000- -
-300000 9 i
2 5 2 100000- \ -
-200000 NOTE: These transients assumed multiple 50000-- operator / equipment ~-100000 f ailures. See TABLE H.
0 , . , , , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. A.9. Steam generator mass inventory-Runaway MFW to two SGs from FP. 800 -
, . . . . . . -1750 700-_ -1500 LDoP A soo. .. ... . . ... too, e -
-1250 500- -
g -1000 l g '00- NOTE: These transients assumed multiple o
, _73o g 300- operof or / equipment f ailures. See TABLE I.
t 200- -
~
10 0 - 0- I' -
^
-~
--O I
- 10 0 -.
-. 250
-700 . . . . . . i 0 1000 2000 3000 4000 5000 6000 7000 8000 lime (s)
Fig. A.10. Steamline mass flow--Runaway MFW to two SGs from FP. 237
s.3 , , , , , , , L2- -
-- 900 9
LOOP A 81- - _ ,,o
/
W
. . . . . . . . . . to o, ,
i 6- l -
.\
-se0 s.9 - -
I
. ?a s.s-.I / --s40 S
/ 2 I
5.7- / - il ' E6-
/
,/
-820
}
53-.
,-,' NOTE: These transients ~-800
,/ assumed multiple operator / equipment s.4 - . 's.....,,,,,,,,,, f ailures. See TABLE H. -
_ 7,, s.3 - -
-760 s2 , , , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. A.11. Steam dome pressure--Runaway MFW to two SGs from FP. 0.9 , , , , , , , 0.8 - 0.7 - 0.6 - NOTE: These transient s - ossumed multiple u
" 03-operator / equipment ~
f ailures. See TABLE H. 0.4 - h 0.s - I 0.2 - , 0.1 - 0.0 - l
-C.1 , , , , , i ,
0 1000 2000 3000 4000 5000 6000 7000 8000 lime (s)
- Fig. A.12.
l TBVs flow area f raction-Runaway MFW to two SGs from FP. 238
l l l l 0.9 . , , , , . . l l 0.8 - - LOOP A 0.7 -- . . . . . .. .. . too, , . 0,6 - - i r.
- l. 0 0.5 - -
o NOTE: These transients j o& ossumed multiple . operator / e ipment I f ailures. See ABLE H. 3 0.3 - - o.2 - - 0.1 - I 0.0 - -
-0.1 , , , , . . .
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. A.13. ADVs flow area fraction-Runaway MFW to two SGs from FP. i l I l I l l l l l-i l l l i l 239
.-_ - - _ _ _ __ _ _ - _ _ _ . . . . ~ . _ _ __ -_ . - _ _ _ _ _ _ _ _ _ . _ _ _ _
i 1 APPENDIX B RUNAWAY MFW TO ONE SG FROM FP t f f 240
300000 - i i i i i i i -660000 I- -
-630000 280000-
- -600000 2 WOO- , -570000 m
- 9 o
6 M M
/ -540000 h 240000-
- -510000 NOTE: These transient s assumed multiple 220m- - operator oquipmeni ~-480000 f allures. ee TABLE H.
- .m 200000 , , , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 8000 Tirne (s) Fig. B.l. Primary system mass-Runaway MFW to one SG from FP. , 0.030 , , , , , , , , , , , THETA 1 0.025-j ~~.~~.~..~"...~-NA2 THETA 3 i - THETA 4 THETA 5 [ I THETA S 0.020-t 8 0.015 - J _ 6 NOTE: These transients l o assumed multiple 0.0:0_ operator e pmen t _ f ailures. ee BLE I. 0.005- .
~~
N "" - 0.000- - l
-0.005 , , , , , , , . , , ,
i 0 500 1000 1500 2000 2500 J000 350C 4000 4:40 5000 5500 6000 Time (s) Fig. B.2. Voiding in the upper head--Runaway MFW to one SG from FP. t t 241
10 . . , , , , , , , ,
. 30 8- .
-25 6- .-20
$, o e
_g i> 3 - 3 ; u
$ 5
-10 r, 3: 3:
NOTE: These transients 2- t assumed multiple . operator / equipment _s f allures. See TABLE H. 0- -
--O
-2 , , , , , , , , , , ,
0 500 1000 1500 2000 2500 J000 3500 4000 4500 5000 5500 6000 Trne (s) Fig. B.3. Pressurizer liquid level-Runaway MFW to one SG from FP. 500c , , , , , , , wrA -800
. . . .. ....* THETA 2
{ MTA 3 p 4000- . ; l
- wrA4 -
-700 a' THCTA $ .C THCTA 6 L
-ooz oV#O -600 { _
* -*EM p
,o3-E
=
j 3000- -X - eQ }
; 500 (7~-
s kg
- I w E . t I, f / /~A
/
NI
*.E 400 5 g
a 2000- : t 'l .! fN 4> _.3 3-s mu*, f -
-i : , \ /t . a
-300 I
S l' l33
.T ! -
./.,.
* .! i../ \
e 3 n
, k': 's f; '/
. y',. s N ""* -200 f --\ 4 0
j 1000-
- - 10 0 0 , , , , , . , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s)
Fig. B.4. Downcomer heat-transfer coef ficients at the core midplane-- Runaway MFW to one SG from FP. l l 242 I i i I
r-590 , , ' ' ' ' ' ' ' ' ' -600 580- -
-580 570- -
-560 LOOP A 8 560- , .... . . , ,
{
-540 550- ,,
- 520
* -- \
a 3 3, g- - -500 ;
- *~ NOTE: These transients
~
ossumed multiple 520-
~
/ operator equipment --'80
/ f ailures. ee TABLE H.
510 -
- .-460 500 , , , , , , , , , , ,
0 500 1000 .1500 2000 2500 3000 3500 4000 4500 5000 5500 6000 Time (s) Fig. B.S. Hot-leg liquid temperature--Runaway MFW to one SG from FP. 12 , , , , , , , , , , , NOTE: These transients assumed multiple 1-operator equipment f ailures. ee TABLE I. 0.s - 9 v 0.6 - 0.4 - i i - 0.2 - o 1 ka L _ 0 . ._ _ I _4 _
-02 , , , , , , , , , , ,
0 500 1000 1500 200C 2$00 3000 3500 4000 4500 5000 5500 6000 Tirne (s) Fig. B.6. PORV flow area fraction--Runaway MFW to one SG from FP. 243
l na , , , , , ,
)
I LDOP A
........ too, . - 24#
1000- . I
-2000 600- g 1600 6 600-o j -noo g
- C 400- i . g t -800 o t NOTE: These transients 2 i cssumed multiple 200-.i operof or / equipment -.g
; f ailures. See TABLE H.
I
- a. . ' . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3. . . . . . . . . . . . . . . . . .
-200
--40 0 60 10 0 16 0 200 260 300 360
' Time (s)
Fig. B.7. Main feedwater flow-Runaway MFW to one SG from FP. 22.5 , , , , , , , , , , , 20- .
-40 v.5 - .
15 - WOPA -
..........wo,, 3o a
gtL u.5 - .
- m 10 -
Q - NOTE: These transients -20 0 8 75- assumed multiple U operator / equipment 2 3 i f ailures. See TABLE I. I S-_ - 2.5 - - 0 --0 l
-2.5 , , , , , , , , , , ,
0 500 1000 1500 2000 2500 3000 3500 4000 4500 $000 5500 6000 Time (s) Fig. B.8. Auxiliary feedwater flow--Runaway MFW to one SG from FP. l l l 244 L
250000 . . . . . . . I
- -500000 LOOP A 200000- .. ...... .. too, s -
- -400000 20000-
[ -300000 [ H NOTE: These transients g"
$ ossumed multiple 2 100000- operator / equipment -
f ailures. See TABLE H. -200000 50000- , , , . . . . . . . . . . . . . . . . . . . . . . . . , , , , ,,,,nn
' ' ', ~ . . . _ , , , , , , ' ,
0 . . . . . . . 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. B.9. Steam generator mass inventory-Runaway MFW to one SG from FP. 1000 , , , , , , , , , , , l ! 900-- --2000 t
/ LOOP A 800-. ,,,,,,,,,,Lg,,, .-1750 700-,
- 500 l Q 600-
~ '28*
! g ~ f 8U'-
~
- NOTE: These transients -'"'
- ossumed multiple d
~
d g '00- Operator / equipment - n g . f ailures. See TABLE H. -750 j 300-
-500 200-30 0 _
.-250
- g. _f.'...k...... ......._ _ _ _______
_ .l. W_ .-0 1
- 10 0 . . . . .
0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000 5500 6000 Time (s) Fig. B.10. Steamline mass flow--Punaway MFW to one SG from FP. 245
e.s , , ,, , , , , , , , e.2 -- . . -900
"^
f/i
- n. ,s'............' ..... . .... soap a s.1 -- 1,
,- ...'..,,/*,
--ses I
g'? s-- --sm x " 7 b 5.s -- h e
--ass [
se-. .., i NOTE: These tr cnsien t s ossumed multiple ~-a25 operator egu !pmen t f ilu r es. ee .ABLE H. u- - (,-ano 5.5 -- , , , , , , , . , , , 0 500 1000 1l'40 2000 2500 3000 3500 -009 4500 5000 6500 6000 Time (s) Fig. B.ll. Steam dome pressure--Runaway MFW t.3 ona SC from FP. t2 , , , , , , , . , , , , 1- - 0.6 - - G v
, 0.6 - -
P w NOTE: These transients
- assumed multiple f 04-operator / equipment f ailures. See TABLE 2.
l 0.2 - - 0- t
-01 , , , , , . , , , , ,
0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000 5500 5000 Time (s) Fig. B.12. l TBVs flow area fraction--Runaway MFW to one SG from FP. 246
. ~ , . . . _ - - __. . . _ - , _-. .
L2 . . . . . . . . . . . 1- toor A
. . . . . ... . . m a t _
0.s - G v 0.6 - NOTE: These transients assumed multiple f
) operator / equipment
~
4 f ailures. See TABLE H. a 0.2 - 0- _
-01 . . . , . . . , . . .
0 500 200 200 2000 2500 3000 3500 4000 4600 5000 5600 6000 h (s) Fig. B.13. F- ADVs flow area fraction-Runaway MFW to one SG from FP. i I~ 247 i l
I I l APPENDIX C RUNAWAY AFW TO TWO SGS FROM FP l l 248
300000 i , , , -c,60000
-G30000 280000- -
-300000 200000- -
370000 o> 2 6 o E E g - 7 f-s4c000 y 240000- -
~ 5' 0 NOTE: These transients ossumed multiple 22c00 -
operator / equipment ~ f allures. See TABLE I. -4s0000
-450000 20C000 . , , , , , ,
0 .1000 2000 J000 4000 5000 6000 7000 8000 Time (s) Fig. C.l. Primary system mass-Runaway AFW to two SGs from FP. s 0.06 , , , , , , , THETA 1 , . . . .. . . . .. THpA 3 0.04- THUA 3 -
- THCTA 4
- -- THC'A 5 THCTA $
0.02- - 8 e i $ l' I 0.00 - a ! 9
-0.02 - -
NOTE: These transients _a g, _ ossumed multiple operator / equipment f ailures. See TABLE H.
-0.06 . , , , , , ,
O 1000 2000 3000 4000 5000 6000 7000 8000 l Time (s) Voiding in the upper head-- aw kFWtotwoSCsfromFP. l l ! 249
=
1 10 , , , , , , , g . -30 8 - 25 7- - e,
} s . 20 3
0 b b 5 3 - 3' 3: is y 4 . 3' NOTE: These transients ossumed multiple - O operator / equipment 2 f oilures. See TABLE I. . s 1- . , , , , , , 0 1000 2000 3000 4000 5000 0000 7000 8000 Timo (s) Fig. C.3. Pressurizer liquid level-Runaway AFW to two SGs from FP. 5000 , , , , , , , THETA 1 -800 THETA 2 THETA 3 p 4000- -
- THETA 4 THCTA 5
-700 **
THCTA 6
~
600 2 t e . O o 3000-
- C 8' '
t* u . g % i j 300
< .- Eo a S .. I e sh g, 2000-E Eh6
-.g%
l
-1
-400 )p
; p
=...,, ,
x .
;E ,
300 g
.y .c z
~ ,. .g e. t ,
~ .a g t -
- o
~
i ' 1000-O e o. O
~%s eN '- ' -
ZQO* D
-10 0 i- 0- , , , , , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) l Fig. C.4.
l Downcomer heat-transfer coefficients at the core midplane-l Runaway AFW to two SGs from FP. l 250 i
600 , , , , , , , 590-- --600 t ,_ 580- - 373 570 - - l g ~~********20 Loor A g 560- --550 v 2 g
, 550- -
g.
-525 f.
-@ 540- -
f v 4
-500
].
33o- 3 g
\
520 -. NOTE: These fronsients -_ ,73 assumed multiple 510 - operator / equipment _
- f allures. See TABLE H. _ ,3o 500- -
490 - , , , , , , ,
-G 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s)
Fig. C.S. Hot-leg liquid temperature-Runaway AFW to two SGs from FP. L2 . . . . . . . I-
-eO O 2 -
!" OVwo l 0.5 - [03-N e -*
- q -
n a
*R"1 l
n 0.6 - o!:* o , ::g , - I. d"o agea f 0.s - pg g a s P8 '. ; 0.2 - - 0 J- - l
- -02 , , , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. C.6. PORV flow area fraction-Runaway AFW to two SGs from FP. l 251 I l
-. - - .-. . - . . - - . - , - , - - ., . . - - - - - -. ,. . - = . , . , - - - . . - .-
7 soo .
, , , , . . . . .nso 700- -
wo, , -1500
. . . . . . . . . . ma, ,
,o,_
-1250 500- ~
f -
-1000 g
o doo- g 1
's
} ~
NOTE: These transients
~ #3 E
3ao_ '\ a f
\ ossumed multiple
\ operator / equipment j
- 5#
f allures. See TABLE I. 200- \, _
\.
10 0 -- g _-250
....,,~~..
o-- g
-8 i . . . , , , , ,
o to 20 ao 40 so so 70 so so too Trne (s) Fig. C.7. Main feedwater flow--Runaway AFW to two SGs from FP. 30 . , , , , ,
~
-60 25- _
-So
~ ~
toora
..........wo,, -4o q
15 - o y _
^30 3 o
C c3 n g" to-
^
-20 0 2 2
*~
NOTE: These transients ~~ 10 assumed multiple operator / equipment o f ollures. See TABLE H. -
--o
--to
-5 , , , . , , , '
0 1000 2000 3000 4000 5000 6000 700o 5000 Time (s) Fig. C.8. Auxiliary feedwater flow--Runaway AFW to two SGs from FP. 252 l l
250000 , , , , , , .
-500000 LOOP A 200000- ---------- toon s -
-400000 NOTE: These transients 150000- ossumed multiple 9 - operator e pmen t . -300000 ?
f f ailures, se ABLE I. 7 h jogooo. .
-200000 50000- -_inonno O , , , . , , . 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s)
Fig. C.9. Steam generator mass inventory-Runaway AFW to two SGs from FP. 900 . . . . . i i 800- --1750 LOOP A 700-- -------*-- LDOPe
, -1500
~
600-
~
500-
! 6 -1000 O i a
- y 400-
~
d l n - 750 M
~
300- NOTE: These fransients ossumed multiple -500 i 200- Operator equipment - l f ailures. ee TABLE H.
~
10 0 - 0-. -0 l
- 10 0 , , . . > > >
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. C.10. Steamline mass flow--Runaway AFW to two SGs from FP. 253
7 * . . . . . _ gnn 6.5 - - Loop A -goo 6- .. ... . ... . Loop a . 5.5 -- --800 5- -
?
f_
-700
'5-R -f
-'.. .! )
J -600 i 3.5 -. : . -_ son NOTE: These fransients 3- assumed multiple - operator / equipment -400 2.5 - f ailures. See TABLE H. h 2
~300 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s)
Fig. C.11. Steam dome pressure-Runaway AFW to two SGs from FP. 1.2 . . . . . . . t-0.8 - - G o.. . NOTE: These transients _ E assumed multiple 4 operator / equipment 1 y f ailures. See TABLE H. 0.4 - - g l 0.2 - - i
' 1 0-
-0.2 . . . , . . . !
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. C.12. TBVs flow area fraction--Runaway AFW to two SGs from FP. 254
1.2 , , , , , , l 1- p LOOP A
.. ... . . . . . too, .
0.5 - d v I 0.6 - - 5 e NOTE: These transients 3 OA- ossumed multiple -
# operator equipmen t f ailures. ee TABLE I.
0.2 - 0-
-01 , , , , , , .
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. C.13. ADVs flow area fraction-Runaway AFW to two SGs from FP. 4 255
f l I 1 i 1 1 i ! APPENDIX D 0.1-m2 MSLB FROM HZP E i t
- j. .
I J 1 1 256
300000 . . . . . 4 -a60000
- -630009 280000- -
- -300001 260000-
. 37o0 0:3 Tn T C'
j 6
,, p
- -540000 ]
s 240000-
- s. . NOTE: These transients . r.iococ, assumed multiple operator / equipment 220000- f allures. See TABLE H. -
, 180000
- -450000 200000- . . . . . . .
0 1000 2000 3000 4000 5000 0000 7000 8000 Time (c) Fig. D.l. Primary system mass--0.1-m 2 MSLB from HZP. 45.0 . . . . . . . 40.0- NA 1 -
. . . . . . .. .. m. :
.-. wA s 35.0- - " " ' -
THETA S THLTA 4 30.0- -i
.h 25.0-1 -
I 20.0-t f 15.0-NOTE: These transients ossumed multiple operator / equipment f ailures. See TABLE I. . t i 5.0 - , L............._ _ m ., _ m .m . . _ _
-5.0 . . . . . . .
0 1000 2000 3000 4000 5000 6000 7000 8000 T h (s) Fig. D.2. Voiding in the upper head--0.1-m2 MSLB from HZP. 257
1 l 17 5 , , , , , , ,
~
15 . 50 l 12.5 , 40 3y " - o 30 ? i 75 - 3 W L. 3
#)
20 j s _ D NOTE: These fransients ossumed multiple 10 2.5 - operator / equipment - f ailures. See TABLE H. 0- . 0
-23 , , , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 8000 Timo (s) Fig. D.3. Pressurizer liquid level--0.1-m2 MSLB from HZP. 5000 , , , , , , , i l wra s .oag -800
- )
l -- _- --- - .MTA grA 2 3 O.; V W Oe W _g M' 9 4000- i _ ], jo3 U -
-700 af
; -m. :4
* *1
. { b! g E s
t 3"' i a.a. ~ 5 6 , i .,
-13.2 -500 ka a O 1 I E r R34 a"
g -
., t}, i 3 400 P
2000-
., p3"* -*
g e (L.w I -300
+-
g l I -
- 7. }~
V f\ l $ I
) '
,,- P Ps% [ fi \l i . ,.., 1
* ~***
5 '
~ '
ww- - N, c-
\p =-- -
m0 o , , , , , , , 0 0 1000 2000 3000 4000 5000 5000 7000 8000 ) Time (s) I l Fig. D.4. l Downcomer heat-transfer coefficients at the core midplane--0.1-m 2 MSLB from HZP. l I t 1 258
seo , , , , , , ,
- - 520' 540-
. -4s0 520- ,
WA
. . . ... . .. . m, a y 2 500--- --44 2
NOTE: These fronsients assumed multiple f& g 480-Operator / equipment E y A y f ailures. See TABLE H. p 4s0-y - -sso g 440- y -
- -320 420- ' . - .. , ,
- -280 400 , , , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. D.5. Hot-leg liquid temperature--0.1-m2 MSLB from HZP. 12 , , , , , ,
~
I
- mi
.5 E's -
0.8 - S 8.b$
*E ?rH 9 -E .
0.6 - N o . k 5 Y
. .. s o - -
0.4 - W 3 g5
>f ocao zoo
- 0.2 -
~~ ' * -- -
0
-0.2 , , , , , , ,
8000 0 1000 2000 3000 4000 5000 6000 '/000 Time (s) Fig. D.6. PORV flow area fraction--0.1-m 2 MSLB from HZP. 259
I l 0.0a , , , , , , , 0.1322r7 0.04- "" -
.......... wo, e - 0.079366 In this transient, the rain 0.02- feedwater flow was set to -
0.0 kg/s initially instead of 9 being ramped down from 10.0 kg/s. - 0.026455 ) o k
- g 0.00 -
M
--0.02sf
-0.02 - -
NOTE: These transients assumed multiple
- operator / equipment _ _c,ny,3g
-0.04 - t ollures. See T ABLE II. .
-0.06 . . , , . . . -0.1322r1 0 1000 2000 3000 4000 5000 6000 7000 8000 Tkne (s)
Fig. D.7. Main feedwater flow--0.1-m2 MSLB from HZP. 22.5 , , , , , , 20- , - 17.5 - - IS -- l too,A - gcm u.s -
.. ... .... . wo, a 3, 6 i o y : ,
d '0 - ! - g
- t. -20 m
$ 7.5 -
l NOTE: These fransients - j ossumed multiple 3_ : operof or / equipment t f ailures. See TABLE H. _-10
! l 2.5 - l -
O -
-O
-23 , , , , , , ,
0 1000 2000 J000 4000 5000 6000 7000 8000 Time (s) Fig. D.8. Auxiliary feedwater flow--0.1-m2 MSLB from HZP. 260
250000 , , , , , , .
-500000 200000- .**,,,**.. -
-400000 l
,,**' mop A t
150000- ........ .. wop a n g -
,...',, -300000 8,,
....- 8 8 ,,..* 2 2
WOOOO- ... -
-200000 NOTE: These transients assumed multiple 50000--
operator / equipment f ailures. See TABLE H. .-100000 0 , , , , , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. D.9. Steam generator mass inventory--0.1-m2 MSLB from HZP. 600 , , , , , , i
- -1200 500- mop a l ..... .... . wo, s -1000 400-
. -800 d$, 300-
- 3, c
y -s00 m o NOTE: These transients g assumed multiple l 5n 200- operator / equipment - n
~'00 l 3 f ailures. See TABLE H. 3 WO- _ -200 l
0 ... _...... h ,
?
--0 l
^
~~#**
- 10 0 , , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 8000 l Time (s) Fig. D.10. [ j Steamline mass flow--0.1-m2 MSLB from HZP. l 261
7 , , , , , , , 6- -
-900
,...,-s.
..........m, 5- ' . . . , . * ' " , -
; i
-se 9 4- i ,-
I. _
- . . . -: .- : 7_
1 . , g a o 3_ i i. - 450 E g 2- ',. --300 8b NOTE: These transients '. .'- assumed multiple t- operator / equipment --so f allures. -See TABLE I. 0- -
--o
-1 . , , , , , ,
o 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. D.ll. Steam dome pressure--0.1-m2 MSLB from HZP. o.06 , , , , , , , 0.05- - o.o4 - - 9 v
'3~
g NOTE: These fransients j ossumed multiple operator / equipment 0.o2 - f ailures. See TABLE H. - S o.ol- - 0.00 - i
-o.oi , , , , , , ,
o 1000 2000 Jooo 4000 5000 6000 7000 soo0 Time (s) Fig. D.12. TBVs flow area fraction--0.1-m2 MSLB from ilZP. 262
0.0s , , , , , , , 0.04-
..........w.
1
, 0.02-5 0.00 3
-0.02 -
NOTE: These transients assumed multiple
-0.04 - operator / equipment f ailures. See TABLE I.
-0.06 , , , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. D.13. ADVs flow area fraction--0.1-m2 MSLB from HZP. 800 , , , , , , .
- -1600 700 -
- -1 00 600 -
- -1200 k 500--
? NOTE: These transients ._ ,oog n
assumed multiple , g 400- operotor / equipmen! - g l f ailures. See TABLE H. _soo , 300- ,
-600 200--
! ~ 100- . -200 0 , , , . , , , 0 0 200 400 600 800 1000 1200 1400 1600 l ! Tlrne (s) f~ Fig. D.14. Break mass flow--0.1-m 2 MSLB from IlZP. 263
l LOOOO , , , i i i i
-245000 100000- -
-210000
$ 80000-- --175000 6
3 k 60000-
~
~
"g N 3 -
-105000 E 40000- -
33 y NOTE: These fronsients -70 # 0 s assumed multiple j; 20000- operof or / equipment - f allures. See TABLE H. ~ 350" 0- -
--0
- 33000
-20000 . . , , , , .
0 200 400 600 800 1000 f200 1400 1600 Trne (s) Fig. D.15. Integrated break flow--0.1-m2 MSLB from HZP. l i-j 1 e 264
APPENDIX E 0.1-m2 MSLB FROM FP l l 1 i t i i l t 265
300000 -
, , , , , , . -660000
-630000 280000- -
-600000 260000-,
--570000
?* o Q
j -
-540000 y u0000- -
NOTE: The=e transients ossumed multiple -NM operator / equipment
,,, f ollures. See TABLE I. _
-480000
-450000 200000 , , , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. E.1. Primary system mass--0.1-m2 MSLB from FP. o.e , , , , , , , , , ,
~
0,40- *
/' wa s -
.......... intia 3
- THttA 3 o.as- - NA 4 -
THtta e THETA 4 o.so- - [ 0.25- -
" 0.20- -
"~ '
NOTE: These transients ossumed multiple ! 0.10 - operator / equipment - f ailures. See TABLE I. ; 4 o.os- - c.00-J -
-J.oS , , , , , , , , , ,
o 2so soo 7so 1000 teso e Uso 2000 2252 2500 250 Time (s) Fig. E.2. Voiding in the upper head--0.1-m 2 MSLB from FP. 266 m~
10 , , , , , , , , , .
-30 8- .
-25 6-
.-20
$ o o
-n y
}
4- . j b h' ~* h 2 . NOTE: These transients 5 ossumed multiple operator / equipment
, i f ailures. See TABLE I. ,_n
--5
-2 , , , , , , , , , ,
0 250 500 no 1000 1250 1500 950 2000 2250 2500 2750 Time (s) Fid . E.3. Pressurizer liquid level--0.1-m2 MSLB from FP. 5000 , , , ,, , , , THETA s -800
....... ..
- THETA 2 THETA 3 g 4000- ,
- THttA 4 THETA s
.-700 a C:
THcTA e
-600 s
} 3oon. .
-500 y% y 5 hl -s00 m 3
2000- l l. .' It . f N :' } 1
.:: -500 $
t l
'y t . 2
*g -
,i , NOTE: These transients .T assumed multiple i
1000- 'h"g' h*( l fP operator / equipment -
-200 'g -
i f ailures. See TABLE I.
., -10 0 0 , , , , , , , 0 0 1000 2000 3000 4000 5000 6000 7000 e000 Th= (s)
Fig. E.4. Downcomer heat-transfer coefficients at the core midplane 5--0.1-m 2 MSLB from FP. 267
800 . . . . . . . . . .
. -600
)
ago- -
- - 57s 140F A 8 --=
m- ..........., E
-s2s 340
- i,
-soo y y
sto- .
-as [
NOTE: These transients assumed multiple _a operator e pmen t
- f allures. ee BLE I.
~
. .e 400 , , , . . . . . . .
o 250 soo 700 1000 taso 1500 950 2000 2260 2000 2750 Time @ Fig. E.5. Hot-leg liquid temperature--0.1-m2 MSLB from FP. oh . . . , . . . . . . 0.a - o.7-0.s-9 v 0.s-02- -
~
NOTE: These transients assumed multiple c2- caerator e pmen t - _f silures. ee BLE H. o.t-0.0
-0.1 , . , , , . , , . ,
o 250 600 750 1000 1250 1s00 17so 2o00 2260 2soo 2750 Time @ Fig. E.6. PORV flow area fraction--0.1-m 2 MSLB from FP. 268
soo . , , , , , , , , , , _ ,3a 700- -
. -soo
..... ... imo, e
- 250 Soo- .
- 200 go- -
300-
-7so f
. NOTE: These transients _soo 200- assumed multiple -
operator equipmen t f ailures. ee TABLE I. , 0-
- N
-o
-too , , , , , , , , , ,
o a e a a a a a a e a a Time (=) Fig. E.7. Main feedwater flow--0.1-m2 MSLB from FP. 22.5 . , , , , , , , , ,
- y. g................................................................... .
, . -n tr.s- ,
3- Loo, a
.......... soo, . 3o 12.5 -
~
-20 f h 1s- -
h NOTE: These tronslenis , I s- . ossumed multiple -., l o erotor e i ment i 1 f ilures. ee LE I. - I 2.5 - i 0
'O
-2.5 , , , , , , , , , ,
o me 800 730 1000 uso 200 Uno 2000 2260 2500 2750 Time (s) Fig. E.8. Auxiliary feedwater flow--0.1-m 2 MSLB from FP. i 4 269
. . _ - , . , . . . ____,..__,,-_,.-..,-,.,,_._._.._,,.-_.-,_.._,,__.m.,.-__.,.__m___,_--m._.m.,__...__ < . , _ , . , . ~ . . . - . - ~ .
anoooo , , , , , , ,
-30000o LOOP A 2000cc- .......... goop e -
l
-dooooo I
scoco- - M -
-300000 $
b 100000-b
-200o00
'..*',,, NOTE: These fronsients ossumed multiple
.,- operator / equipment 3o000-f ailures. See TABLE H. -
-tococo i
o , , , , , , , o o 1000 2000 3000 400o 6000 5000 7000 8000 Time (s) Fig. E.9. Steam generator mass inventory--0.1-m2 MSLB from FP. soo - , , , , , , , , , , -yso l N- - LOOP A
--soo l . . . . .... . . too, .
I
-1250 3co- -
- -1000 o 400 -
I B d -750 g 300- - l
- -500 200- NOTE: These transients -
assumed multiple operator / equipment -250 mo- ~ f ailures. See TABLE H. - o-- * * * \ --o l - 10 0 , , , , , , , , , , o 200 000 750 1000 1230 1soo 1750 2000 2200 2s00 2750 Time (s) Fig. E.10. Steamline mass flow--0.1-m2 MSLB from FP. 270 t
. - - - - . . . - . - - . . _ , - , . - . . - - . , . . - . . . , . . - , , , - - - - - - - - . - - . . - . , - - - - - - . - - - - - - ~ . - - - - --
, 7 , , , , , , , i i ,
-soo s- LOOP A
.......... too, e
--7so 5- .,,
\
--soo 4-_ \.,, ,g.
a
. ...so ;
2
--300 2-- NOTE: These fronsients assumed multiple operator e pment 1-
- f altures, se BLE I. --mo 0- .
\ --o t
o 2so 600 750 do 1250 do 1No 20'00 22'50 2SD0 2750 Time (s) Fig. E.11. Steam dome pressure--0.1-m 2 MSLB from FP. t2 , , , , , , , , , , 1-NOTE: These transients o.s- ossumed multiple - e operator e pment U f ailures. ee BLE I. o.s - o.4 - I o.2 - I o-
=ol , , , , , , , , , ,
o S to 5 20 25 30 36 40 46 50 55 Time (s) Fig. E.12. TBVs flow area fraction--0.1-m2 MSLB from FP. 271
t.2 , , , , , , , , , , l-wa .
.....-... goop a 0.s - -
I 9 v 0.s - - NOTE: These fronsients ossumed multiple operator e pment o.g. f allures. ee BLE I. . 1 [ 0- - f
-0.2 . , , , , , , , , ,
0 ISO 800 750 1000 1250 1500 175 0 2000 2250 2000 2750 i Time (s) Fig. E.13. ADVs flow area fraction--0.1-m2 MSLB from FP. 1000 , , , , , , , i
- -2000 s00- . --050
-1500 4
s00-
-1250 R
-1000 t!
* ' ~
NOTE: These fransients i assumed multiple -750 l l operator e pmen t i f allures. ee BLE I.
-500 3o0
- -250 0 , , , , , , i 0 0 00 100 10 0 200 250 300 350 400 Time (s)
Fig. E.14. Break mass flow--0.1-m2 MSLB from FP. 272 i
s0000 , , , . . . i e0000- ,
-175000 70000- ,
- 50000 M~
. -125000 b, d
~
50000-
. -100000 .
~
40000-
. -7s000
~
30000-
-50000 20000-NOTE: These transients ~-25000 10000-, ossumed multiple operator / equipment
- n. ,
f ailures. See TABLE H. .
.o -10000 , , , , , ,
350 400 0 50 100 20 200 250 300 Time (s) Fig. E.15. Integrated break flow--0.1-m 2 MSLB from FP. l 273
I i l APPENDIX F 0.1-m2 MSLB WITH TWO OPERATING RCPS FROM HZP i I I 1 l l 274
30C000 , ,-J50000
' 050CC3 280000-L
- -600000 i
2G0000- -570000 r= 2 o 6 N U g - suo0a y 240000-NOTE: These transients 3ionno ossumed multiple operator / equipment 220000- f ailures. See TABLE I. -
, -48000.1
- 450000 200000 , , , . ,
0 1000 2000 3000 4030 5000 6000 7000 8000 Timo (s) Fig. F.1. Primary system mass--0.1-m2MSLB with two operating RCPs from HZP. t 8.0 f , , , , , , , , , , I"E'" ' - 7.0 - .. .. . . .. . THETA 2 THC'A 3 l - THCA 4 6.0 - THUA $ THCTA 6 NOTE: These fransients
~
50-8 assumed multiple j ,,o _ operator / equipment _ y f ailures. See TABLE I. u 3.0 - l 2.0 - to-I o.0 --
- k. ' " ^ - - - - - - - -
-10 , , , , , , , , , ,
O S00 1000 500 2000 2500 3000 3500 4000 4500 5000 5500 Time (s) Fig. F.2. 2 Voiding in the upper head--0.1-m MSLB with two operating RCPs from HZP. 275
to . . . . . . . . . . 33 s- - 25
- s. -
. 20 E E l
} -
- 15 t l 3 s -
$ b E
~
NOTE: These tronsients h 2-assumed multiple
/ operator equipment -
t f ailures. ee TABLE I. ~5 0- -
-O
-s
-2 . . . . . . . . . ,
0 Ef.C 1000 1 BOO 2000 2'00 J000 3t00 4000 4l,00 5000 5500 Time (s) Fig. F.3. Pressurizer liquid level--0.1-m2 MSLB with two operating RCPs from HZP. MMC . . . . . . . ws . -s000
.. ....... . ws :
wma p 40000- - _ g --7000 .;. ws.
*S000 E
f30000- s
~****
i, NOTE: These fronsients 6 k ossumed multiple y _ operator e nt .m ,, L f o llures, e LE I. 20000- - i ( l I y w
-3000
}
i -
-2000 10000- . -
l l -
-1000 i 0 . , , , . . . e 0 1000 2 00 J000 4000 8000 6000 7000 e000 Time (s)
Fig. F.4. Downcomer heat-transfer coefficients at the core midplane 5--0.1-m 2 MSLB with two operating RCPs from HZP. l , . _ _ _ . . -_-- - - - - - - -- - - - - - - - - - - - -
560 . . . . . . . .
-525 540- -
-500 LOOP A l
8 520--! . . . . . . . . . . wo, . -- 475 8 i 2 - 3' 2 { o c -450 }* 500- - h NOTE: These ironsienis -425 h assumed multiple operator / equipment 5 1 4a0-, f allures. See TABLE H. -_4oo 6
-375 460- -
_ 330 440 , , , , . , , , , . 0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000 5500 Time (s) Fig. F.5. Hot-leg liquid temperature--0.1-m2 MSLB with two operating RCPs from HZP. 0.7 , , , , , , a i i i 0.6 - - 0.5 - - 0 *ni
- E $ 0.4 - *;;;
c
- g[w -
4 o G-5 a.s - ~5*. -
'. EE 3 I 0.2 - n.od. -
w,c'
.. E , 3 Hwe=
omaa ! 0.1 - Zoow - l J- t o.O - l
-0.1 . . , , , , . . . .
0 600 1000 1600 2000 2600 J000 3600 4000 4400 6000 6600 i Time (s) Fig. F.6. PORV flow area f raction--0.1-m2 MSLB with two operating RCPs from HZP. 277
35 , , , , . i 70 30- . Loopa
. . .. .. . . . . Woe O -60 25- -
-50 20- -
6
-s0 $
e g 15 - . 8 NOTE: These transients -30 C 8 assumed multiple 3 2 g_- operator / equipment j f ailures. See TABLE H. --20 5-. -_ ,n o-- o
-5 , , , , , , ,
- -10 o as so a too as no 1m soo Time (s)
Fig. F.7. Main feedwater flow--0.1-m2 MSLB with two operating RCPs from HZP. US . . , . . , , , , . 20- .----------------------- ---------------- - ---------------------.
~
17.5 - -
-40 15 - Loopa -
i 9 _: .. .. . . . . . . woe , 33 N cn 12.5 - - l 6 i o i , ; y 10 - j - g t n -! -20 1 75-
- NOTE: These transients
! ossumed multiple -
- operator / equipment l 5-.i f oilures. See TABLE I. -- ,a l I
- I 2.5 - -
l 1 0- -
-0
-2.5 . . . . . . . . . .
0 500 1000 1500 2000 2$00 3000 3500 4000 4500 5000 5500 l Time (s) , Fig. F.8. l Auxiliary feedwater flow--0.1-m2 MSLB with two operating RCPs from HZP. 278
250000 . . . . . . . l
-500000 l
200000- .
- ,,. 4o0000 I
'..**..- LoopA N 000- ,.' -
m ,,,,,,,,,,L,,,, y . ,,.',, 300000 h g 1 m 000
-200000 NOTE: These transients assumed multiple _
- 00~ -
operofor equipmen1 -100000 f ailures. ee TABLE U. 0 , . . . . . , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. F.9. Steam . generator mass inventory--0.1-m2 MSLB with two operating RCPs from HZP. > 500 , , , , . . . 4 i i l - -1050 I ' 450 - LOOP A
,-goo 400- . . . . . . . . . . too, .
350- - 750 t ~ Q 300- l m -600
- 2. -
c, y 250- y o o C ~43o C g 200-' p E
- E 15 0 -, -300 NOTE: These transients l 10 0 - ossumed multiple -
operotor e pmen t _ iso 50 I f ailures. ee BLE I. -
' --0 o_:......t
-50 , . . > > > . . i -
0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000 5500 Time (s) l Fig. F.10. l Steamline mass flow--0.1-m 2 MSLB with two operating RCPs from HZP. l 279
I 7 , , , , , , , , , ,
-900 -
6- - LOOP A
.. . . .... .. Logy g
~ ** l 5- -
'~., l
\..---.,' . . ',,,. j
-600 o -
i NOTE: These transients e kg : assumed multiple operator / equipment O
'a
-- '5 a
3-i f ailures. See TABLE I. E
- i. p
{ I, -300 n. 2- -
~
' ' .................--- - **.... ,,.. ,so
~. _
0-- --O
-t . . , , . , . . , ,
0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000 5500 Time (s) Fig. F.11. Steam dome pressure--0.1-m2 MSLB with two operating RCPs from HZP. 0.0e , , , , , , , 0.07- - 0.06- -
.n 0.05- -
U O 0.04 -- - 2 4 I 0.03- - 8 0.02-NOTE: These transients ~ ossumed multiple operator / equipment g,g , _ f allures. See TABLE H. _ 0.00- -
- 0.01 . , , , , , ,
0 10 20 30 40 50 60 70 80 Tme (s) Fig. F.12. TBVs flow area fraction--0.1-m2 MSLB with two operating RCPs from HZP. 280
- 0.06 , , , , , , , , , ,
0.04- -
. . . . . . . .. . wop a y 0.02- -
D o k 0.00 - e 8
-0.02 - -
NOTE: These transients _o,o4_ ossumed multiple - operator / equipment f ailures. See TABLE H.
-0.06 . , , , , , , , , ,
o soo 1o00 moo acoo 2600 300o 3600 4000 4600 sooo 6600 Tavie (s) Fig. F.13. ADVs flow area fraction--0.1-m2 MSLB with two operating RCPs from HZP. 800 , , , .
- -1600 700-
- -1400 600-
' - -1200 500- - d!, - -1000 o B 3 g 400- - g
-800 g NOTE: These transients 2 assumed multiple ~
j 300~ operator / equipment _,on f ailures. See TABLE H. 200- -400 WO- --200 0 , , , , 0 0 10 0 200 300 400 500 600 Time (s) l Fig. F.14. Break mass flow--0.1-m2 MSLB with two operating T. cps (from HZP). 281
140000 , , , , 120000- - -270000 100000-
~
- -225000
^
m 6 @ 3 s0000- -
, -180000 g O E M
60000- -
-135000 l -
--90000 I
{ I40000- i
~
NOTE: These transients ossumed multiple 0 _ operotor equipmenf ._0 f ailures. ee TABLE I.
-20000 , , , ,
0 10 0 200 300 400 500 600 Time (s) Fig. F.15. Integrated break flow--0.1-m2 MSLB with two operating RCPs from HZP. 282
i APPENDIX G DOUBLE-ENDED MSLB WITH UNIS0 LATED AFW TO BROKEN SG FROM HZP l l t l l i 283
300000 - , , , , , . -560000 i
- -630000 280000-
- o0000a 260000- . -570000 o
3 g -
-540000 $
- m. .
- -510000 NOTE: These tronslents
,,,,_ ossumed multiple .
- o erotor e -4s0000 f ilures. ee LE I.
- -450000 200000- . . . i i , .
0 1000 2000 3000 4000 5000 6000 7000 s000 Trne (s) Fig. G.I. Primary system mass--double-ended MSLB with unisolated AFW to broken SG from HZP. CAO , , , , , , IMU"1 0.55- . ..... .... THCA 2 THCA 3
- THUA 4 0.30- --
THcAs - THUA 4 0.25- - J OJ0- - 6: NOTE: These transients 4 15 - ossumed multiple i operator equipmen t '
""- f ailures, ee TABLE E.
l 0.05- - 0.00-
-0.0s , , , , , , ,
O S00 1000 1600 2000 2500 3000 3500 Time (s) Fig. G.2. Voiding in the upper head--double-ended MSLB with isolated AFW to broken SG from HZP. 284
-.-+
15 , , , , i , 12.5 -, -,, 30 -
-3o 1 73- -
j >
-20 .3 s- -
. _,o 2.5 - -
NOTE: These fronsients assumed multiple operator / equipment 0- f ailures. See TABLE H. --o
-2.5 , , , , , ,
0 500 1000 1500 2000 2500 3000 3500 nme (s) Fig. G.3. Pressurizer liquid level--double-ended MSLB with unisolated AFW to broken SG from HZP. l l l 20000 l l e000-l g a000-a000 NOTE: These fransients assumed multiple g_ operator / equipment f ailures. See TABLE H. p 4000-
@ 3000-h --
m. c , , , . , , , o acoo sooo sooo 4000 acoo sooo sooo ecco 11E M Fig. G.4. Downcomer heat-transfer coefficients at the core midplane-- double-ended MSLB with isolated AFW to broken SG from HZP. 285 i i
M . , , , , , 560 550- _
-520
, -480 b .
. .. ... ... . . wo, e C
S00- '. --440 s 2 m.- . -400 g y E s2 NOTE: These ironslent: .
-360
,3 assumed multiple N. operator / equipment
-m d
f ailures. See TABLE H.
\.3 m- t
-280 400 .
. , , , . . . - _74o M , , , , . ,
O S00 1000 1500 2000 2500 3C00 3500 Time (s) Fig. G.5. Hot-leg liquid temperature--double-ended MSLB with unisolated AFW to broken SG from HZP. 0.06 . , , , , , 0.05- - 0.04- - n U
, 0.03- -
k I 0.02- _ 3 NOTE: These transients assumed multiple operator / equipment f ailures. See TABLE H. 0 01- ,
' L -
0.00
-0.01 , , , , i ,
0 500 1000 1500 2000 2500 3C00 35 0 Time (s) Fig. G.6. PORV flow area fraction--double-ended MSLB with unisolated AFW to broken SG from HZP. t 286
0.06 , , , , , , 0.132277 WO' ' - 0.04-
- .. . ... . . mop e - 0.079366 In this transient, the main feedwater was set to 0.0 kg/s initially instead of being ramped down from 10.0 kg/s.
0.02- .
- - 0.026455 6 o Q Q.00 g
--0.02s455j
-0.02 -
NOTE: These transients ossumed multiple __g,n7,3,, operator / equipment _o.04 f oilures. See TABLE H.
-0.06 , , . .
, , -0.132277 0 500 1000 1500 2000 2500 3000 3500 Time (s)
Fig. G.7. Main feedwater flow--double-ended MSLB with unisolated AFW to broken SG from HZP. 22.5 , , , , , , l 20
- i.o 17.5 -
15 - - tooP A -
..........wo,. 33 u- -
Q u
* ~ $ _ 20 6 7.5 -
5-_ -_ ,g NOTE: These transients u- ossumed multiple _ l operator / equipment f ailures. See TABLE H. 0-- --S
. 2.5 , , , , , ,
0 800 1000 1500 2000 2500 3000 2500 Time (s) Fig. G.8. Auxiliary feedwater flow--double-ended MSLB with unisolated j AFW to broken SG from HZP. l 287 i
250000 . . . . . . . 500000 LOOP A 200000-- -- -------- LDOP O -
-400000 150000--
.~.
,' 300000 [
n ,,,, , g 8 ..,.- 2 s 100000-
-200000
[ NOTE: These transients assumed multiple 50000 operator / equipment _
~
f ailures. See TABLE I. , 100000 0 . . . . . . . 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Tme (s) Fig. G.9. Steam generator mass inventory--double-ended MSLB with unisolated AFW to broken SG from HZP. e00 -
, , , , , , 37.,,
700- -
'530 wCP a
.. ... . .. . - wor e ,
-1250 500- g 400- -
- toco g
o 1 d L -750 g 300-. - G : f 2 h NOTE: These transients 500 200- ossumed multiple operator / equipment f ailures. See TABLE I. _-250 0-L* ---- -- --- - ------- - - --- --
. - - = --3
-100 , , , . . .
O S00 1000 1600 2000 2500 3C00 3500 Time (s) Fig. G.10. Steamline mass flow--double-ended MSLB with unisolated AFW to broken SG from HZP. 288
7 , , , , , ,
- -900 6-woe .
..........wo,,
s- ...... ~'** '
.. _-m
....... . - ****_,,,.. {
4-1 ~300 c i y
- .... *\ &
- 4so e s-
- 9 5
I.
' "~"'~"" --* E 2'
NOTE: These transients assumed multiple operator / equipment -"so 8-~ f ailures. See TABLE I. 0-- 3
-1 . , , , , ,
0 500 1000 1600 2000 2500 3c00 36c0 Time (s) Fig. G.11. Steam dome pressure--double-ended MSLB with unisolated AFW to broken SG from HZP. 0.06 , , , , , , 0.05- - 0.04-m t$ 0.03-0 k i 0.02- NOTE: These transients - A assumed multiple operator / equipment f ailures. See T ABLE H. - 0.01-0.00
-0.01 , , , , , , .
0 500 1000 1600 2000 2500 3000 3500 Time (s) Fig. G.12. TBVs flow area fraction--double-ended MSLB with unisolated AFW to broken SG from HZP. 289
' o.06 , , , , , , ;
I gu. wo* .
..........wo,.
m o.02- -
)
5 ' O d o.00 - 5
-o.o2 - -
NOTE: These transients ossumed multiple
-o.o4 - operotor / equipmen1 -
f ailures. See TABLE H.
-0.os , , , , , , .
o 500 1000 1500 2000 2500 3000 350o Time (s) Fig. G.13. ADVs flow area fraction--double-ended MSLB with unisolated AFW to broken SG from HZP. teco , , , , , , ,
-3200 1400- -
-2800 200- -
-2400 k
cn 1ooo- - N 6 -2000 S,, g 800- - g NOTE: These transients N h" gon_ ossumed mumple - 2 operator / equipment -noo f ailures. See TABLE H. l O -
-300 1
200- . 4oo o , , , , ~, , , o o 200 400 soo soo 1000 uoo 1400 1600 Time (s) Fig. G.14. Break mass flow--double-ended MSLB with unisolated AFW to broken SG from HZP. 290
i t MMO , s s , , s s
- , -270000 120o00
--225o00 2 000--
Si 6
-iaono y 3
l 3 a0000- - - j E E a a j eoooo_- --us000 j 3 4000o-- --90000 NOTE: These transients h = ossumed multiple 20000- operotor / equipment --45000 f ailures. See TABLE I. 0-- --o
-20000 , , , , , , ,
0 200 400 600 800 1000 1200 M00 1600 Time (s) Fig. G.15. Integrated break flow--double-ended MSLB with unisolated AFW to broken SG from HZP. l l l 291
- - ,,,., , , ------,,-m.- - -
.,-,---,.,.-e.,w,,,.,--m, ,y,w-y, mr,, , , ,, -r- - - ,--
l APPENDIX H DOUBLE-ENDED MSLB WITH TWO STUCK-OPEN MSIVS FROM FP l l 292
320000 , . . . . .
- -690000 300000-- -
-660000
-630000 280000- -
- -600000 9 7 o
6 n 260000- ,
--570000 E h S
-540000
## ~
NOTE: These transients assumed multiple - 5t0000 operator e pment g ogo, f ailures. ee BLE I. _
-480000
- -450000 200000 . , , , , . ,
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. H.l. Primary system mass--double-ended MSLB with two stuck-open MSIVs from HZP. 12 , . i i . . l THETA 1
. .. . . . .... THETA 2 ,
. . . TMCTA 3
- TH('a 4
- THUA 5 THETA 6 I
0.s - 5 - 0.6 - LC 6 f 0.4-NOTE: These transients assumed multiple . 0.2 - operator e pment - ! f ailures. ee ABLE H. l 0-L - l
-0.2 , , . . . .
0 500 1000 1500 2000 2500 3000 3500 1 Time (s) Fig. H.2. Voiding in the upper head--double-ended MSLB with two stuck-open MSIVs from HZP. 293
10 , , , , , ,
- -30 8- .
- -25 6-- .-20
$, o
=
- 13 1
.5 4- .
3 5 -10 b g - g 2- - NOTE: These transients assumed multiple operator equipment ~~0 o. f ailures. ee TABLE B.
- --5
-L , , , , , ,
0 500 1000 1500 2000 2500 3000 3500 Time (s) Fig. H.3. Pressurizer liquid level--double-ended MSLB with two stuck-open MSIVs from HZP. 5000 , , , THETA 1 -300
. . . . . . . . . . i HL TA 7
- NETA 3 4000- - NCTA 4 -
-MO a.
7 - - _. NtTA 5 g-mcTA. E 600 :- e,
; g g
3000-8 h { NOTE: These transients t E y assumed multiple -400 }} 2000-operator e ont ~ 8 f ailures. ee LE I. [P
~
[. -300 g x '
.f _
%- ~* 3l 1000- .f -
a
-10 0 0- , , . . , . -3 0 1000 2000 3000 4000 5000 tiOOO 7000 8000 Time (s)
Fig. H.4. Downcomer heat-transfer coef ficients at the core midplane-- double-ended MSLB with two stuck-open MSIVs from HZP. l 294 i
s75 . . . . . .
- -540
! 550- - I - 495 525-8 C
~
500-
. . . . . . . . . . wo, , y
-405 ,
m_ _ ' - NOTE: These transients _3so 4 *- ossumed multiple - I operator / e uipment 3 _t
. P -
f ailures. See ABLE I. j se g 425-
- - 270 ago_ .
375-
- 18 0 350 , , , , , ,
0 500 1000 1500 2000 2500 3000 3500 Time (s) Fig. H.5. Hot-leg liquid temperature--double-ended MSLB with two stuck-open MSIVs from HZP. 05 . . . . . . 0.8 - - l 0.7 - - l 0.6 - - l 9 0.s - - 0.4 - - 0.5 - - NOTE: These transients 0.2 - ossumed multiple - operator / equipment [ o,, _ f allures. See TABLE H. _ O.0 -..m - - -
-0.1 , , , , , ,
0 500 1000 1500 2000 2500 3000 3500 Time (s) Fig. H.6. PORV flow area fraction--double-ended MSLB with two stuck-open MSIVs from HZP. 295 i
r l 33 , , , , , , ,
-70 ,
1 30- -
.. . . .. . . .. wo, s -60 l
25- -
-50 f m- -
d!, -
-40 f
15 - -
-30
=
j NOTE: These transients 8 S-. ossumed multiple --20 2 operator / equipment f ailures. See TABLE H. 5-_ __g 0-. o
-5
---to 0 26 50 75 100 12 5 16 0 f75 200 Time (s)
Fig. H.7. Main feedwater flow--double-ended MSLB with two stuck-open MSIVs from HZP. zz.s , , , , , , l 20- -
. 4o 17.5 - -
15 - LOOP A
.. . .... . . . wo, e -30 9 _
12.s- - a y 10 - -
-20 $
g j 7.5 - - 5" NOTE: These transients ossumed multiple 5-- operator / equipment -. ,0 l f allures. See TABLE I.
- l. 2.s -
0- -
--O i
-23 , , , , , ,
0 500 1000 1500 2000 2500 3000 3500 Time (s) Fig. H.8. Auxiliary feedwater flow--double-ended MSLB with two stuck-open MSIVs from HZP. 296
250000 , , , , , , .
-500000 l
LDOP A 200000- ---------- loops -
-400000 150000- -
-300000 [
O O 2 2 100000- -
-200000 NOTE: These transients ossumed multiple operotor / equipment f ailures. See TABLE H. -
0-
-100000
~. _
O , , , , , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. H.9. Steam generator mass inventory-double-ended MSLB with two stuck-open MSIVs from HZP. 4000 , , , , , , , , , 3500- _ LOOD &
---------- Loor s
-6000 2500- -
=
6 2000- --4500 3 o I D M 1500- - n f
-30%
1000- NOTE: These fransients - l ossumed multiple i operator / equipment _ ,3, 500- f ailures. See TABLE H.
~
0-- 0
-500 , , , , , , , , ,
0 10 0 200 300 400 500 600 700 800 900 1000 Time (s) Fig. H.10. Steamline mass flow--double-ended MSLB with two stuck-open MSIVs from HZP. 297
7 , , , , , , , , , 6- -900 LDop a
.......... LOOPO 5-
~
~*
p 4- --600 9 A i S M 3-
- - 450 2
3 i 2--- -~300 NOTE: These transients assumed multiple I-operator / equipment f ailures. See TABLE L -- 85 0 o-- -.o
-t , , , . , , , , ,
o too 200 soo 400 soo soo 700 soo soo sooo Time (s) Fig. H.11. Steam dome pressure--double-ended MSLB with two stuck-open MSIVs from HZP. Om , , , , , , 0.06- - 0.05- - e U 0.04- _ O 0.0 3 -. - e 8 0.02- - NOTE: These transients assumed muffiple 0'0 ' ~ operator / equipment f ailures. See TABLE L
~
o.co -
-0.01 , , , , . ,
0 600 1000 1600 2000 2500 3000 3500 Time (s) l Fig. H.12. i TBVs flow area fraction--double-ended MSLB with two stuck-open MSIVs from HZP. 298 l l l i
l am , , , , , , 0.04- - j . .. . . . . . . . too, . 0.02- - l 0 0.00 - 9
-0.02 - -
NOTE: These transients assumed multiple operator / equipment
-0.04 - f ailures. See TABLE I. -
-0.06 , , , , , ,
o soo 10 o 0 soo 2000 2600 30o0 asco Time (s) Fig. H.13. ADVs flow area fraction--double-ended MSLB with two stuck-open MSlVs from HZP. 5MO , , , , , , ,
-10000 4000- -
-8000 3000- -
-6000 l 6 o l D y y 2000- -
-4000 o
NOTE: These fransients j ossumed multiple operator / equipment S" 1000-_ f ailures. See TABLE H. --2000 I l 0-- h -0
-1000 - , , , , , , ,
--2000 0 200 400 600 800 1000 1200 1400 1600 Time (s)
Fig. H.14. Break mass flow--double-ended MSLB with two stuck-open MSlVs from HZP.
, 299
2s0000 . . . . . . i + >
-500000 200000- - ::
-400000 1s0000- [
-300000-g g20000- .
J
-200000 so000- NOTE: These transients . .
ossumed multiple - SNO s
}
operator / equipment f allures. See TABLE I. 0-- --0
- --'
- 0'
-s0000 . . . . . . . . .
O 10 0 200 300 400 500 000 700 800 900 1000 Time (s) Fig. H.15. Integrated break flow-double-ended MSLB with two stuck-open MSIVs from HZP. l l 1 I l 300
i I f i APPENDIX I ONE STUCK-OPEN TBV FROM FP f l I i I t 301
7. 30C000 , , , . -660000
-J3000')
2a0000- -
-30000.)
26C000- - 90000 cn 2 6 o uc001 - 240000- - 510000 220000 NOTE: These transients _ ossumed multiple 4soona operator / equipment f allures. See TABLE E. 450000 200000 . . . . . . . . 0 1000 2000 3000 4000 b000 6000 7000 2000 Time (s) Fig. 1.1. Primary system mass--stuck-open TBV from FP. 0.0s , , , , , , , , , , , THETA 1
. .. . ...... THtfA 2 0.04 -- THETA 3 -
- THETA 4 THETA S THETA 4 0.02- -
8 0.00 -
-0.02 - -
NOTE: These transients i ossumed multiple
-0.04 - operator / equipment -
t e,ilures. See TABLE E.
-0.Os . . . . . . . . . . .
0 500 1000 1500 2000 2500 3000 3500 4000 4600 5000 5500 6000 . Time (s) I Fig. 1.2. j Voiding in the upper head--stuck-open TBV from FP. l [ l 302 i i
10 . . . . . . . . . . . g.
. -30 8- -
-2s 7- -
l
$ 6-- . -20 g 1 i b 5- -
3 g -
-s ) 4- -
3- NOTE: These transients - '" assumed multiple 2 operator / equipment . f ollures. See TABLE I. 3 i. 0 . . . . . . . . . . . 0 0 500 1000 200 2000 2500 3000 3600 4000 4600 5000 5600 8000 Time (s) Fig. I.3. Pressurizer liquid level--stuck-open TBV from FP. 5000 , , , , , , , ! Ncta i -800
. . . . . . . . . . NOA 2
! -- NCTA 3 4000- - NETA 4
- l 9 -
NETA $ M e 70o . C
- "E mcT*
- g C l k= .
- E'j .soo h e
- . .=2 o
8 3000-a 81b - O
-s.. -sco j i
e jM - a* y - C n $ wi -400 O 2000- '.. ~~\-'.... " f**b --
...," f....---,,'- y 3 g :3
>,e= -300 o
m s" ~~.'*"* ;; OMaO f g .; ZOO
- 1000-
-b \ ,, p -- -
-200 h
'3
*( ..
0 , , , , , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Trne (s) l Fig. 1.4. j Downcomer heat-transfer coefficients at the core midplane-- stuck-open TBV from FP. 303 l
590 , , i i i + ' ' '
-600 585- _
-590 Sao- NOTE: These ironsienis -
ossumed multiple -5a0 m- OPerotor / equipment 2 f ailures. See TABLE II. .- -- go, , _ C
, - toop e
- 570 D 510- -
E e - 2 565- -560 f - g_ _-550 4
.P 3 555- --540 550--
--530 545-_ ,
--520 540 , , , , , , , , , , ,
0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000 5500 6000 T'me (s) Fig. 1.5. Hot-leg liquid temperature--stuck-open TBV from FP. tz , , , , , , , , , , , g-OC Eos:i 0.s -
- E Wj -
ye&q Q E a-0.6 - * - O 5 *[NS E I 7,DO*E *
) 0.4 -
w
- .=
a u 3 o e o. o Z 0 O *= o.2 - - 0 -- - - l -02 . , , , , , , , , , , i o too woo moo 200o 26oo sooo 3600 4000 4600 60 % 660o sooo Time (s) Fig. I.6. PORV flow area fraction--stuck-open TBV from FP. 304 l
800 , , , , , . ,
-1600 700 -
Lo0P A
.. .. ...... Loor e -1400 600- -
j -1200 T 500- - Q
$ -1000 e, 3
- g 400- -
y
~"
a NOTE: These ironslenis g j 300-ossumed multiple - j operotor / e pmen t _soo f ailures. See ABLE I. 200- ~
-400 10 0 -_ --200 0 , , , , ,' , , 0 0 10 0 200 300 400 SCO 600 700 800 Time (s)
Fig. 1.7. Main feedwater flow--stuck-open TBV from FP. 22.5 , , , , , , , , , , , 20- -
-40 17.5- -
15 - Loop A -
. . ...... --- Loor e -30 12.5 -
o 3 l g 10 - - C n
- 20 M
e j 7.5 - . j NOTE: These transient s 8~. ossumed multiple -_io operator equ i pmen t 2.5 - f ailures. ee TABLE H. - l 0 --O
-15 , , , , , , , , , , ,
0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000 5500 6000 Time (s) Fig. I.8. Auxiliary feedwater flow--stuck-open TBV from FP. 305
I NM i . . . . . . ;
-500000 LOOP A 200000- - toop a -
-400000 150000- -
)
-300000 @
fg_ NOTE: These transients cssumed multiple T 8 2 operator / equipment 2 20000- f ailures. See TABLE H. -
-200000 50000-, -
-100000 D i . . . . . . 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s)
Fig. I.9. Steam generator mass inventory--stuck-open TBV from FP. 800 -
. . . . . . . . . . i -1750 700 -, -
-1500 (DOP A
.. ........ tooP .
-1250 500- -
-1000 dN-f NOTE: These transients 7g g . ossumed multiple _ 73o operator / equipment 3"~ ~
1 f ailures. See TABLE I. '
-500 200. -
l l
~
10 0 -
~
0-- --O
-WO , . . . . . . . . i .
0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000 5500 6000 Time (s) Fig. 1.10. Steamline mass flow--stuck-open TBV from FP. I 306
a.50 , , , , , , , , , , ,
- _ - - - - _. .920
. s.25- % -
6-LOOP A
---------- Loop g 5.75- ,1
--840 9
a S [I 3.3o.
-. -- am e 5 -760 E
E25- - NOTE: These transients '. 5-, j ossumed multiple \ -_7,o l o erotor e pmen t i f ilures. ee ABLE H. \ 4.75- \- _ -680 4.50 , , , , , , , , , , , 0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000 5500 6000 Tune (s) Fig. 1.11. Steam dome pressure--stuck-open TBV from FP. 1.2 , , , , , , ,
- g. .
0.8 - i 9 v - g 0.6 - k 0.4 - I 0.2 - t i NOTE: These transients assumed multiple 0- operator equipmen t - f ailures. ee TABLE I.
-0.2 , , , , , ,
0 250 600 750 1000 1250 200 1750 2000 Time (s) Fig. I.12. TBVs flow area fraction--stuck-open TBV from FP. 307
4
, 6 4 , , , , g g g 4
I- toop a -
. . . .. . . . . . LDoP S 0.8 -- -
l 9 v NOTE: These fronsients 06- assumed mulliple - n E operator / equipment 4 f ailures. See TABLE H. 0.4 - - 9 0.2 - 0- -
-0.2 , , , , , , , , , , ,
o soo 1000 1600 sooo 26oo Jooo 3600 4000 4600 sooo 6600 sooo Tme (s) Fig. I.13. ADVs flow area fraction--stuck-open TBV from FP. 800 , , , , , ,
-1600 700- -
600- -
-1200 k h I -
7 N g, NOTE: These transients .ioco g y assumed multiple , o 400- operator e pmen t - o 5n f ailures. ee ABLE I. -800 5m S 300 - S
-600 200-- _
--400 10 0 - *-200 t
l 0 0 200 300 400 500 k' 600 i 700 0 10 0 Tme (s) Fig. 1.14. Break mass flow--stuck-open TBV from FP. 308
mmo , , , . . .
- -ssoooo soooo-
. -120000
^ soooo- -
. -100000 ,
e, 4 coco-
- e,
. -80000 soooo-
. -60000 A #
20000- NOTE: These transients - ossumed multiple operator / equipment f ailures. See TABLE H. - 10000- - -20000 o ido 250 3do 4bo sdo 650 700 Time (s) Fig. 1.15. Integrated break flow--stuck-open TBV from FP. I 309 i
}
su h l APPENDIX J ONE STUCK-OPEN TBV WITH ONE STUCK-OPEN MSIV FROM FP 310 ) l I s.
r-300000 -
. . . . . . i -660000
- -630000 280000- -
. -600000 260000- , -570000 9 9 m m 0 _
-540000 j 2 -
240000-NOTE: These transients -510u0 assumed multiple operator equipment ~ 220 u0-_ f ailures. ee TABLE H. -480000
- -450000 200000 . . . i i . i 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s)
Fig. J.l. Primary system mass--stuck-open TBV with stuck-open MSIV from FP. 0.06 . . . . . . . . . THETA 9
.......... THETA 2 0,04 THETA 3
- THETA 4 THETA 8 THctA e 0.02- -
.5
}
0.00
' ~
NOTE: These transients ossumed multiple operator equipment
,0, _ f ailures. ee TABLE I. _
-0.06 . . . . . . . . .
0 250 000 750 1000 1250 1500 1750 2000 2250 2500 Time (s) Fig. J.2. Voiding in the upper head--stuck-open TBV with stuck-open MSIV from FP. 311
6 , , , , , , , , , 5- -
-5 4- -
p v NOTE: These transients ossumed multiple 8 3_ ,.,a 4 operator / equipment s 3 f ollures. See TABLE I. j 2- -
-5 1
0-- - o
-l , , , , , , , , ,
o 250 500 750 1000 1250 1500 175 0 2000 225o 2500 Time (s) Fig. J.3. Pressurizer liquid level--stuck-open TBV with stuck-open MSIV from FP. 5000 , , , , , , . um i -Boo
..... . .... mw a MTA 3 9
4000-- - MTy --700 -f
-m.
b
-soo g
3000-k h\ -
-500
\
k' b $ l s ! 2mo- i NOTE: These transients
~
*l j
l - f ossumed multiple -300
- [;'
operator / equipment T ' ' f ollures. See TABLE I. .T
,g-
_i icoo. { %! ^I
-200 ,t -
_i Li' . o , , , , , , , o o 1000 2000 3000 4000 5000 sooo 7000 8000 Time (s) Fig. J.4. Downcomer heat-transfer coefficients at the core midplane-- stuck-open TBV with stuck-open MSIV from FP. 312
se0 _ . . . . . . . . .
-s00
$go.
- -580 syn _ -
LOOP A -500 8 C
........u,,,,
u 500-
- -540 f.#
"~ NOTE: These transients ~
assumed multiple operator e ment _m jlh T
. f ailures. ee LE I. y .y *~ ~
3.
. , -500
$30_ N- .,
s20 -
-480 510
~
~#
0 250 500 750 1000 1250 1500 f750 2000 2250 2500 Time (s) Fig. J.5. Hot-leg liquid temperature--stuck-open TBV with stuck-open MSIV from FP. EM . . . . . . . . . 0.04- -
, 0.02 -
i 0 0.00 l J NOTE: These transients assumed multiple operator equipmen t 4,, , f allures. ee TABLE H. _
-0.04 . . . . . . . . .
0 250 500 7bo 1000 1250 1600 1750 2000 2250 2500 Time (s) Fig. J.6. PORV flow area fraction--stuck-open TBV with stuck-open MSIV from FP. 313
soo , , , , , , ,
-tooo yoo. .
_ LOOP a
. .... ...... ino, o
_m so0- -
-1200 Qa m- -
O -1000 3 g' soo- - a g-
. -soo .
- m. .
h NOTE: These transients -soo assumed multiple ' 200 operator equipmen t . f ailures. ee TABLE H. -400 10 0 . -
-200 0 , , , , , '. , o
.o too 200 soo 400 500 soo 700 soo Time (s)
Fig. J.7. Main feedwater flow--stuck-open TBV with stuck-open MSIV from FP. 22 3 , , , , , , , , , 20-
-4o 17.5 - -
is- uwa - o . .......... inor 8 -30 12.5 - - 3 s 0 *~ ~
-20 E 7#~ ~
NOTE: These transients ossumed multiple 5-. operator e pmen t -.,a f o llures, se BLE I. 2.5 - o o
-2.5 , , , , , , , ,
4 0 250 500 750 1000 1280 1800 1750 2000 22's0 2soo Time (s) Fig. J.8. Auxiliary feedwater flow--stuck-open TBV with stuck-open MSIV ! from FP. ! 314
250000 , , , . . . .
-500000 LDOP A 200000- -
LocP B -
- -400000 150000- -
-300000 $
M h 1 icoooo. -
- -200000 NOTE: These transients
,,,* ossumed multiple
,,', operotor / equipmenf f ailures. See TABLE I.
50000-- -iO0000 0 . . . . , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 h (s) Fig. J.9. Steam generator mass inventory--stuck-open TBV with stuck-open MSIV from FP. 800 _
, , , , , , , . . -1750 700-- --1500 m
.......... toop a
-1250 i
500- -
-1000 0 400- - O 3 3 E -750 d n 300- NOTE: These fransients -
n j ossumed multiple j operator / equipment -500 200- f allures. See TABLE H. - 3o0 .- 250 L-l y 0- - 1--------~~~---------------~h - 0
-10 0 . . . , , . . . .
0 250 500 750 1000 1250 1500 f750 2000 2250 2500 l h (s) Fig. J.10. , Steamline mass flow--stuck-open TBV with stuck-open MSIV from l FP. 315
7 , , , , , , , , , ,
-800 e .
LD0P A
..... . .. m n 5-
~
~#
4
...~~..'. -s00
....,'- "~-~~-........... - .. , .. q
-40 e
3- . g l N U -
.-s00 E 2
,_ _ NOTE: These transient: __ is, assumed multiple operator e pmen t f ailures. ee ABLE I. _,,
0 250 S$0 7$0 10bo 1250 1500 fNH) 2d00 22'50 2500 Time (s) Fig. J.11. Steam dome pressure--stuck-open TBV with stuck-open MSIV from FP. 1.2 , , , , , , , , , t. 0.8 - - G v 0.s - - 1 g 0.4 -
\
l 0.2 - - NOTE: These transients
- o. assumed multiple .
operator e pmen t f ailures. ee ABLE H.
-0.2 , , , , , , , , ,
i 0 250 500 750 1000 1250 1500 f750 2000 2250 2500 Time (s) Fig. J.12. TBVs flow area fraction--stuck-open TBV with stuck-open MSIV l from FP. l '316
1.2 , , , , , , , , , 1- LOOP A
~
.......... LOOPS 0.8 - -
m 0 0.6 - - NOTE: These transients assumed multiple 0.2_ operator equipment _ f o llur es. ee TABLE I. 0-
-0.2 , , , , , , , , ,
0 250 500 750 1000 1250 1500 175 0 2000 2250 2500 Time (s) Fig. J.13. ADVs flow area fraction--stuck-open TBV with stuck-open MSIV from FP. 800 , , , , , , , , ,
-1600 700- -
600- _
-1200 k 500-- _
E -1000 $ o y 400- l
-800 $ M 300- .
3 NOTE: These fronsients -s30 ossumed multiple 200- operator equipment _ f allures. ee TABLE B. -400 10 0 -. -.-200 X 0 2$0 500 SD 10'0 0 tSO 1500 150 2000 22'50 2500 Time (s) Fig. J.14. Break mass flow--stuck-open TBV with stuck-open MSIV from FP. 317
14oo00 , , , , , , , , , 12cooo.
.-270000 icoooo-- .-225000
?n ?
6 o 5 acoco-- . -180000 y 0 0 2 m j soooo-- . -135000 j 1' moco-- _-soooo 3 NOTE: These fransients 2 5 assumed multiple 2o000-- operaior / equipmeni .-moo f ailures. See TABLE I. o- - -
-o
-20000 , , , , , , , , ,
o 2So Soo 75o 1000 1250 1500 175 o 200o 2250 2500 Trne (s) Fig. J.15. Integrated break flow--stuck-open TBV with stuck-open MSIV from FP. } c 318
APPENDIX K 0.002-m2 HOT-LEG BREAK FROM FP i 319 l l
I l 300000 -
, , , , , , . -660000
-630000 280000- -
-600000 l
se0000-. - _m 1 M
-s40000 j m an,_ NOTE: These transients .
ossumed multiple operator / equipment -5'0000 fa11uras. See TABLE I. 220000- -,,,,,,, 43oooo 200000 , , , , , . . 0 1000 2000 3000 4000 5000 0000 7000 8000 Time (s) Fig. K.l. Primary system mass--0.002-m 2 hot-leg break from FF. U i 4 a i i THETA 1
. . .... .. .. TMUA 2 ,
THCTA 4
- THETA 4 THCTA 4 THETA 4 0.8 - -
l
.5 l -j 0.s - -
u: 0.4 - - NOTE: These fronsients assumed multiple 02- operator / equipment - f ailures. See TABLE I. 0-J - t i l -02 . . . . . . l 0 1000 2000 3000 4000 5000 6000 7000 l Time (s) Fig. K.2. Voiding in the upper head--0.002-m2 hot-leg break from FP. 320
6 , , , , , , 5- -
- 15 4- -
I e 3- . -10 g 3 .5 ti
)
2' NOTE: These fronsients assumed multiple -5
*N operator / equipment
- i. f ollures. See TABLE I. .
o- --o
-1 , , , , , ,
o 1000 2000 3000 4000 5000 6000 7000 Time (s)
- Fig. K.3.
Pressurizer liquid level--0.002-m2 hot-leg break from FP. 5000 , , , , , , , THETA 1 -500
.. . . . . . .. . wo THETA 3 2- <om-- _
-m a-wa .
-600 3 E
3000- - i m
~ A' wg jlI l ~"
a i a
~'**
5- 2000-
- i\ .
t - r l j { \ , 3 -
', , ,\ l -300
$m 1000- . L_ :
; }
. ::::::; 7,-
n , 1oo NOTE: These transients
- assumed multiple -10 0 operator / equipment f ailures. See TABLE H.
l o 1000 2doo 3doo 40'00 5doo 6doo 7doo sooo [ Time (s) l Fig. K.4. Downcomer heat-transfer coef ficients at the core midplane-- 0.002-m2 hot-leg break from FP. 321
800 , , , , , .
- -600 S80- -
- - 570
~
560- ** g '
.. ........ Loo, e 54o e v
I 540-. -
- 510 4, ,_- .-4s0 3 3 E.l _ _eo f[
NOTE: These transients 800-assumed multiple operator / equipment -420 4 o_ f ailures. See TABLE H. . _ -3,0 400 , , , , , , 0 1000 2000 3000 4000 5000 6000 7000 Time (s) Fig. K.5. Hot-leg liquid temperature--0.002-m2 hot-leg break from FP. 0.06 , , , , , , 0.04- - 0.02- - v 0.00 - i -0.02 - . I NOTE: These tronsients assumed multiple , operator / equipment
-0 O'- f ailures. See TABLE I. -
-0.Os , , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 Time (s) Fig. K.6. PORV flow area fraction--0.002-m2 hot-leg break from FP. 4 322
s00 , , , . , , ,
-1600 700-- -
LOOP A
.. . ..... .. Loop g
-1400 000- --
~
-1200
$a M- -
?
l
-1000 h o
f Q m-
-800
},
M 300- . 3
-600 NOTE: These fronsient s 200- ossumed multiple operof or / equipment _-400 f allures. See TABLE H.
E~ -
--200 0
0 100 200 300 400
. D SCO 600 700 s00 0
Time (s) Fig. K.7. Main feedwater flow--0.002-m2 hot-leg break from FP. 0.0s . . . . . 0.132277 0.04- -
- - 0.079366 LOOP A 0.02- . . . . . . . . . . goo, ,
- - 0.026455 d., o g 0.00 -
g M M l j -
--0.02s4ssj
-0.02 -
NOTE: These transients ossumed multiple operator / equipment
--0.07ssee
~0 O'- f ailures. See TABLE I.
-0.0s , . . . , . -0.132277 0 1000 2000 3000 4000 5000 6000 7000 Time (s)
I- Fig. K.8. Auxiliary feedwater flow--0.002-m2 hot-leg break from FP. i 323
i~ IS0000 , , , , . , .
-S00000 Loop A 200000- ---------- tooe s -
-400000 150000- -
} -
-s00000 $
M 1o0000- -
-200000 50000-. NOTE: These transients -
.iooooo assumed multiple Operator / equipment f ailures. See TABLE I.
0 , , , , , . , 0 0 1000 2000 3000 4000 5000 6000 7000 e000 Time (s) Fig. K.9. Steam generator mass inventory-0.002-m2 hot-leg break from FP. 800 , , , , , , - tyso 700-, -
-1500 LOOP A
.. .... . . .- too, e
-1250
$00- -
-1000 Q
6 400- &
- a 0 -
-1s0 g g 300- -
2 NOTE: These fronsients j 200 ossumed multiple - 5# operator / equipment f ollures. See TABLE I. 10 0 - --250 0-k- -
,j . - --O
- 10 0 , . , , , ,
0 1000 2000 3000 4000 5000 6000 7000 Time (s) Fig. K.10. Steamline mass flow--0.002-m2 hot-leg break from FP. 324
I s.75 , , , , ,
- ~845 s.5o- -
LD09 A
~
s.25- - _.... Loop a -900
~
s-1.( -ess 5.75- k -
- -sto w -
3.so-1 --755 E s.2s -- 3 s- - \ *'y'
.pga NOTE: These fronsients 4.75- assumed multiple ! -
equipment t,, -s75 operator 4.5o- f allures. ee TABLE R. i - t. t -eso 4 ss-
-5s5 4 , , , , , ,
o 1000 2000 sooo 4000 5000 sooo 7000
. Time (s)
Fig.K.11. Steam dome pressure--0.002-m hot-leg break from FP. L2 , , , , , , 1-o.s - e t$ - o.s-
.I a.4 NOTE: These transients -
- B ossumed multiple operator equipment f ollures. ee TABLE I. -
o.2 - 0-
-02 . , , , , ,
o 1000 2000 sooo 4000 sooo sooo 7000 Time (s) Fig. K.12. TBVs flow area fraction--0.002-m 2 hot-leg break from FP. 325
12 , , , , , , 1- (DOP A
..........wo,,
0.s - - G V I 0.s - -
^~ ~
NOTE: These transients assumed multiple operator equipmen t 0.2 - f ailures. ee TABLE H. - 0- -
-02 . , , , , .
0 1000 2000 3000 4000 5000 6000 7000 Time (s) Fig. K.13. ADVs flow area fraction--0.002-m2 hot-leg break from FP. 200 . . . . . . .
-400 ps- -
-350 an- -
-soo G5- -
NOTE: These transients -250 assumed multiple
} m.
l operator e ulpment f allures. ee ABLE H. 5 d is. ~
- 15 0 50- _L f -
-100 25-- '-50 0 . , , , . , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) i Fig. K.14.
- Break mass flow--0.002-m2 hot-leg break from FP.
326'
..:.s~-
450000 , , , , . . MOM -~ --900000 350000-, -750000 300000-
- -600000 3r 9 d 250000-g
- --450000
$ 200000--
3 1 g 150000--
-300000 p
, m.
' 300000- NOTE: These tronslent: - I assumed multiple operator / equipment -
- N 00 50000- f allures. See TABLE I.
0-- --O
-50000 , , . . . .
0 1000 2000 3000 4000 5000 6000 7000 Time (s) Fig. K.15. Integrated break flow--0.002-m2 hot-leg break from FP. 327
l APPENDIX L STUCK-OPEN PORV WITH STUCK-OPEN ADV FROM FP ( 328
30C000 , ,
.30000 J300CJ 280000-
-600000 260000- -
-570000 P 9 ca eb 0 i; y -54c000 y 240000- -
'5'0000 NOTE: Thes? fransients assumed multiple 220000 operator e pment f ailures, ee ABLE H. -480000
-450000 200000 . . . . . . .
0 1000 2000 3000 4000 5000 6000 7000 8000 Timo (s) Fig. L.1. Primary system mass--stuck-open PORV with stuck-open ADV from FP. 1.2 , , , , , , , THETA 1
.. . . ...... THETA 3 ,
THETA 3
- THETA 4 THCTA 5 THCTA 4 0.8 -
5 0.6 - h p ! 0.4 - NOTE: These transients - ossumed multiple operator equipment f altures. ee TABLE E. - 0.2 - 0- ' -
-02 . , , , , , ,
0 1000 2000 E00 4000 5000 6000 7000 8000 Time (s) Fig. L.2. Voiding in the upper head--stuck-open PORV with stuck-open ADV from FP. 329
m , , , , , , ,
-30 9 ,
8 -
-25 2 7 -
g Y y 3 (,
. -20 3
.4 b b -
3 5 _ E NOTE: These iroristenis 4- assumed multiplo - operotor equ' pmeni f allures. ee T/BLE I. 3- - -10 2 , , , , , , . 0 1000 2000 3000 4000 S000 6000 7000 6000 Timo (s) Fig. L.3. Pressurizer liquid level--stuck-open PORV-with stuck-open ADV from FP. 5000 , , , , , , , utTA , 800
. . . . . . ...
- THUA 2 NETA 3 p 4000-
_ - )) 4,
-700 .y.
g unA . l -600 m i 6 8 3000-j 1 - (C ( I
- 5 8
i
,A p,\
d,. lh,i ' {h \ i' l 3 y
.- 2000-1
'I t,
IJ f i I
.\ l ,i 1
l -
-400
'E - f:\ ,\ ,
h)[, :, -l39:(,, Q s; }p -300 f , 1 e .
;, . r- ,
m,4
*I 1000-NMN rv %
;h c- -
-200 -
[ i g _t NOTE: These transients ossumed multiple -10 0 operator equipmen t f ollures, ee TABLE L 0 , , , , , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. L.4. Downcomer heat-transfer coef ficients at the core taidplane-- stuck-open PORV with stuck-open ADV from FP. 330
s00 , , , ,
-595 580- -
560 560- -
-525 8
i.co, A 540-j
.......-..u,o,,
E e 520-
-4so j g
h -
-455
# m- - h g
f. 1 400-
-42n
]
6
-385 460-NOTE: These transients -
- ossumed multiple -350 4'0-operotor / equipmenf f ailures. See TABLE I. -
- 315 420 . . i ,
. . i 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s)
Fig. L.5. Hot-leg liquid temperature--stuck-open PORV with stuck-open ADV from FP. 0 300 . . . . . . ,a 0.504- - m 0.502- - U l O k 0.500 - , i' l $
- 0.498- -
l NOTE: These transients o.4,e_ ossumed multiple _ operator / equipment f allures. See TABLE H. l 0A94 , , , , , . . o ioco 200o acoo . coo sooo sooo moo sooo Time (s) Fig. L.6. PORV flow area fraction--stuck-open PORV with stuck-open ADV from FP. 331
1000 . , . . . , , , ,
-2000 LOOP A 300- . .. . .. ..... toop a .-750
-1500 m M- -
6 -1250 o I R
-1000
$g m -
M- -
-750 NOTE: These transients
~
assumed multiple - 5M 200- operator / egulpmeni _ f ailures. See .ABLE I.
-2m l
l 3 0 15 0 250 3de 400 550 650 7h0 8b0 9hD 1000
- T*.me (s) l Fig. L.7.
Main feedwater flow--stuck-open PORV with stuck-open ADV from FP. 22.s . . . . . . . g_ .
-40 17.5 -
is. LOOP A -30
~
12.5 - 6 . ......-.. LDor s
~
'0 - E
, -20 ,
f 7.5 - NOTE: These transients ossumed multiple
~
3- operator / equipment --iO ( 1 f allures. See TABLE H. 2.S ~ 0 - -O
-2.5 . . . . .
W 0 1000 2000 3000 4000 6000 7000 e000 TIME (s) l Fig. L.8. Auxiliary feedwater flow--stuck open PORV with stuck-open ADV from FP. 332
. .. . - _ - _ _ _ _ _ _ _ _ - _ - _ _ - _ _ l
i l l l 250000 , , , , , , ,
-500000 LOOP A 200000- .. ... .. . .. too, e -
l -
-400000 NOTE: These transients assumed multiple 33nono_ operator / equipment .
- f ailures. See TABLE H. ~~
g -
,,,',, -300000og
$ ,'...., a O
2 2 100000- ,...- -
**,,.. -200000 S##- -
.,ooooo 0 , , , , , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s)
Fig. L.9. Steam generator mass inventory-stuck-open PORV ~ with stuck-open ADV from FP. 800 , , , , , , . -us0 7M-. --1500 LOOP A
.......... too, e
-1250 I 500- -
f 6
-1000 g
o 4oo_ _ d -
-750 g i
a 300- - e I j NOTE: These transients j assumed multiple -500 2w- operator / equipment l f ailures. See TABLE H. 10 0 - --250 0-- t.
--- - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - --O
-10 0 , , , , , , ,
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. L.10. Steamline mass flow--stuck-open PORV with stuck-open ADV from FP. 333
7 , , , , , , ,
-s00 l 6- toopA
......... 1DOP 5
\
-750 5- -
4-_ _
-600 $
3 E E E
" NOTE: These fronsients a s-ossumed multiple -~'80 operator / equipment E" f allures. See TABLE H.
2-- --300 g-- --150 0 , , , , , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. L.ll. Steam dome pressure--stuck-open PORV with stuck-open ADV from FP. tz . , , , , , , g- - 0.8 - - 9 v 0.6 -- g k e NOTE: These transients j 0 '- assumed multiple - 4 operator / equipment f f ailures. See TABLE H. 1 0.2 - - 0-
-02 , , , , , , ,
0 1000 2000 3000 4000 6000 6000 7000 5000 Time (s) Fig. L.12. TBVs flow area fraction--stuck-open PORV with stuck-open ADV from FP. 334
t.2 , , , , , , i l 1 wo, a
. . . . . . . . . . w o, .
0.8 - l t G v 0.6 - g k 2 0.4 - NOTE: These transients - S assumed multiple operator / equipment 0.2 -
, f ailures. See TABLE I. -
i: 0- '---------------------------~~------~~----------------------------- -
-02 , , , , , , ,
o tooo 2ooo sooo 4000 sooo sooo 7000 sooo Time (s) Fig. L.13. ADVs flow area fraction--stuck-open PORV with stuck-open ADV from FP. 200 , , , , , , ,
-400 sys. -
- -350
~
15 0 - I NOTE: These transients 300 assumed multiple Q 125 - operator equipment g _ f ailures. ee TABLE H. -250 }
- z g 100- -
g n -200 , E 8 2 n- - 2
- - 15 0 50- _
\7_ --10 0 25-. -
-50 0 , , , , , , , 0 0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s)
Fig. L.14. PORV break mass flow--stuck-open PORV with stuck-open ADV from FP. 335 l 1
350000 . . . . . . .
-750000 300000- -
-600000 250000- .
^
m 3
& o j
- 200000-- . -20000 y CE si sm-a
. 8
-300000 2
g - 3 x m, 100000-g i -- 2 NOTE: These transients -150000 A 50000- ossumed multiple - operator / equipment f ailures. See TABLE H. 0- .
-0
-50000 . . . . . . .
0 1000 2000 3000 4000 5000 6000 7000 8000 Time (s) Fig. L.15. PORV integrated break flow--stuck-open PORV with stuck-open ADV from FP. e . . . . . . . _-So LOOD & 33 _
.. ..... ... too, s .
30- -
-60
)? **~ NOTE: .These transients assumed multiple o
g 20-operator / equipment . -e $ f ailures. See TABLE I. C
= .
h 15 -!
~30 j
P to- - E -15 < 3 . o..'................- ............................ ..- ..o
-5 . . . . . . .
o 1000 200o sooo 4000 sooo sooo 7000 sooo Time (s) Fig. L.16. ADV break flow--stuck-open PORV with stuck-open ADV from FP. I, 336
120o00 , . . , , , , _m 100000- -
-210000 80000-- g,, --175000
.......... w ,a g gg,
- M0000 6 9 o
e 8 -105000 3 40000 . . _ . . . NOTE: These transients . $" ossumed multiple operotor / equipment - 2000 f ailures. See TABLE I. soooo- .
-35000 a_
--35000
-20000 . . . . . . .
O 1000 2000 3000 4000 6000 sooo 7000 sooo Tme (s) Fig. L.17. ADVs integr'ated break flow--stuck open PORV with stuck-open ADV from FP. i 337
APPENDIX M STUCK-OPEN PORV FROM HZP 338
assooo , , , , , , , , , i 2s0000- -
-570000 255000- -
-555000 250000- -
3e000-- -
-S40000
.. mo000- -
bm
-525000 j
) 235000- -
-510000 230000- -
225000-. -
.m 220000- -
-4s0000 215000 , , , , , , , , , , ,
o ano soo 20 1000 teso 1s00 oso 3000 sano asco suo 3000 Time (s) Fig. M.l. Primary system mass-stuck-open PORV from HZP. O.30 . . . . . . , , THETA 1
. ..... .. .. THLTA 2 ,
- TH[TA 3
- THCTA 4 THCTA $
THCTA 4 0.20-8
'4 -
0.15 - w$ o O.10 - . 0.c5-
' - ' ' - -' ~~~-
0.00
-0.05 . . . . . . . .
0 500 1000 1500 2000 2500 3000 3500 4000 4500 Trne (s) Fig. M.2. Voiding in the upper head-stuck-open PORV from HZP. 339
u-
-36
- m. .
..so Q
] s-
-25 i
b g- -
. -20
_a 4- . _g 1 . . . . . . . . . . . e ano soo no sooo uso isoo oso sooo sano asoo smo aooo h (s) Fig. M.3. Pressurizer liquid level--stuck-open PORV from HZP. 8000 . . . . . nEg -500
.......... DEm 2 THEM 3 g 4000- -
f -Q -
-700 a oem.
-soo
- m. l
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\.' , f 5 l 2000-i V ;
5 ; - g , y . i Is -
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o , . . , . o o soo 1000 1soo 200o smoo acoo Time (s) Fig. M.4. Downcomer heat-transfer coefficients at the core midplane--stuck-open PORV from HZP. 340 1
W . . . i 4 . i 4 I- 550-
,g
( - 540-
-500
'. LDOP A g :
i, ---------- Loop s C
$ 510 - E
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- -425 3 3 $
480-
, -400 470-
- -375 4so. -
450 i . . . i i i . 0 500 1000 1500 2000 2500 3000 3500 4000 4500 Time (s) Fig. M.S. Hot-leg liquid temperature-stuck-open PORV from HZP. 040G , , . . . . . . 0.504- - 0.502- - v G 2 M 0.500 - l 0.498- - 0.496- - 0.494 . . . . . . . . 0 500 1000 1500 2000 2500 3000 3500 4000 4500 Time (s) Fig. M.6. PORV flow area fraction--stuck-open PORV from HZP. 341
l e . . . . . . 5 m. -
.. .... ... . too, s
-in 4- .
1 l 1 d, R 3- - 3o D
-s 2- -
0-- --0
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0 260 600 750 1000 1250 1500 1750 2000 2260 2500 2750 3000 Tme (s) Fig. M.7. Main feedwater flow-stuck-open PORV from HZP. 0.0G . . . . . . . . 0.132277 0.04- -
- 0.079366 LDOP A 0.02- ~
......---- 20P e l
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--0.079366
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)
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Fig. M.8. Auxiliary feedwater flow--stuck-open PORV from HZP 342
mm , , , , , l
- -500000 LOOP A 200000-
---------- toor s
- -400000 150000-3
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e . i 0 l j100000--
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l Steam generator mass inventory-stuck-open PORV from HZP. a , , , , , , , , , j , , l 7- ~ Loop A
--O l
.. . .. . . . . . wo, e 1
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-5 2- -
l t- . 0--
--O
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0 250 500 750 1000 1250 1500 1750 2000 2250 2500 2750 3000 Time (s) Fig. M.10. Steamline mass flow--stuck-open PORV from HZP. 343
s.175 , , , , 6.15 0 - ' - . ....._,----- w,, -
-----.o...........
moe s _ ,,g 6.125 - -
--sa5 3 6.10 0 -
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} 0.02- -
l i 0.01- - i e,qq . -
-0.01 , , , , , , , , , , ,
0 260 600 MO 1000 1260 1600 1760 2000 2260 2500 2760 3000 Tme(s) Fig. M.12. , TBVs flow area fraction--stuck-open PORV from HZP. 344
I am . . . . . . . . mA i o,u . .
.. ... ... . . m e 0.o2 -
E O k em - e 9
-o.02- .
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-o.06 . , . . . . . .
0 500 1000 1500 2000 2500 3000 5500 4000 4500 Time (s) Fig. M.13. ADVs flow area fraction-stuck-open PORV from HZP. 50 . . . . . . . . _ f -105
. -90 40-
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}
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. -is o- -
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o soo 1000 1500 2000 2500 3000 3500 4000 4500 Time (s) Fig. M.14. Break mass flow-stuck-open PORV from HZP. 345
100000 , , , , , , , ,
~
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^
120000- . -270000 100000- - -225000 m
$ 80000- . -180000
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- -90000 20000- -
-45000 0- -
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-20000 , , , , , , , ,
0 500 1000 1500 2000 2500 3000 3500 4000 4500 Time (s) Fig. M.15. Integrated break flow-stuck-open PORV from HZP. l 346
i I I REFO'4T NUM*E R fan ,ned Dy r/DC. ### Vof Ne, ef aarJ I U S. NUCLE A J 5.E1UL ATMV COMMISSION feRC n ea , FORM 335
' NUREG/CR-4109 M',"3ja'- BIBLIOGRAPHIC DATA SHEET LA-10321-MS sE NsTRvCriONsOq -E REVERSE f ,
J LEAVE BL ANE 2 VITLE AND 5LOTITLE TRAC-PF1 lyses of Potential Pressurized-Thermal-Shock Trcnsients t Calvert Cliffs / Unit 1 j , ,,,,,,,o,,,co ,,,,,,o A Combustion Engineering PWR f oON1,, vEAR g1984
. ,.u T ,.OR is, Jpe Gregory D. Spr gs, Jan E. Koenig, Russell C. Smith / * '" " "S
vEAR uONr. f
/ February l1985 8 PstOJECY;T ASK WORK UNIT NvWBER 7 v E*. FORM NG ORGANilAf EON NAME D MasLING ADDRESS ffw8verle Codes Los Alamos National boratory ,,,,,,,,,,,,g Los Alamos, NM 8754 A 7315-4 DOR E SS f f=8wde lp Cosej 11a TYPE OF REPORT 10 5+ON50 RING ORG AN12 A180N NAVE AND M AILIN Division of Accident Eval tion Informal Office of Nuclear Regulato Research * "' ' "' "~'"~ '
U.S. Nuclear Regulatory Comm sion Washington, DC 20555 12 SUPPLEMENT ARY NOTES A 13 A85TR ACT f200 words or ieni Los Alamos National Laboratory part . ipa ed in a program to assess the risk of a pressurized thermal shock (PTS) to t cactor vessel during a postulated overcooling transient in a pressurized water reac' (PWR). We provided the thermal-hydraulic analyses of three general accident ca ories: steamline breaks, runaway-feedwater transients, and small-break loss-of- o nt accidents. These postulated accidents included multiple operator and equip ent allures. Results were provided to Oak Ridge National Laboratory (ORNL) wh plan o determine the probability of vessel failure and accident occurrence fo an ove 11 assessment of PTS risk. Our study was performed for a Combustion En neering , Calvert Cliffs / Unit 1, using the Transient Reactor Analysis Code RAC-PF1). We found the results of the analyses to very sensitive to the initial condi-tions of the plant. If the pla was initiall at hot-zero power (compared to full power), the decay heat was muc less, which ma it possible for the same accident initiator to produce significa tly lower downco r temperatures. However, routine opsrator actions may reduce ti consequences of y of these simulated accidents if the prescribed pressure-te perature relationsh s are followed.
- i. oOcuvENT AN ALvs.s . E v.ORos OEsCR.PTOas . a *v,*ggi-Unlimited
'6 SECURITY CLAS$1FICATeoN f r. . -,
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'7 Nvv8tR 06 PAGE5 18 PRICE
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