ML20206D070
| ML20206D070 | |
| Person / Time | |
|---|---|
| Site: | Bellefonte  |
| Issue date: | 03/31/1987 |
| From: | Bayless P, Chambers R, Dobbe C EG&G IDAHO, INC. |
| To: | NRC OFFICE OF NUCLEAR REGULATORY RESEARCH (RES) |
| References | |
| CON-FIN-A-6354 EGG-2471, NUREG-CR-4741, NUDOCS 8704130238 | |
| Download: ML20206D070 (121) | |
Text
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1 s NUREGICR 47410 EGG 2471'
} March 1987 II- Feedwater Transient and Small Break Loss Paul D. Bayless Charles A.Dobbe j.
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of Coolant Accident Analyses for the nosanna chambers Bellefonte Nuclear Plant
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e , s Available from Superintendent of Documents U.S. Government printing Office post Office Bot 37082 Washington, D.C. 20013 7982 and National Technical Information Service Springfield, VA 22161 NOTICE This report was prepared as an account of work sponsored by an agency of 5
the United States Government, Neither the United Sates Government nor any agency thereof, nor any of their employees, makes any warranty, expressed or implied, or assumes any legalliability or responsibility for any third party's use, or the results of such use, of any information, apparatus, pmduct or proc.
ess disclosed in this report, or reprewnts that its use by such third party would not infringe privately owned rights.
i NUREGICR-4741 EGG 2471 Distribution Category: R4 l
FEEDWATER TRANSIENT AND SMALL BREAK LOSS OF COOLANT ACCIDENT ANALYSES FOR THE BELLEFONTE NUCLEAR PLANT l
l l Paul D. Bayless Charles A. Dobbe
! Rosanna Chambers Published March 1987 EG&G Idaho, Inc.
Prepared for the Division of Accident Evaluation Office of Nuclear Ro0ulatory Research U.S. Nuclear Regulatory Commission Washin0 ton, D C. 20555 Under DOE Contract No. DE AC07 761D01570 FIN No. A6354
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l ABSTRACT Specifie sequences that may lead to core damage were analyicd for the llellefonte nuclear plant as part of the U.S. Nuclear Regulatory Comminion's Sescre Accident l
Sequence Analysis Program. 't he Rl! LAP 5, SCDAP, and SC11\P/Rlil AP5 com-puter codes were used in the analyses. I he two main initiating esents investigated were a low of all feedwater to the steam generators and a small cold leg break low of coolant accident. The transients of primary interest within these categories were the TNILil' and S D sequences. Variations on sprems asailability were also imestigated.
Powible operator actions that could pres ent or delay core damage were identified, and two were imestigated for a small break transient. All of the transients were analy/ed until either core damage began or long term decay heat remos al was established. 'I he analpes showed ' hat for the sequences considered the injection flow from one high-prenure injection pump was necewary and sufficient to present core damage in the l
! absence of operator actions. Operator actions were able to present core damage in the S:D sequence; no operator actions were asailable to present core damage in the
'I N11.II' sequence.
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IlN No. A6354-Sescre Accident Sequence Analpis li i
EXECUTIVE
SUMMARY
Nine postulated tramients initiated by either a total system would deprewurize. The lower sptem pres-loss of feedwater to the steam generators or a small jure at the time of sessel melt through would cold leg break were analyzed for the llellefonte nuclear reduce the estent of direct containment heating.
power plant. The transients varied in the sptems and in-sewel natural circulation during the TNil.II' compments as ailable. 'I he analyses were per formed to transient was imestigated by using a three-channel determine those combinatiom of operating equipment model ofIhe core and upper plenum (Case 17-1c). A that lead to a succewful recosery of the plant or to core natural circulation How was established in Ihe reae-damage. Two of the transients that resulted in core for sewel after the core began to unemer, with damage were analyzed through that portion of the s apor leas ing Ihe middle of Ihe core being cooled in transient. the upper plenum and returning to the core through The analyses were performed at the Idaho the lower powered periphery of the core. This How National lingineering 1.aboratory as part of the tramferred heat from the core to the upper plenum United States Nuslear Regulatory ComminionN structures, slowing the heatup of the core. When Sescre Accident Sequence Analysis Program. The the fuel rod cladding in the inner channek bal-transients were anal >/ed using the Rlil.AP5/ looned, howner, this 110w was disrupted.1 he flow NIOl)2, SCI)AP, and SCDAP/Riii.AP5 computer area reduction camed by the ballooning diserted codes. These codes calculate the sprem thermal- flow from the middle of the core toward the periph-hydraulies. core damage and integral core damage cry. Thk flow diversion roersed the flow in the and sptem thermal hydraulics, respectisely, outer region of the core; the flow throughout the llellefonte is a prewurized water reactor doigned core was then from the bottom to the top.
by liabcock and Wileos and being built by the A TNil.II' tramient that included failure of the fennewee Valley Authority. The two transients that reactor coolant pump shaft seah (Case 17 2) was were analy/ed through the core damage phase of analy/cd up to the time of core damage. T he pur-the transient were the I N11.II' and S,1) sequences. pme of this analysis was to determine if hasing a 1he Accident Sequence lisaluation Program has leak near the loop seah would promote the clearing identified these as two of the rhk-dominant of the liquid from the loop seah, thus reestabihh-sequences for prewurized water reactort ing natural circulation through the loopt 1)opite l he I hlt it' sequence is characteri/ed by the low of ming sery comervathe salues (designed to pro-all ae power for I to 3 h, the low of ausiliary feedwater, mote clearing of the loop seah) for the leak rate and and no operator actiom. The SCI)AP analpis of this failure time, the loop seah did not clear.
tramient (Case 17 le) predicted that fuel rekication A ihird irausient was simulated in w hich all feed-would begin about I h after the transient began. t he water was im', but one high prewure injection SCI)AP/ Rli!.AP5 analpis (Case l' Id) predicted that pump was able to supply water to the reactor cool.
the hquid in the loop seah, w hich had been prnenting ant sprem (Case I? 3). The asalfability of the one sapor from nowing through the coolant loops, would high preuure injection pump was sufficient to pre.
How into the reactor sewel, caming a partial core sent the core damage that had occurred in the quench and reestablkhing loop natural circulation iNil ll' sequence. The sptem reached a stable How. With the loop natmal eirculation, the corc heatup leed and-bleed configuration, with water being i calculated to be much slower, so that fuel relocation added by Ihe high preuure injection pump and woukt not occur until about to h after the tramient steam being remmed through the power operated began. relief sake. 'ihe borated water storage tank con.
~lhe natural circulation How through the loops tained enough water to maintain the feed and, aho camed the temperature of the reactor coolant bleed operation for about 24 h. A similar tramient splem piping to increase. Iaihue of the reactor was analy/ed that awumed that the power operated coolant sptem piping could occur in the piping or relief sahe could not be med(Case 144). Again, the in the steam generator tubo long before the core high preuute injection proented core damage.
began to relocate. l'adure of the steam generator flowner, the now through the safety relief sahn tubn woukt prmide a path for fiwion products to roulted in more cycin of the relief sahes during b> paw the containment. With a failure of the reae- the tramient than were npected for their lifetimes.
tor coolant sprem boundary prior to the core melt- Ihe small break low of coolant accidents were ing and nentuall) failing the reactor sewel, the initiated by a 5 cm (2 in.) break in one of the cok!
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legs. The first tramient awumed that no actise sys- damage for the S:D sequence. Ily following the tems to mitigate the tramient were asailablet only Abnormal Tramient Operating Guidelines, the l the two core flood tanks were a ailable (Case S 1). plant was deprewuri/cd so that low prewure injee- >
The core began to heat up after about 2l00 s. Core tion was mailable to cool ihe core. No such opera- l damage (cladding osidation) was calculated to for actiom were available for ihe TMl.11' sequence, i begin near 2750 $ when the masimum dadding the tramients that did roolt in core damage gener-temperature reached 1000 K (l340'12). ally had shm heatup rato. The slow heatup rato alkm A small break low of wolant accident was imeu much of the cladding to In midited before any rdoca-tigated in w hieh all feedwater sptems were unasail- tion of the fuel rtxh begim. Ihe otemise midation l able and only one high prewure injection pump will aho affeet the containment ropome after the core was mailable (Caw S 2h The core liquid locl was material mehs through the reactor seswl.
increasing after about 3 h, with an adequate supply Natural circulation llow in the reactor sewel ,
of makeup water to keep injecting liquid for at least reduecd the core heatup rate. Ily tramferring more {
12 more hours. T he fuel rod dadding temperaturo energy to the structures in the upper plenum, the ;
remained at or below the saturation temperature fuel rod dadding temperaturn increawd more ;
throughout Ihe tramient, slowly. I he slower heatup rate will dda) mre dam- :
T he S;D t ramient is a small break low of coolant age. 'I he higher st ruct ure temperatures in the upper i aeddent with no high.prosure injection and no plenum will re uli in fewer fiwion products being operator actiom (Caw Sdk Ihe core unetwered retained in the reactor seuel, llallooning of the dadding disrupted the inaewel I while still abme the preuure of the core flood tanb. Although the prewure later durcased, the natural circulation 0aw. in the 131l ll' wquence, dad- (
injuihm or hquia nom ine m ,e nooa ian a was dino.dh,ooinginihemiaak.ofiheo,eimiumaine ;
not able to proent socre core damage. I uel reloca- coolant flow area, forcing How toward the wre lyriph. l tion would oceur about 12,000 s atter the start of cry. 'l hintopivd the natmal circulation ihm tetur ning l the tramient. Ilecame of the slow heatup rate, to the core frem the opper pienum. Adihtional ddap ,
otemise midation of the dadJing was predicted, in core damage that wouki rouh f rom continuai natu-such Ihat about We of Ihe /irealoy in the core had ral circulation in the reastor sowl were then lost. ,
midi /ed before fuel rehwation began. The loop t he dearing of the loop seah allowed the entire l wah contained enough liquid throughout the core reactor coolant sptem to heat up with the wre. !
heatup that no loop circulation was rentablished. While redudng the core heatup rate, the high tem- ,
Operator actiom were aho imotivatal for the S;D peratures ht the loops would came the piping to fail, by either creep ruptureor meking,long before l transient (Caws S4 and S 5L Pnweduro bawd on the Abnormal tramient Operating Guidehno were naid- the wie began to relocate. 't he powible failure of l cled thing thew pnxeduro, wre damage uas emily the steam generator tube would prmide a path for !
moidal, t he plant was rennered with the liquid lod flulon produsts to b> pan the containment. 't he in the reastor sowl alxwe the core and with the km- failure of the reactor coolant spiem wouki aho allow the setem preunte to decreaw prior to reae. I prewure injation pumps on. Socral short dadding temirrature nomiom occurral, but the grak d.nl- tor senel failme. T he olent of direct wntalmnent ding temperature was noer far atxne the stead >4 tate heating at the time of reactor send failure would operating temperatuie. be reduced, table liS.I summati/o the f ramient analpes. ihe analpo demomtratal the importance of gyr.
Ihe low of feedwater and small Neak tramient forming integral calculatiom of the mee damage and analpo showed that one high prewure injection splem thermal hydraulie behaslor. Ily ming the pump was sullicient and necewary,in the abwnee MlHP/Rl!I APS omie, the effats of the huip wal of operator asilom, to proent core daman for dearing and hiop drentation on the one and uru(tme thew tramients. Stable long term deca) heat temperaturo und of the balk = ming on the natural dr.
termnal um etablkhed with the core cooled w hen m!ation ihm in the reactor sowl cooki be knotivated, one high.prenure injntion pump was mailable, wherem they muki not with separate calmlathms of in the abwnse of the high preuure injullon, the spiem thermal h>drauho (RI l AP3) and the mte operator actiom were susceufut in prcenting core damage INCinP).
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l ACKNOWLEDGMENTS T he authors esprew their appreciation to Dr.11.11. Agrawal of the Nuclear Regula-tory Comminion, the sponsor of this work. We also appreciate the support of the Tennewee Valley Authority,especially Ken Keith and Mark .\1 iller,in prosiding infor-n.ation on the llellefonte plant and its espected operation, in resiewing the resulh of the analpes, and in awisting in the deselopment of the operator action scenarios.
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CONTENTS i
Alls I R AC I . , ii I I!NI ( U IIVI! SLININI ARY . iii ACKNO\VI.lilXi\ll N IS ,. si ACRON Y NIS . . Niii IN 1 ROI)UC IION , , I I til l) Wall:R I R ANSiliN I S . . 3 l
l I \ll 11' Analpes . , , 3 Rl:1. Al'5 Analpes . . ., , . ,, 3 :
1 SCI)Al' Analpis .. . 14 l l
SCI)Al'/ RI l Al'5 Analpes ,, , . , , 18 I Nil 11' Sequence with RCI' Slial't Seal I eakage . .. , ,. 31 1 Of W with One lil'l l'Hinp Asailalile , ,. , , ,. , ,, . 34 i
I SllliltilaI) ol I cesladtel ll'iillsiefil Nesults , ,,. . . . ... 3')
S\l Al 1. IIRl! AK l OSS Ol' COOL. AN I ACCll)l:N I S , . 42 Sinall lireak i OCA witti 101 W anti No I:CC .. .. . . 42 l
Sinall liteak l()C A uilli l Oi W afiti l'aitial lil'l , , ,. . . 44 I
l S,1) Sequt nee Alialpis , , , , ,, , ,. . , , 47 S.l)()perator Action linestigations . ,, , ,, , ,. ,, $(i I l Al()(i lor Ihe S,1) Sequellee , ,, ,. ., $(i l
(1petator Actioli Case 1 ,. ., , ., ,,,, , ,,, ,, 57 l ()petalot Actioll ('aw' 2 , .. . ,,, . . . , $9 Snnunary of Sinallilical l()CA Resuln , ,, . . .. . . (i2 cal ( lli AllON \1 t tN( l RI AINill S . , ,, , , , , . .. , , , (i$
l I
( CON ( l(ISIONS . . , , ., , ... , ... , , .. (iN l
l Rl lI kl NCl S , . .. , , . . ,,, ,, ..., ,,,, 70 1
\l'I'l NillN A ( O\ll'lill R C01)l I)I SCRil'IlONS ,, ,, , ,, ,,. .. . A.1
\l'I'l NI)l\ ll -lit i I l l(IN II l'l AN I lil SCRll'1lON , .,,, , ,, ,, , , 11 1 Al'I'l Nill \ (' \lul)I l 1)! S( Ril' lit)NS , . ., ,. , 01 ui l
FIGURES Prewuriier pressure for the single-channel REl AP5 TNILil' calculation 5
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- 2. Steam generator pressures for the first 300 s of the single- channel 6
REI AP5 TNil.II' calculation . . . .. ... . .. ... . .. ..
Steam generator pressures for the single-channel RELAP5 TNILil' calculation . . .. 6 3.
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- 4. Steam generator liquid mawes for the single-channel REL AP5 TNILil' calculation . .
Prewuriier collapsed liquid lesel for the single-channel REl.AP5 TNILil' calculation . . . . 8 5.
- 6. 1.iquid temperatures in the hot and cold legs for the single- channel 8
REl.APS TNil.II' calculation .. . ... . . .
Ilot leg maw tiow rates for the single-channel REl.AP5 TNil.II' calculation 9
- 7. . .
N. Slaw flow rates at the core b) pan inlet and outlet for the single-channel 9
REl.AP5 TNil 11' calculation . .. . . .. . . .. .
9 Void fractions in the reactor sessel upper head and upper plenum for the single channel Rlil Al'5 TNil.II' calculation . . . . . .. .. . .. .. . 10 Void fractions in the core for the single-channel Rl! LAPS TN!!.ll' calculation 10
- 10. .
- 11. I uel rod cladding surface temperatures for the single-channel 12 Riii APS Thlt.it' calculation . . . . . . . .. ... . . . . . ..
- 12. Vapor temperatures in the hot legs for the single-channel 12 REI APS INil 11' calculation . . . . . . . . ... . .. . .. .
Prewurlier prewure for the single- and three-channel REl.AP5 'INil.it' calculations . 13 j 13.
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- 14. I uel rod cladding surface temperatures for the single- and three-channel 13 i REl. AP5 ~I N11.It' calculations . ... . .. ... . ... .. . ..... .
- 15. Core outlet man flow rates for the three channels for the three channel 14 Ri l.AP5 'I NILil' ealculation . ... ... . .. . .. . .. . . .. . .. ...
l'uel rod cladding surface temperatures for the SCI)AP I Nil.ll' calculation . . . . . . 16 16.
l'oel awembly coolant flow areas for lhe SCI)AP 'I Nil.II' calculation . . .... . .. 16 17.
lotal hydrogen generation for the SCI)AP 'l Nil.ll' calculation . . . . . . . . . . ...... 17 18.
- 19. Soluble hwion product (cesium and iodine) release rate from the f uel rods for lhe IN NCl)AP I Nil.ll' ealculation . ... . .... .. . . . . .. . ... . . . ...
- 20. l'uel rod cladding surface temperatures at nodes 1. 3, H. and 10 for ihe 19 single channel SCI)AP/Riil AP5 INil.ll' calculation .. ........ .. . ... ....
s iii 1
- 21. Total hydrogen production for the single-channel SCDAP/RELAP5 TN!Lil' calculation . . . .. . ..... . . . .. 21
- 22. Noncondensible quality at core nodes 8. 9. and 10 for the single-channel SCDAP/REl APS TNill!' calculation . . .. .. .. . .. . 21
- 23. Vapor temperatures in hot legs A and 11 for the single-channel 4
SCDAP/RELAP5 TNil l!' ealculation . ..... .. . . ...... . . . ... -22
- 24. Vapor temperatures in the Il-loop hot and cold legs for the single-channel SCDAP/RELAP5 TN1Lil' calculation . . . . .. ... . . .. .... . .. 22
- 25. 1 uel rod cladding, hot leg nozzle. and steam generator tube temperatures for the single-channel SCDAP/RELAPS TN1Lil' ealculation . ... ... . .. ...... 23 1 26. Pressuri/er prewure for the single-channel SCDAP/RELAPS Thill!' calculation .. ..... 23
- 27. Collapsed liquid lesel in the pressurizer for the single-channel SCDAP/RELAPS TNill!' calculation .. .. .. .... . ... .. . . . .. 25 f 28. Noncondensible quality in the containment for the single-channel i
SCDAP/RELAP5 TNill!' calculation .. . .. . . .... . ... .. . .. 25 l
29 Core and upper plenum flow patterns b Aire and after ballooning for the j three-channel SCDAP/REl.AP5 TN!!.II' calculation . . . .. ..... . .. . .. 30 l 30. I uel rod cladding surface temperatures at the top of Ihe core for the three-channel and single-channel SCDAP/RiiLAPS TN11.II' calculations . . . . . . .. . 31
- 31. Pressuri/er prewure for the pump seal leak and base case REl.APS TN!Lil' calculations . .. .. . . .. . . . . .. . .. . . 33 4
- 32. Pressurizer collapsed liquid lesel for Ihe pump seat leak and base case REl.AP5 TNILil' calculations .... .. .. ... .... ..... . .. . .. .. 34
- 33. 1 uel rod cladding surface temperature at the top of the core for Ihe pump seat leak and base case REl AP5 T N!! 11' calculations . .. ...... . . . .. .. 35
- 34. I oop A hot leg man flow rate for the REl.AP5 pump seal leak calculation . . . ... . .. 35 4
- 35. 'Iotalleak flow through the RCP shaft seals for the REl.APS pump seal leak calculation . . , , . ..... . . ... ......... ....... ... 36
- 36. Prewuriter piewure for the i OI W transient with one llPI pump available . . . . . . . . .. . 38 ;
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- 37. Prewuriter prewure for she i OI:W tramient with one llPI pump and no I PORV asailable . .. .. .. . . . . ... . ......... . . .. ...... . .... 38 l
- 38. Reactor coolant system maw imentory change for the i OI:W tramients with one llPI pump mailable . . . . . . . . . . . ..... ... ......................... . 39 i
j 40. Collapsed liquid leselin the core for the small break i OCA with 1.OI:W and no !!CC . . . 43 i, is
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- 42. Fuel cladding surface temperatures for the small break LOCA with LOFW and no ECC . . . ... . . . ... . . ....... . .. . . ...... ... ... 45
pump asailab!c ... ... ... .... ... .... . . ..... . .... ..... 46 l
j 44. Collapsed liquid level in the core for the small break LOCA with LOFW
, and one HPI pump available . . ........ . ....... .. ... ... . .. . 46 l
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- 45. Pressurizer pressure for the S,D transient .. . . . ......... ... .. ... 49
- 46. Collapsed liquid lesel in the core for the S,D transient . . . . ............... . . 49
- 47. Fuel rod cladding surface temperatures at nodes I,5,8, and 10 for the S,D transient .. 50
- 49. Control rod guide tube temperatures at nodes I,3. 8, and 10 for the S:D transient . . .. 51
- 31. Volume-aserage piping temperatures in the broken cold leg for the S2 D transient ... ... 53
- 52. Niass flow rate through the break for the S D transient .... ......... .............. 54
- 53. Void fraction in the solume upstream of the break for the S,D transient ............. 54
$4. Liquid solume in the CFTs for the S,D transient . . .... ... . ..... ..... ...... 55
- 55. Noncondensible quality in core nodes I,4,8, and 10 for the S,D transient . .... .. .... 55
- 56. Pressure in steam generators A and 11 for the S D transient . .... .... ............ 56
$7. Pressuriier pressure for S,D operator action case ! .. . .. .. .. .. .. .. ...... ... .. .. .... . .... ........ 59
- 58. Pressure in steam generators A and il for S,D operator action case I . . . . . . . . . . . . . . . . . 60
- 59. Fuel cladding surface temperatures in nodes 8,9, and to for S,D operator action case i . . . 60
- 60. Upper plenum and hot leg vapor temperatures for S,D operator action case i . . . . . . . . . . . . 61
- 61. Alass flow rate through the break for S,D operator action case t . . . . . . . . . . . . . . . . . . . . 61
- 62. Collapsed liquid lesel in the core for S,D operator action case i . . . . . . . . . . . . . . . . . . . 62
- 63. Fuel rod cladding surface temperatures in nodes 7,8,9, and 10 for S D operator action case 2 . . . . . . . . . . . . .... ...... . . ... .. ................... .., .. 63 11 1 . General schematie of the licliefonte reactor coolant system . . . . . . . . . . . . . . . . . . . . . . . . 11-4 x
I 11-2. Diagram of the ATOG response to the S,D transient . . . . . .. Il-7 1
11-3. Diagram ofIhe ICC portion of the ATOG response to the S,D transient . 11-8 B-4. Definition of the ICC regions as determined by the core exit temperature . . , 11-1 0 C-1. Nodalization of reactor coolant loop A for the RELAPS calculations . C-4 C-2. Nodalization of reactor coolant loop 11 for the RELAP5 calculations . .. . . . C-5 C-3. Nodalization of the reactor vessel for the single-channel RELAPS calculations . C-8 C-4. Nodalization of the core for the three-channel RELAP5 calculation . . C-10 C-5. Cross-sectional view of the regions in the three-channel core models . ... . .. . C-1 l C-6. Nodalization of the reactor sessel for the SCDAP calculation .. . . C-12 C-7. Crow-sectional siew of a typical N! ark C fuel assembly . . . C-13 C-8. Nodalization of the core and upper plenum for the three-channel SCDAP/RELAP5 calculation . . . .. . .. . ... C-15 C-9. Nodali/ation of the high point vents, hot leg dump-to-sump, and relief
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satse dow nstream piping for the S,D operator action calculations . . . . C-16 l l
C-10.Nodalization of the split steam generator for the S,D operator action calculations . C-17 TABLES
- 1. Sequence of esents for the RELAPS calculation of the TNILil' sequence . . . .. 4
- 2. Sequence of esents for the SCDAP calculation of the TNILil' sequence . . .. . 14
- 3. Sequence of esents for the integral and separate code cateulations of t he 'I NILil' sequence . . . . .. . . .. 26
- 4. Sequence of esents for the single- and three-channe(integral analyses of t he T Nil.II' sequence .
... .. . ... . .. 29
- 5. Sequence of esents for the pump sealleak and base cases .. .... . . 33
- 6. Sequence of esents for the i OFW with IIPI transients .. . . .. . .. 37 7 Summary of the loss of feedwater transient analyses . .. . ..... .. . 40
- 9. Sequence of esents for the S2 D transient . ... . .. . .. . .. .. . 47
- 10. Sequence of esents for the two S,D operator action cases . . . .. ... . .. . 58 l1. Summary of the small break i OCA analyses .. .... . . . .. . ..... .. 64 si
=--
11-1. Component and control system setpoints for llellefonte . . ... ...... .. ... . 11-5 Il-2. Comparison of computed and desired steady-state parameters . . . ......... .. . 11- 6 C-l. Correspondence between the physical and mathematical components in the primary loops for the RELAPS model of lleliefonte . . . . . . . . . . . . . . . . . . ... .. . C-6 C-2. Correspondence between the physical and mathematical components in the A. and ll-loop steam generator secondaries for the RELAPS model of llellefonte . . . . . . . C-7 C-3. Correspondence between the physical and mathematical components in the reactor sessel for the RELAPS base case model of llellefonte ... .. .... ........ . .. C-9 C-4. Correspondence between the phnical and mathematical components in the core for the RELAPS three-channel core model of llellefonte . .. . . .... .... . . . C-11 C-5. Relation between the core axial nodes and the height abose the bottom of the fuel for the three code models .... ... .. . . .. . .. . .... C-14 sii
ACRONYMS AFW ausiliary feedwater ATOG Abnormal Transient Operating Guideiines ll&W llabcock and Wilcox IlWST borated water storage tank l
CDV condenser dump vahe '
CIT core flood tank ECC emergency core coolant or cooling ECCI emergency core cooling injection ESFAS emergency safety feature actuation system FSAR Final Safety Analysis Report ilPI high-pressure injection ICC inadequate core cooling ICS integrated control system INEl. Idaho National Engineering Laboratory I.OCA loss of coolant accident LOFW loss of feedwater I.Pl low-pressure injection 1.W R light water reactor NIADV modulating atmospherie dump vahe N1FW main feedwater SIS main steam 51SIV main steam isolation salve NNI non-nuclear instrumentation NSSS nuclear steam supply system PORV power-operated relief sahe PWR pressurized water reactor siii
4 9
-i RCI' reactor coolant pump ,
i RCS reactor coolant system SASA Sescre Accident Sequence Analysis SG steam generator I
SRV safety relief sahe TVA Tennessee Valley Authority l
i i
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0 4 r l
4 i
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a t
i i
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XIV
FEEDWATER TRANSIENT AND SMALL BREAK LOSS OF COOLANT ACCIDENT ANALYSES FOR THE BELLEFONTE NUCLEAR PLANT INTRODUCTION So cral specific t ransients t hat may lead to core dam- these analyses, core damage is defined to begin age in the llellefonte Nuclear Plant were analyzed. The when Ihe fuel rod cladding temperature reaches analpes were performed at the Idaho National Engi- 1000 K (1340 F), the temperature at which oxida-neering laboratory (INEL) as part of the United States tion of the cladding begins in the core damage cal-Nuclear Regulatory Commission's Soere Accident culations. The TMLil' and S D transient analyses Sequence Analysis (SASA) Program. This program were extended through the severe core damage was established to caluate postulated reactor aecidents phase of the transient, until the assumption of a mer a broad spectrum of accident sequences. These rodlike geometry in the core was no longer salid.
sequences may estend beyond the current design basis The analyses of these two transients also included in terms of sptem failures, core damage, and release of identifying and investigating operator actions that fission products to the emironment. The objectives of may delay or prevent core damage.
the SASA Program are to evaluate nuclear plant The other feedwater transients imestigated were a response for accident sequences that could lead to par- loss of all feedwater with one HPI pump available, tial or total core meh, to determine the timing of signif- toth whh and without ihe power-operated relief valve icant cents, to determine the magnitude and timing of (PORV) available, and a TMLil' sequence in which fission product release from the fuel rods and the leakage through and failure of the reactor coolant hydrogen generation rate, and to oaluate the effect of pump (RCP) shaft seats was modeled. llesides the S D operator actions on accident mitigation. sequence, other small break IDCA transients included The analyses also proside information that can the loss of all emergency core cooling (ECC) and auxil-be used in Ihe resolution of outstanding soere acci- iary feedwater ( AFW) systems, and the loss of all ECC dent issues. Issues identified in NUREG-0956, and AFW systems with the exception of one HPl
" Reassessment of the Technical llases for Estimat- pump. Two further analyses of the S 2D sequence imes-ing Source Terms,"I for w hich information is pro- tigated operator actions to mitigate the transient.
vided by these analyses include hydrogen The analyses were performed using three com-generation, the release and retention of fission puter codes. RELAPS/ MOD 23was used to calcu-products, the estent of direct containment heating, late Ihe plant response from transient initiation up and the effects of in-sessel natural circulation. to core damage for all the sequences, SCDAP4 cel-The transients imestigated were based on two culated the core damage portion of one of the initiating events, a loss of all feedwater to Ihe steam TMLil' analyses, based on input extrapolated generators and a small [5.1-cm (2.0-in.) diameter] from the RELAPS calculation. SCDAP/
cold leg break loss of coolant accident (LOCA). For RELAP5 5was used to perform an integral calcula-each initiating event, several specific transients tion of the TMLil' and S D2 transients from were investigated in which different systems were initiation through severe core damage. The opera-auilable. T.he transients of primary interest in each tor action calculations were also performed using group were the loss of all feedwater with concurrent SCDAP/RELAPS. Descriptions of these codes are loss of all ac power, w hich is designated the TM Lil' provided in Appendix A.
sequence, and the small cold leg break with no llellefonte is a PWR designed by liabcock and high. pressure injection (IIPI), which is designated Wilcox and being built by the Tennessee Valley Author-the S D sequence. These two transients were ity. It has a rated core thermal power of 3600 MW selected because they are two of the risk-dominant distributed among 205 fuel assemblies. The plant has sequences for pressurized water reactors (PWRs).2 two hot legs, two once-through steam generators, and Each of the transients was analyicd from initiation four cold legs. The loop with the pressurizer is referred until either core damage had begun or until long- to as loop A; the other loop is loop II. Two core flood term decay heat remosal had been established. For tanks (CFTs) inject liquid into the reactor vessel I
4 downcomer. The 11PI pumps inject into the cold legs, were assumed to remain constant. End-of-cycle and the low-pressure injection (1.PI) pumps deliver fuel conditions were used to maximize the fission water to the reactor vessel downeomer. The anxiliary product buildup in the core. The input models used y
feedwater is injected near the top of the riser portion of for Ihe various codes are described in Appendix C.
the steam generators. Appendix il contains more infor- The following sections present the feedwater mation on the llellefonte plant, including a diagram of transient analyses, the small break LOCA analyses, the plant. a discussion of the uncertainties in the calculations, The plant was assumed to be operati ng at full conclusions, and references. Appendices A, II, ,
power when each of the transients began. All fune- and C provide further descriptions of the computer tioning plant systems were assumed to be operating codes, the llellefonte plant, and the input models, i
at best-estimate capacities. Relief salve setpoints respectisely.
l i
4 2
l l
- This series of transients was initiated by a com- that core damage began. The core thermal-plete loss of feedwater (LOFW); that is, neither hydraulic conditions at the end of the calculation main nor auxiliary feedwater How was available to were used as input to a SCDAP analysis of the core either steam generator. The concurrent failure of damage portion of the transient. Alore recent anal-various other systems and components defined the yses have used the integrated SCDAP/RELAPS transients analyzed. computer code to analyze the system thermal-The base case transient was t he Tht L B ' sequence hydraulics and the core behavior from transient ini-with no operator actions considered. The ThlLB' tiation through severe core damage. hiost of these sequence is defined as the loss of all offsite ac analyses used one-dimensional models of the sys-power, the failure of the diesel generators to provide tem. Analyses of the transient through the initial onsite ac power, and the failure of the steam-driven core heatup have been performed using three-auxiliary feedwater pump to supply water to the channel models of the core and of the core and
, steam generators. No ac power is restored for 1 to upper plenum. Each of these analyses is discussed 3 h. As the transient begins, power is lost to the in more detail below.
control rod drives and to all the pumps. The reactor scrams, and the main feedwater and RCPs begin to RELAPS Analyses. The Th1LB' sequence was coast down. Feedwater flow is quickly reduced to analyzed using the RELAP5 computer code with zero as the main feedwater valves close. The turbine both single- (Case F-la) and three-channel stop valves close, and the pressure in the steam gen- (Case F-1b) models of the core. The three-channel crators increases until the relief or dump valves model was used to approximate a two-dimensional open; the pressure is maintained fairly constant model of the core.
thereafter. As heat is transferred from the reactor coolant system (RCS) to the steam generators, the liquid in the steam generators is boiled. When the Single-ChannelCore. The sequence of signif-icant events for theTAILB' sequence (Case F-la)is hquid is gone (after about 300 s), the steam genera-summarized in Table 1. At the beginning of the tors are no longer able to remove the decay heat transient, the reactor scram was initiated, the tur-being generated in the core. The RCS begms to heat bine stop valves and main feedwater valves began to up and pressurize. The PORV and safety relief valves (SRVs) connected to the top of the pressur-close, and the RCPs began to coast down. liigh tier controlIhe pressure. Afler the RCS reaches sat- loop flow driven by the RCP coastdown provided adequate heat transfer through the steam generator uration, a high-pressure boilof f ensues, resultmg in core uncovering and heatup. The pressure remams tubes to boil off the secondary liquid inven:ory by 350 s. (The transient assumes no AFW injection.)
above 15.76 h1Pa (2285 psia) throughout the Steam generated in the steam generator secondaries i uncovering and heatup of the core. The transient proceeds to severe core damage since there is no way was relieved to the atmosphere via the secondary to replace the fluid bemg lost from the RCS SRVs and modulating atmospheric dump valves through the relief valves at this high pressure.
(NIADVs). During this boiloff period, the heat The next section desenbes several analyses of the transfer across the tubes decreased as the tube sur-TNILB' sequence in greater detail. The sections that face area in contact with liquid decreased. After the
- follow present analyses of a Th1LB' sequence in w hich RCPs coasted down, natural circulation resulting leakage through and failure of the RCP shaft seats was from the density difference between the fluid in the considered, and of two total IDFW t ransients in w hich downcomer and Ihat in the core began in the loops.
one llPI pump is available. The last section summa- When the heat transfer rate across the tubes dropped below the core heat transfer rate to the rizes the results of these analyses.
coolant, the primary system began to heat up. The increasc in primary coolant temperature caused the TMLB' Analyses liquid to expand and the primary system pressure to increase. The primary system pressure increased to This section describes several analyses of the the PORV open setpoint of 15.93 N!Pa (2310 psia)
ThlLB' sequence. The first analysis used the at 250 s, where the pressure was stabilized by steam RELAPS computer code. This calculation encom- discharge through the PORV. When the pressurizer passed the transient from initiation up to the time filled with liquid, the PORV could no longer relieve 3
Table 1. Sequence of events for the RELAPS calculation of the TMLB' sequence Time Event .
(s)
Scram signal 0.0 Reactor coolant pump trip 0.0 Main feedwater valves begin to close 0.0 Thrbine stop valves begin to close 0.1 Thrbine stop valves closed 0.2 Main feedwater valves closed 2.0 Control rods fully inserted 3.1 Power-operated relief valve initial opening 250.
Steam generator B dry 250.
Steam generator A dry 350.
Primary saturates (hot legs) 950.
Primary safety relief valves initial opening 1000.
Natural circulation ends 1100.
Fuel cladding heatup begins 1800.
Fuel cladding temperature reaches 1000 K (1340*F) 2550.
< Dryout at bottom of fuel 2880.
Calculation terminated 3000.
the pressure and the primary system then pressur- maximum cladding temperature reached the calcu-ized to the SRV open setpoint of 17.24 MPa lated oxidation initiation temperature of 1000 K (2500 psia) near 1000 s. The primary coolant tem- (1340*F). The calculation was extended to provide j perature increased until the saturation temperature the core dryout time of 2880 s for use as a bound-t was reached at 950 s. Steam then filled the upper ary condition for the SCDAP analysis. Since the portions of the hot legs, with the liquid draining cladding temperature in the two-phase region is into the reactor vessel and lower elevations of the near the saturation temperature, inclusion of the loops. The liquid in the vessel was boiled off by zircaloy oxidation energy in the upper part of the fission product decay heat until the core uncovered core would have little effect on the core dryout and began to heat up near 1800 s. At about 2550 s, time.
the peak fuel cladding temperature reached 1000 K Figure I presents the pressurizer steam dome (1340*F). The SCDAP code begins to calculate pressure. The pressure initially dropped as the reac-cladding oxidation at that temperature. Since tor power decreased faster than did the heat trans- ,
a RELAPS does not model the exothermic cladding fer in the steam generators. By 40 s, the capacity of oxidation, the temperatures calculated above the steam generators to remove decay heat had I
decreased below the rate at which core decay heat 1000 K (1340* F) are too low. The RELAPS calcula-tion was terminated at 3000 s, about 450 s after the was being added to the coolant and the RCS began 4
~. - - _ . - . _ .
l 2 i
-2600 i
j
. SRV open -2500
- ~
17
^ m D D g -
SRV closed _ -2400 m Pressurizer f u
- Core dryout . e 16 -
PORV open .
s PORV closed s e n e n e e
. L L J
o., .
-2200 a.
15 -
-210 0 14 O 1000 2000 3000 Iigure 1. pressuriier pressure for the single-channel RELAp5 TNILil' calculation.
J to heat up, resulting in a pressure increase to the liquid. Since there were no other systems or compo-
, PORV opening pressure of 15.93 N1Pa (2310 psia) nents that the operators could use, there were no i at 250 s. The RCS pressure cycled between the actions that could be taken to prevent or delay core l PORV opening pressure and the valse rescat pres- damage.
1 sure of 15.76 NIPa (2285 psia) until the pressurizer The pressure response of the two steam genera-filled with liquid at 700 s. The PORV was then tors during the first 300 s of the transient is shown
, unable to control the pressure, and the RCS pres- in Figure 2. The combination of the initial high I surized to the SRV opening pressure of 17.24 N1Pa heat transfer rates and the closed turbine stop (2500 psia) at 1000 s. The SRVs cycled seven times salves resulted in a rapid pressure increase in the between the opening pressure and the closing pres- steam generators to the secondary SRV opening sure of 16.48 51Pa (2390 psia) between 1000 and pressure of 8.87 N1Pa (1287 psia) by 5 s. The sec-2250 s. The second cycle of the SRVs, beginning at ondary SRV closed at 10 s when the steam genera-
, 1100 s, was noticeably longer than the others tor pressure reached 8.18 N1Pa (ll87 psia) and was because of a high steam generation rare from ihe not challenged again during the t ransient, as seen in l boiling in the core during this period. The steam the long-term pressure response show n in Figure 3.
generation rate dropped below Ihe PORV capacity Pressure relief was provided by the NI ADVs, whose after 2400 s and the primary system pressure opening and closing setpoints were 8.41 N1Pa I
decreased, dropping sharply after the core uncov- (1220 psia) and 8.07 N1Pa (1170 psia), respectisely. l cred completely at 2880 s. The NIADV cycling time increased significantly Operator actions to potentially mitigate the tran- after the steam generator secondaries dried out at sient or delay core damage were seserely limited by 250 s and 350 s foj loops B and A, respectively,
- he boun_dary conditions of the sequence (no ac because of the too of liquid available for boiling.
- power and no feedwater). The PORV could be The liquid masse in the steam generators are pre-j opened by Ihe operators during the transient. How. sented in Figui-
- 4. The difference in dryout times '
, ever,it was already open most of the time until after between the two loops was the result of erroneously g core damage had occurred, llecause of the small specifying different closing times for the main feed-capacity of the PORV, opening the valve earlier in water valves. This error was corrected in subse-the transient would not be sufficient to depressurize quent calculations. The A-loop secondary mass the RCS and allow the core flood tanks to inject was higher as a result of the different closing times
, 5 I
9- a 1300 SRV open A A-Loop O B-Loop
-1250 8.5 -
n WADV open
^
g O
0.
A w _
-1200 m 2 c.
WADV closed )
8 - -
I
-115 0 y l
$ SRV closed y e e L L
- 0. -
-110 0 n. l 7.5 -
l N -1050 1
t t 7
0 10 0 200 300 Time (s)
Figure 2. Steam generator pressures for the first 300 s of the single-channel RELAPS ThtLil' calculation.
9 i -1300 WADV open A A-Loop O B-Loop
-1250 8.5 ,
n l
^
- O l
~ '
. 91200 2
v v
Q.
e e
' O~ .
. -115 0 3 3 y WADV closed y e o L L
- n. . -110 0 n.
7.5 - ^-100P 8'8 * "'
i generator dry i B-loop steam -1050 generator dry t t 7
0 1000 2000 3000 ligure 3. Steam generator preuures for the single-channel REl.AP5 TNILil' calculation.
I r
}
J 6
15000 i i
! )
A A-Loop 3oono O B-Loop j -
-25000 10000 -
n
- 'Q v
.x 20000 v
}
' E -
15000 m o
2 @
3
- 5000,- -
-10000 5000
, m - ,* * *
'O O 200 400 600 Time (s)
Figure 4. Steam generator liquid masses for the single-channel RIILAp5 TMLil' calculation.
by approximately 1360 kg (3000 lbm), or about hot and cold leg fluid was the driving force for loop loro of the initial inventory of 14061 kg natural circulation.
(31000 lbm). This mass difference had a small The hot leg mass flow rates are shown in effect on the timing of the events after the steam Figure 7. Natural circulation was established fol-generators dried out. lowing the RCP coastdown, approximately 200 s Figure 5 presents the pressurizer collapsed liquid after transient initiation. Ilot leg fluid saturation in lesel. The lesel followed the primary system pres- conjunction with soiding in the top of the hot leg sure by decreasing as the primary system cooled interrupted natural circulation, and the lack of a ,
prior to 45 s and then increasing as the primary temperature difference between the hot and cold icg system heated up and expanded until the pressur- fluid prevented its reestablishment. Cycling of the izer became liquid solid at 700 s. The level was rees. SRVs following i100 s produced normal flow in the tablished after 1200 s when vapor generated in the A-toop hot leg and reverse flow in the B-loop hot core by boiling was draw n into the pressurizer. The leg as coolant flowed toward the pressuriier.
spikes in the level following 1200 s corresponded to Figure 8 shows the mass How rate at the inlet and the cycling of the SRVs. Vapor was condensed as out!ct of the core bypass. The flows were in the the pressure increased w hile the SRVs were closed, normal (positive) direction for the initial i100 s of raising the level; liquid flashed as the pressure the transient. When natural eirculation in the loops dropped after the SRVs opened, lowering the level, ended at i100 s, the flow reversed in the core ;
The pressurizer lesel decreased steadily after the bypass. The liquid in the bypass channel was cooler j SRV cycling ended at 2250 s as the liquid drained than that in the core, and the density difference l
- back into the hot leg. produced flow from the bypass into the core inlet !
) Figure 6 presents the hot and cold leg liquid tem- w hile pulling warmer water from the upper plenum peratures for both loops. The difference in cold leg into the bypass. The density in the core and core
- temperatures during the initial 500 s of the tran- bypass equalized by 1800 s, and the bypass flow sient was caused by the difference in steam genera- was essentially stagnant for the balance of the tran-tor dryout times discussed previously. The hot leg sient.
fluid was subcooled until 950 s. The cold leg fluid The soid fractions in the reactor vessel upper reached saturation at i100 s,150 s after the hot leg head, upper plenum, and core are presented in Fig-fluid. The early temperat ure difference between the ures 9 and 10. The upper head and upper plenum i
7
12 i
- -35 10 -
^
E
. 30 C -
v SRV open v q8 - -
-25 g
e e 36-- j - -20 3a a T T -
J - -15 -3 4
- -10 2
0 1000 2000 3000 Time (s)
Figure 5. Pressuriier collapsed liquid lesel for the single-channel RELAP5 TNILil' calculation.
{ 640 i i i
m 620
- -I n v
x -
650 b- v 8 e
' L 3 h a E 600 -
g
' u
- e 600 h
e A A-Loop hot log
{
e I-
- 580 O B-Loop hot leg -
O A-Loop cold leg U X D-Loop cold leg 1
' ' -550 560-0 1000 2000 3000 Time (s) l'igure 6. Liquid temperatures in the hot and cold legs for the single-channel RELAP5 TNILil' calculation.
1 1
1 8
1500 g , ,
A A-Loop 3000 0 0-Io:p m m en
" N
\ 1000 - -
E j
v
- 2000 .o SRV cycling V e End oi purnp p-g coastdown #
' 500,- -
1000 '
y 3
- r i
.* \0 b
1 C o
0-- h e 0 0 0 m m
2 0 2
-500' ' ' - -1000 0 t000 2000 3000 Time (s)
Iigure 7. 110 leg maw flow rates for the single-channel REl.APS TNil.II' calculation.
1500 , ,
A Inlet - 3000 0 Outle t
^ 7 1000 -
o> E
.x v () - 2000 .o v
C
- = 9 O +-
0 500 - -
1000
)
o Natural circulation ends y C O C
m "- -'"l M 0-- + - -. zs o m m- n.J '
F- " ~ T ~" TI " "' E3 2
-500 ~ ' ' - -1000 0 1000 2000 3000 I igure 8. Niaw flow rates at the core bypass intet and outlet for the single-channel REl.AP5 TNil.It' calculation.
9
1 , ,
5l 7;C =n n n .
0.75 -
I g -
C O
O o
O l M Height above 1 0.50 -
core outlet T A Upper head I
'o O 2.9 m (9.5 f f)
> 0 1.8 m 6.0 ft m 4 X 1.0 m 3.2 fi 0.25 -
0 0.3 m (1.0 ff)
~
0 0 1000 2000 3000 Time (s) l'igure 9. Void fractions in the reactor sessel upper head and upper plenum for the single-channel Rlil.AP5 Thll.ll' calculation. l
E E 1 i '
l qi$
A Node 1 O Node 2 O Node 3 X Node 4 0.75 - _
C Node 5
+ Node 6 f
= l o l l 0
& 0.50 -
? b o 1 9 0.25 -
0 O 1000 2000 3000 Time (s) l'igure 10. Void fractions in the core for the single-channel RI!!.AP5 TN11.It' calculation.
10
above the hot leg centerline [1.3 m (4.2 ft) abose analysis are compared with those of the single- l the top of the core] soided rapidly after the liquid channel analysis below. _ l reached saturation near 950 s. Liquid draining The RCS pressure response, presented in back into the vessel from the hot and cold legs Figure 13, shows good agreement between the cal-maintained a relatisely comtant soid fraction in the culations. The difference in response prior to core until the loops completely drained near 1800 s. 1500 s was caused by the difference in the steam Afler the loops drained,Ihe remaining liquid in the generator secondary side liquid inventory during
- core boiled off rapidly, producing the top-down the transient. As discussed in the previous section, i dryout of the core shown in Figure 10. the A toop steam generator secondary had about Anhow n in Figure iI, the fuct rod cladding sur- 10% more mass than desired in the single-channel face temperatures increased rapidly following dry- calculation. Figure 13 shows that the pressure led
- out, with the highest temperature calculated to the calculated RCS pressure from the single-1 occur 3.1 m (10.0 ft)abose the coreinlet. Theelad- channel case until 1500 s. Iloiling began in the core j ding temperature was calculated to exceed the initi- at 1500 s in both cases, and the two calculations ation temperature for potentially significant produced nearly identical primary system pressure
/irealoy oxidation [1(XK) K (1340 F)] at about responses between 1500 s and 2420 s. At 2420 s, 2550 s at the hot spot. Cladding temperatures cal- the three-channel core calculation showed that the l
eulated by REl AP5 are less than those calculated primary system would repressurite to the SRV by SCDAP after a temperature of 100t) K (1340 F) opening pressure of 17.24 AlPa (2500 psia), result-
) has been reached because the heat generated by the ing ia an additional cycle of the SRVs not caleu-oxidation reaction is not considered in REl.APS lated with the single-channel core model. The calculations. The REl.AP5 calculation was contin- difference in the initial steam gcaerator liquid ued beyond 2500 s to better define boundary condi- imentories may hase contributed to the additional tions needed for the SCDAP core damage analysis, e>cle. The primary system pressure increased The core heatup resulted in superheated steam in slow ly in t he t hree-channel core case after t he S RVs j the primary loops, anhow n in Figure 12. The loop closed at 2550 s. The pressure decreased after core heatup was anmmetric, with the highest tempera. dryout occurred at 2650 s in the three-channel core ;
tures occurring in the A-toop hot leg, the path case,230 s earlier than in the single-channel case. i between the core and the open pressuri/er PORV. The earlier core dryout was caused by the smaller l initial steam generator liquid imentory. At 2850 s, 1
a rapid pressure increase to 18.3 N1Pa (2650 psia)in Three-Channel Core. An analysis was per- the three-channel core case was caused by water in formed to determine if simulating two-dimensional the loop seals draining into the vessel and partially flow effects in the core would significantly aber Ihe quenching the core. The high sapor generation rate calculated plant respome to the TNil.II' sequence. from the partial quench of the core briefly exceeded A REl.AP5 calculation of the TNil 11' sequence t he SRV eapacity and resulted in the observed pres-
! was repeated, replacing the single-channel core sure increase. Clearing of the loop seals is discussed with a model containing three paralid channels in the section on calculational uncertainties.
l (Case I-lb). Each channel represented a separate Cladding surface temperatures from the single-
- radial region of the core with the fuel model in the and three-channel core cases are presented in
! channel using the corresponding a erage fuel Figure 14 for the core bottom, midplane, and top.
assembly power for that radial region. The center 1he three-channel core case results are from the j and middle channel were at about the same radial highest power (Region I) core channel. Prior to power factor (1.10 and 1.07, respectisely), with the cladding temperatures reaching l(XX) K (1340 F),
i outer channel at a lower aserage radial power factor at 2500 s in the base case and 2630 s in the three-I (0.72). The three channels were split asially at the channel core case, the cladding temperature
- same slesations as Ihe single-channel model. Ihe respome was similar. The heatup rates for the ihree-j parallel channels we e connected with crossflow channet core case were somew hat tower than for the j junctions between s ilumes at the same elesation to single-channel case because of steam recirculation i approsimate a two dimensional model of the core, in the core. As shown in Figure 15, the core outlet 4 N! ore details of the three-channel nodalization are flow in the Iwo center regions was upward with flow j contained in Appendix C. The calculation was in the outer region downward after the core began
] allowed to run to the same transient time (3(XX) s) as to uncoser at about 1750 s. The flows were driven i the single channel calculation. The results of this by the fluid density difference between the hotter i
i 1
1 11 i
l - - - _ - . - _ - _ . _ _-
1500 , ,
1 A Node 1 )
- O Node 2 ;2000 O Node 3 -
X Node 4 '
n l
- Q C Node 5 l'-
v + Node 6 v 3
e ',
~ 1500 ,
y 2
} 1000 D 1 L L !
e e a a E
1000 h ;
l-- '
C E E E -- u O II^-[_,
[
i , 500 500 0 1000 2000 3000 !
Time (s)
I'igure 11. I uel rod dadding surface teinperatures for the singfe-channel RiiLAP5 TAfI.II' caleufation.
900 , ,
O A-Loop hot leg O B-Loop hot leg 1100 0 A-Loop cold leg X B-Loop cold leg s
^ m 6 goo' . -
1000 vb {
, e e
, u L 1 3 '
9C0 ?
I fw 0 L
m e l C.
E 700'- -
s00 E e
s t--
__ r, -
700 lQO3 5 77~E '
=
i 600 0 1000 2000 3000 Time (s) l-igure 12. Vapor temperatures in ihe hot legs for the single-channel RiiLAP3 T All.it' calculation.
i 12
... ~ - . . ._ - .-
l l
20 , ,
A Single channel core l -
2800 O 3 channel core i
n ^
( )- U gn 18 .-
2603 -
i v
2 v
e ,
' L
' U 2400
- L 16 - -
1 Q.
Loop seats 2200
! clear 14 i 0 1000 2000 3000 l nme (s) i l'igure 13. I'rewuii/er prewure for the single- and three-channel RI:I.Al'$ TNil It' calculations.
1500 , ,
Single channel core
- - A Dottom 32000 O Midplane O Top
^ 1250 -
3 chonnel core
~
^
5 X Bottom D L
O Midplane
+ Top
)) 1500 '
u l
3 3
] 1000 -
D
- l L
i e 5-i Q-E - -
-1000 E l
- 750 -
d' t-
[% 5 [ <
',X j
, ll_-
.. - [ "m"" - o -m Z
! 500 ' ' , 500 j 0 1000 2000 3000 i Time (s)
J I~igure 14. l'uct rod cLiddmg surface temperatures for the single and threewhannel Riii.Al'5 Thll ll' calculations, i
e i
13
_m.-. -
_ . . _ - . - _ , - -r-- -,m., . - - . _ - - . . - . - -e - - - , - --,----r-- - - . - - - - - -, - - - --
1000 ;g , ,
A Region 1 2000 O Region 2
^ O Region 3 7
- 1500
.y .O v 500 -
- 1000 v u 500 u i
Loop seals clear lI l'
1 o l%l *O L .rt _ .m n 0 ~ my wuNQ 0 t;-
m Core uncovery ,
M
) o h 2 -
-500 y I -
' ' -1000
-500 i 0 1000 2000 3000
- Time (s) l'iyure 15. Core outlet mass fhm rates for the three channels for the three-channel RIil Al'5 TNil.ll' calculation.
inner core regions and the relatisely cooler outer shell and relocation of material, and fission prod-core region. 't he cooler steam flowing from the uet release.
outer region to the center kept the center channel temperatures lower than the I-D calculation tem-peratme. 'T he flow inercase at 2S50 s resulted from the loop seal clearing, w hich caused inercased boil- Table 2. Sequence of events for the ing in the core. As shown in l'ipure 14, the quench SCDAP calculation of the TMLB'
^
was nearly core. wide, cooling all clesations sery sequence rapidly. I he quench occurred, howeser, after clad-I ding temperatures had exceeded the 1000 K (1340^1hladding oxidation temperature; the addi- Time tional energy from this reaction was not included in _. _ _ _ _ Even t (s) 1 the calculation. The calculation does illustrate the Calculation begins 2500 !
l sensitisity of the loop hydraulies during the core heatup and the ef fect changes in loop conditions I uel cladding oxidation begins 2550 i
can base on the core response to the transient I uel cladding ballooning 2750 t
SCDAP Analysis. The SCDAP computer code g:uel cladding failure 2840 was used to calculate the core damage portion of the i Nil.ll' transient (Case I-le). Iloundary con- Steam starsation obsersed in upper core 3100 l
i ditions for the calculation were prosided by the single-channel REl.APS cateulation. 'lable 2 con. ZrO3 rupture 3550 l
l tains the sequence of esents for this part of the tran- 3550
/r.U-O relocation uent, l'he SCDAP calculation was begun at 2500 s.
Iise aspects of the core damage will be discuwed: Dryout at bottom of fuel 4100 cladding sut face temperature, cladding ballooning .
and failure, cladding osidation and the resultant Calculation terminated 4600 i
hydrogen generation, rupture of the cladding oside 14
1 Table 2 shows that the core dryout time was fuel bundle cross-sectional flow area is presented at 1200 s later in the SCDAP calculation than in 113 the same axial elevation s used in Figure 16. The RELAPS calculation. The difference was probably fuel rod ballooned to ihe maximum extent allowed the result of the different thermal-hydraulies by SCDAl* (approximately 90% flow area redue-models used in the two codes. SCDAP employs a tion). The assumption of a maximum balloon size simplified thermal-hydraulics model, which is based on results of ballooning experiments.Il assumes no interaction between a fuel bundle and The oscillations prior to ballooning obsened in the the rest of the reactor sessel and the RCS; only a curve corresponding to node 8 in Figure 17 were
- fuel bundle and the associated portion of the upper related to oscillations in the fuel rod internal pres-plenum are considered. The difference could hase sure. llecause of these oscillations, the timing, an impact on the cladding oxidation calculations, magnitude, and shape of ballooning and the timing j affecting temperatures and all other temperature- of fuel rod failure and fission prc fuct release could
- dependent parameters, such as ballooning, clad- hase been slightly affected. The cladding failed at ding failure, fission product release, and hydrogen node 8 at 2840 s because of increased stress and
! generation. decreased strength, w hen the cladding temperature j Figure 16 presents the fuel rod cladding surface was 1370 K (2010 F).
1 temperature calculated at three of the ten eleva- Figure 18 presents the cumulatise hydrogen gen-tions. The cladding temperature began to increase crated in the core. Zirealoy oxidation and the l immediately in the top 90% of the core. The heatup accompanying hydrogen generation is calculated at the bottom of the core began at 2700 s, as seen in by SCI)Al' when cladding temperatures exceed I
Figure 16, when the midpoint of node I became 1000 K (1340"F). At 2550 s, oxidation began at uncosered. The calculation was terminated at nodes 7,8, and 9. Hydrogen generation rates then 4600 s when the maximum cladding surface tem- increased as more of the zirealoy reached the perature was about 2900 K (4760 F). Significant threshold oxidation temperature. At 2880 s, the
, relocation of fuel and cladding had been calculated maximum hydrogen generation rate was caleu.
I by that time. The existing data base,6,7,8 howeser, lated. The heat generated by the oxidation reaction l has allowed SCDAl' code assessment 9 only up to was nearly 22% of the core decay heat at that time, temperatures of 2400 K (3860'F). Calculated elad- After 2880 s, the liquid level in the core had i ding surface temperatures reached 2400 K (3860 F) decreased, resulting in decreased steam mass flow
- at 3900 s. In subsequent discussions, results cateu- and oxidation rates. A lack of steam, or steam star-lated at 3900 s will be gisen, as well as the results sation, implies that little or no oxygen is available i from ihe end of Ihe calculation (4600 s). Since only at Ihe cladding surface for oxidation. At 3100 s, no
, a single aserage-power fuel assembly was modeled steam was flowing past the cladding in the upper l for this calculation, all parameters representing half of the bundle. Oxidation ceased at these eleva- i core-wide effects were obtained by muhiplying the tions as a result of steam starvation, and no heat
] calculated fuel assembly parameter by the number was generated by the oxidation reaction. If steam i of fuel awemblies in the core (205). There were, had been present, accelerated oxidation would have j however, both higher- and lower-power fuel assem- been calculated by SCDAP beginning at 3250 s
! blies in the core. Cladding failure with the accom- when cladding surface temperatures reached i panying fiwion product release and the onset of 1850 K (2870*F). As a result, cladding tempera-
! hydrogen generation would occur sooner in the tures would hase increased rapidly. Ilowever, with
! higher-power rods and later in the lower-power rods no steam present, the temperature rise slowed and i
than in aserage-power rods.10 Fission product the fuel rods did not form a thick oxide shell. The
! release and hydrogen generation would therefore effect of steam starvation can be seen at 3100 s in
, begin earlier, but at a slower initial rate, than is Figure 18 by the decreasc in the slope of the curve.
indicated by this analysis. When dryout at the bottom of the fuel was caleu.
The fuel rod cladding ballooned at 2750 s in lated at about 4100 s, oxidation ceased completely.
node 8. The effect of the ballooning on the elad- llecause of the steam-staned environment, only
, ding surface temperature can be seen in i igure 16 about 41 kg (90 lbm) of hydrogen was produced in i in the cune corresponding to node 8. llecause the the core as a result of the oxidation of approxi-i heat transfer surface area increased, heat transfer mately 4% of the cladding. I
- to the coolant increased and the cladding surface When the cladding temperature increased to
! temperature rise slowed. A corresponding decrease 2125 K (3365 F) at 3500 s at node 7, the zirealoy l
in flow area can be seen in Figure 17,in which the inside the oxide shell began to liquefy and dissolve 15 1
.-- _. . -- -~ _ .
3000 , , , .
Zr0 rupture -4000 1
m Steam starvation ' ^
M l'-
v v 2000 -
--3000 e Ballooning , e L L 3 3
-2000 E.
Q-E.
O-Zr-U-O relocation E -
E e 1000 E e j H -
A Node 1 -1000 e 2
O Node 7 O Node 8 o
0 2500 3000 3500 4000 4500 5000 Time (s) l Figure 16. Fuel rod cladding surface temperatures for the SCDAP TNILB' calculation.
l 0.03 i i i 0.3 l , r-us" . . . . . . ,
O O O O 'q Q Q Q y O Ob l u u u l
^ l ^
n u l E 0.02 - -
1 v -
0.2 v 1 g 2r-U-O relocation g )
e 8
- L L i o A Node 1 a l O Node 7 3 O Node 8 i
O 0.01, -
1.n.
-- 0.1 -
La-
! O C C C C C C C C C]
I J Bal coning '
0.00 O.0 2500 3000 3500 4000 4500 5000 Time (s) i Figure 17. Fuel assembly coolant flow areas for the SCDAP Thlt.B' calculation.
l j
l 16
60 i i i i
-12 0 ^
9v o E
.v Dryoul -
V
[ -10 0 C
I C
.9 40 - Steam starvation - 0 0 -80 L
.e c m -
-60 g h C m
o 20 -
-40 y Waximum generallon rate 5 I .
-20 $
1
- O ' ' ' '
O 2500 3000 3500 4000 4500 5000 Time (s)
Figure 18. Total hydrogen generation for the SCDAp TMLil' cateulation.
the outer portion of the fuel pellets. At 3550 s, the ted, the axial power and temperature profiles are relatisely thin ZrO2 shcIl ruptured at node 7 incorrectly calculated by SCDAP following clad-because of increased stress and decreased strength ding oxide rupture. Cladding oxidation, hydrogen at a temperature of about 2160 K (3430 F). A hot generation, and fission product release calculations mixture of liquefied fuel and cladding relocated could therefore be affected. At 4250 s, the tempera-downward; SCDAP only calculates the relocation ture at node I exceeded 2125 K (3365 F), and the of liquefied material. A step increase in the clad- fuel and cladding mixture that had relocated there ding surface temperature can be seen in Figure 16 from higher elevations remelted and dropped below at 3550 s at node I, where the liquefied mixture the bottom of the core. A corresponding increase in solidified. The rapid temperature rises seen in Fig- flow area can be seen in Figure 17. The cladding ure 16 at 3630 s,3640 s, and 3705 s at node 1 also oxide layer was subsequently breached at nine of I
resulted from relocation of material from higher the ten axial elevations, and downward flow of fuel locations (nodes 5, 4, and 9, respectisely). The and cladding eventually resulted in the relocation l resulting decrease in flow area can be seen at node I of S4% of the original tirealoy below the bottom of in Figure 17 at 3550 s and at 3630 s. At 3900 s, the fuel rods. Of the original UO2 , 0.8% had when cladding temperatures first reached 2400 K dropped below the bottom of the fuel rods and (3860 F), the cladding had been breached at seven I.6% had relocated to lower positions on the rods.
! of the ten axial elevations and 60% of the zircaloy No interactions between molten core material had melted and relocated from its original pc3ition. and the liquid in the lower plenum were modeled in Most of the relocated tirealoy had dropped below the version of the SCDAP code used for this analy-the bottom of the core, but some remained frozen sis. When the core material began relocating into at node 1. Of the original UO2 ,1.4% had been the lower plenum, steam created by this hot fuel relocated at that time. About 43% of that was and cladding mixture would enter the core. This below the bottom of the core. Ilowever, code-to- steam then could have further oxidized the clad-data comparisonsl2 have shown that the amount ding, producing more hydrogen.
of UO relocated is underealculated by SCDAP by The calculated release rate of soluble fission a factor of Irom six to seven. The power generated products (cesium and iodine) frem all of the fuel in the relocated fuel is carried along with the fuel. assemblies in the core to the coolant is presented in i Since the amount of relocated fuelis underealcula- Figure 19. SCDAP models the release of cesium l
17
2.0 i i i i e
. 0.004 0
6
?
- O 0.003
- e . ,
e e
gmM 6 n M >= \
D 1.0 -
oE 0.002 5 _D fo*J 8v k sur.i r.i .
a.
g g Fu.I dl.solullon E C"' M M M O.000 O.0 2500 3000 3500 4000 4500 5000 l
Time (s)
! Figure 19. Soluble fission product (cesium and iodine) release rates from the fuel rods for the SCDAP TMLB' calculation.
and iodine from a failed fuel rod with a burst com- fractional release rates from the fuel for each of
( ponent and a diffusion component. The burst com- these elements. Initial fission product inventories
! ponent is a model of the initial release when the are provided in Appendix B. At the end of the cal.
cladding fails. The diffusion component models culation, the fractional release of these four ele-l the release of the remaining cesium and iodine. Fig- ments was i1.4%. ,
ure 19 indicates that cladding failure and the l accompanying burst release occurred at 2840 s. SCDAPlRELAPS Analyses. Integral analyses of About 0.001 kg (0.0021bm) of soluble fission the TMLil' sequence were performed using the products and 7.1 kg (15.7 lbm) of noncondensible SCDAP/RELAP5 computer code to investigate fission products (xenon and krypton) were released the differences between separate and integral calcu-in this burst. About 4.985 kg(10.99 lbm)of cesium lations of the system thermal-hydraulics and core and iodine remained in the gap. Subsequent release damage. The SCDAP/RELAP5 code calculates of fission products from the fuel to the fuel- the system thermal-hydraulics, core damage, and cladding gap and from the gap to the coolant was fission product transport and deposition. For these highly temperature-dependent. As the temperature analyses, both single- and three-channel models of increased, more fission products were released the core and upper plenum regions of the reactor from the fuel and from the cladding. The step sessel were used. These analyses are discussed in increase at 3705 s corresponded to the time at the following sections, which the ZrO2shell ruptured at node 9. Shortly before this rupture, fuel had been dissolved by mol- Single. Channel Analysis. The single-channel ten zirealoy. At liquefaction (melting or dissolu- analysis of the transient (Case F-Id) used the same tion), SCDAP assumes an instantaneous release of model as the single-channel RELAPS calculation all the fission products from the affected fuel to the discussed presiously, but with ten axial nodes in the fuel-cladding gap. At 3900 s, when temperatures core rather than six. Three separate SCDAP com.
first reached 2400 K (3860*F),9.1% of the cesium, ponents were used to model the core, representing iodine, xenon, and krypton that was initially in the the fuel rods, control rods, and instrument tubes.
fuel had been released from the fuel in the 205 fuel The transient response of the plant using this model assemblics in the core. The code calculates identical w 11 be described. This will be followed by a 18
comparison of the analyses of the TMI.II' entrance of the liquid into the core caused a rapid sequence using separate and integral codes. removal of the stored energy in the fuel. The heat Figure 20 presents the fuel rod cladding surface transfer was so great that the pressure increased i
temperature for four of the asial nodes. The heatup abose 21.4 MPa (3100 psia), and the SCDAIV started near 2260 s. The temperature at the bottom REl.AP5 code had difficulties keeping the calcula-of the core began to increase abose the saturation tion proceeding. Accordingly, the model was temperature near 3600 s. The temperature decrease changed between 3820 and 40(X) s to produce a near 3840 s reflected a partial quench of the core, slower delisery of the liquid from the loop seats to This quench was caused by the clearing of liquid the reactor sewel. This change did not alter the from all four of the loop seals. Water from the loop overall transient behasior; it served to extend the seals and the downcomer entered the lower half of heat removal from the core and consequently the core, quenching the fuel rods. Steam cooling in reduce the peak prewure. The heatup after Ihe loop the upper half of the core reduced the cladding tem- seal clearing would be the same in either case.
peratures below 1000 K (1340'F), stopping the osi-The assumption that the rodlike geometry of the dation of the cladding. Iollowing the quench, the core was not disrupted by the partial core quench is temperatures increased at a much slower rate, alid. The quench front extended into the lower because the entire RCS was heating up, not just th half of the core. The cladding in ihis part ofIhe fuel
- reactor vessel. The heatup rate at the end of the calculation, about 0.06 K/s (0.ll'F/s), if contin- r ds had alta.med a maximum temperature of only I
l ued, would reach the melting temperature ofiirca- 1369 K (200$*F). The cladding at this temperature loy [2125 K (3365'F)] in about 4 h. Since the was not heasily oxidized and would not be embrit-I
- TMI.il' sequence awumes ac power is lost for I to tied, so that shattering of the fuel rods into a debris
- 3 h, the plant may be recosered before this occurs. bed as the quench progressed would not occur. In l The clearing of the loop seals began in the A. the top part of the core, where the cladding may loop cold leg that is connected to the pressuriier have been sufficiently osidized to be embrittled, the spray line. The liquid from the cleared loop seals cooling was by steam only, with no entrained lig-flowed into the reactor sessel and the core. The uid. The temperatures decreased more slowly than
]
i
- 2000 , , , ,
000 Loop seals clear A Node 1 O Node 3 rj O Node 8
! X Node 10 2500 m .
m M k'-
V 1500 -
- v e e 2000 L i 3 O
.s
' O 6
$. 1 1500 l 1000 g
- i H H i
-1000 l n . . . - - -
' ' i 500 1
500 ~ i 0 2500 5000 7500 10000 12500
.l Time (s) j I~igure 20. I uct rod eladding surface temperatures at nodes I,3,8, and to for the singleshannel SCDAp/ Rill Ap5 TMLir calculation.
t l 19 4
,c -.,- _ - - . . - - - -. , . _ . , , -_ ---- - ...- ..
In a quench, so that again formation of debris started and before the loop seals cleared because would not be expected. the pressurizer relief valves were open. This drew A sausage-type balloon occurred near 3425 s, flow from the reactor sessel through the A loop to extending oser the top 40% of the fuel rod clad- the pressurizer, causing it to heat up while the B-ding. This balloon caused a 57% reduction in the loop hot leg fluid stayed near the saturation tem-coolant flow area. The ballooning had no notice- perature. When the loop seals cleared of liquid, a able effect on the cladding heatup rate or on the large slug of hot sapor was driven from the core core flow. The cladding ruptured near 3490 s, initi- into both loops, causing a large increase in temper- l ating the release of fission products from the fuel ature after 3800 s. The temperature then decreased !
rods to the coolant. Over the course of the tran- as cooler vapor was leasing the core. The tempera-sient,less than 1% of the xenon, krypton, cesium, tures in the two hot legs were nearly identical for lodine, and tellurium that was initially in the fuel the rest of the transient as the natural circulation rods was released to the coolant. flow through the loops kept the fluid well mixed.
Oxidation of the zircatoy cladding began near Figure 24 shows the Il-loop hot and cold leg l 2955 s. The amount of zircaloy oxidizing contin- vapor temperatures. The cold leg fluid began to
! ued to increase until the partial quench of the core heat up after the loop seals cleared. The tempera.
occurred, reducing the temperatures low enough ture difference between the hot and cold leg fluid l
that no oxidation was taking place. As the tempera- remained nearly constant after about 6500 s,indi-l tures increased above 1000 K (1340*F), oxidation cating that the vapor natural circulation flow was began again. Figure 21 shows the total hydrogen stable. The temperatures in the A loop cold leg l produced during the transient. At the end of the were nearly the same as those in the 11 loop i i calculation, approximately 121 kg (267 lbm) of throughout the transient.
hydrogen had been generated. The average temperatures of the fuel rod clad.
- Figure 22 shows the noncondensible quality in ding at the top of the core, the A-loop hot leg noz-
! the top three volumes of the core. The nonconden- zie, and the top of the A loop steam generator tubes
! sible quality is defined as the mass fraction of the are presented in Figure 25. The steam generator t vapor phase that is not steam. The quality was t ube temperature remained near the saturation tem-( increasing before the loop seals cleared as the perature of the RCS until the loop seals cleared.
! amount of cladding being oxidized was increasing. The hot leg nonle temperature lagged behind the
! After the loop seals cleared, the small amount of cladding temperature by about 1000 K (1340'F) i hydrogen that was being generated was carried before the loop seals cleared. When loop flow was throughout the RCS by the loop flow. The loop seal reestablished after 3900 s, the three structures l
clearing also resulted in a four fold increase in the began to heat up at about the same rate, showing core flow rate, which would make h unlikely that the effectiseness of the vapor natural circulation flow in transferring the core heat throughout the the oxidation reaction would become steam-limited. RCS. At the end of the calculation, the cladding Some relocation of the control rods occurred. temperature was 1288 K (1859'F), the hot leg noo The control rods reached the melting temperature zie temperature was 1105 K (1529'F), and the of stainless steel, about 1700 K (2600'F), near steam generator tube temperature was 1182 K 3760 s. The stainlew steel then flowed down the (1668'F). The steam generator tubes were hotter control rods until it cooled enough to refreezet than the hot leg noule, esen though the vapor tem-none of this material dropped below the bottom of perature was higher at the noule than at the tubes, the core. The molten control material that had been because the tubes are much tlunner than the hot leg containc 1 by the stainlen steel, however, has a nonle, much low er melting temperat ure t han docs stainlew Figure 26 presents the RCS preuure response.
steel, and it flowed dow n into Ihe lower plenum, in The pressure increased to a maximum 5alue of the lower plenum, the energy was remosed from the about 19.61 Mpa(2844 psia) shortly after the loop molten material by the coolant. About 40% of the seals cleared as the liquid was boiled in the core. As control material in the core dropped into the lower stated earlier, the calculated peak pressure was plenum. actually higher, but the loop seal clearing was .
The cffeet of the loop seal clearing can be seen in extended (and the peak pressure consequently I the hot leg vapor temperatures presented in reduced) to enable the calculation to continue.
Figure 23. The A-toop hot leg temperature was Except for that period of time, the SRVs were able greater than that of loop 11 after the core heatup to control the preuure. The prewure decreased 20 l
__ __ g -. _ _ _ _ 4 - a s_
l 15 0 i i i i
. -300 Partial core quench l
250 r -
10 0 -
^ 200 E G .
.O O
$ . 15 0 g O O I 2 50,-
100
- 50 i
0 0 0 2500 5000 7500 10000 12500 Time (s)
Iigure 21. iiital hukogen produstion for the single thannel SCD\P/ RI.l.AP5 I Nil.II' calculation.
0.15 i i i i A Node 8 1 O Node 9 I x 0 Node 10 E
O i
h 0.10 - -
e M
C l e "O
@ 0.05 - -
, U C
1 ooo.. _ J \w - -
--==e 0 2500 5000 7500 10000 12500 1 Time (s) l'igure 22. Noncondensible quality at core nodes 8,9, and to for the single-channel SCDAP/Rl!!.AP5 'I All.ll' calculation.
i 21
i 2 00 , , , ,
A Loop A O Loop B 1400 -
- -2000 ^
- m l M v
b v
- 1200 - - 8 d
3 3
$6 - -1500 $u 1 1000
{
E E e e >
, 9- &
800 -. - -1000 Loop seals clear 600 "
. 0 2500 5000 7500 10000 12500 l Time (s) l'igure 23. Vapor temperatures in hot legs A and 11 for the single-thannel SCDAP/ RLI.AP5 D1111' calculation.
1 1
1600 , , , ,
J A Hot leg O Cold leg O cold leg 1400 i -
-2000 n 1
v x b v
g
- 1200 - -
- 4 3 3 5L -
1500 $L
[ 1000
{
E E e e j
I H
800-- Lyjgr**
h ~ ~1 h
i 600 U 0
2500 4
5000 7500 10000 12500
] Time (s) l l'igure 24. Vapor fernperatures in the Il-loop hot and cold legs for the single channel SCDAl'/Rl!I.Al'$ I Nil.II'
{ .,
calcul.tuon.
I 22 1
! 2000 , , , ,
000 l
A Clodding (Node 10)
O Hot leg (Loop A)
O steam generaior (Loop A)
, 2500 l M b v 1500 - :
Loop seals clear - V e e L -
2000 L 3
s-1
- D O l
L L I
[ -
1500 1 E 1000 - -
E e .
e F-
-1000 a
A S00' O 2500 5000 7500 10000 12500 Time (s)
Figure 25. Fuel rod cladding, hot leg nonie, and steam generator tube temperatures for the single-channel SCDAP/RELAPS TNILil' calculation.
20 , , , , !
-2800 l
18 - - --2600 m ^
U U
- n. -
2
- E
-2400 v 16 -
S 3 g
-2200 $
u
- Q, b 14 j Loop seals clear
.-2000
-1800 12 O 2500 5000 7500 10000 12500 Tim. (s)
Figure 26. Pressuriier pressure for the single-channel SCDAP/RELAP5 TNILB' calculation.
i 23
sufficiently so that the PORVs were again cycling, caused by two factors, the slower heatup rate and a and not held open, by about 4350 s. different ballooning model. The slower heatup was The pressurizer collapsed liquid level is shown in caused by the different thermal-hydraulics in the Figure 27. Some liquid was held up in the pressur- codes (discussed previe'isly), in that there was more izer when the loops drained, and it was slowly flow through the core, and hence more cooling, in boiled as the relief valves cycled. Liquid flashed to the SCDAP/RELAPS calculation. The new bal-steam as the pressure was reduced, and the open tooning model (used in the SCDAP/RELAPS cal-valves introduced hotter vapor from the hot leg to culation) tends to allow longer balloons.
the pressuriier, where it could heat the liquid. The The major difference between the two calcula-large pressure increase associated with the partial tions was the clearing of the loop seals. The clear-quench of the core resulted in the SRVs remaining ing of theliquid from theloop seals occurred just as open for about 100 s. This was long enough to boil the peak cladding temperature was approaching the remaining liquid. 1850 K (2870*F), the temperature at w hich acceler-The noncondensible quality in the containment ated oxidation of the cladding begins. With no lig-is shown in Figure 28. The quality continuously uid in the loop seals or the downcomer to inhibit decreased after the PORV first opened near 280 s the flow of steam through the loop piping, a strong because steam was being added to the air in the natural circulation flow was established with heat containment. A single volume was used to model rejection in the steam generators and loop piping the containment, and it received the flow from the and heat addition in the core. This loop flow redis-PORV and SRVs. No heat structures were modeled, tributed the energy in the system. Before the partial so that there were no surfaces on w hich steam could quench, most of the energy being generated in the condense. Condensation of steam would increase core was kept in the reactor vessel, heating the core the noncondensible quality. About 7ro of the non- and the vesselinternals. After the quench, the core condensible gas in the containment at the end of heat was distributed throughout the system.
the calculation was hydrogen. Because a much larger structural mass was then The events in this transient were nearly the same available as a heat sink, the core heatup was much as those in the RELAP5 analysis, but the timing slower than before the loop seals cleared. The loop was different. The sequence of events for the inte- piping and steam generator tubes heated up at gral and separate code analyses is contained in nearly the same rate as the core. The heatup rate at Table 3. The steam generators dried out later in the the end of the calculation indicated that it would SCDAP/RELAP5 calculation for two reasons. take nearly four more hours to reach the tempera-The primary factor was a difference in core power ture at which zirealoy would begin to melt [2125 K during the first 100 s ofIhe transient. The SCDAP/ (3365 F)] it would take almost eight more hours to RELAP5 calculation included the fission product reach the temperature at which zirconium dioxide and actinide decay power, while the RELAP5 cal- begins to melt. Since the heatup rate is very slow, culation was also able to account for the fission about 0.06 K/s (0.1l'F/s), it is expected that the ,
power. Therefore more heat was generated in the cladding would be completely oxidized, so that no l RELAPS calculation until about 100 s, after w hich relocation of fuel rod material would occur until both codes calculated the same core power. A see- the zirconium dioxide melted. This is consistent ondary factor was a larger liquid imentory (9'Jo) in with experimental esidence that has shown com-the steam generators in the SCDAP/RELAP5 plete in-placc oxidation of the cladding with heatup steady-state calculation. This difference in timing rates below 0.8 K/s (1.4 F/s).33 extended to most of the esents occurring before the The transient progression would change dramati-partial core quench; natural circulation ended cally before any fuel relocation occurred. Figure 25 520 s later, the core heatup started 460 s later, and showed that the reactor vessel nonles and the cladding oxidation began 450 s later. Core dryout steam generator tubes were at temperatures only occurred about 300 s earlier in the integral analysis 100:0 200 K (180 to 360'F) telow the fuel rod clad-than in the SCDAP calcolation because of the dif- ding temperatures. The melting temperature of ference in thermal-hydraulics models (discussed in Ineonelis about 1500 K (22405), so that the steam the SCDAP analysis section). generator tubes would melt long before the core The ballooning of the fuel rod cladding extended reached temperatures at which relocation of fuel
' rod material would begin. Similariy, the strength of oser a longer section of the fuel rods, but with a smaller flow area reduction, in the integral calcula- the hot leg noules would be diminished by the ele-tion than in the SCDAP calculation. This was vated temperatures, so that their sursival up to the 24
15 i i i '
40
^
E ,o - Top of pr...urtz.r O 3
e 30 _
> e e -
.e
? -
20 ?
3 cr 5 - 3 cr
. J J
-10 0 ' ' ' '
O O 2500 5000 7500 10000 12500 Time (s)
Figure 27. Collapsed liquid levelin ihe pressuriier for the single-channel SCDAP/RELAP5 ThlLB' calculation.
1 i e i i x
- 0.8 0
3 Cr
.e
.O
.- 0.6 -
to C
e "O
C O
o C 0.4 -
o -
Z l
l 0.2 ' ' ' '
O 2500 5000 7500 10000 12500 Time (s) l Figure 28. Noncondensible quality in the containment for the single-channel SCDAP/RELAP5 ThlLB' calculation.
l 25
Table 3. Sequence of events for the integral and separate code calculations of the TMLB' sequence Time (s)
Integral Separate Event Analysis Analyses Scram signal 0.0 0.0 Reactor coolant pump trip 0.0 , 0.0 hiain feedwater valves begin to close 0.0 0.0 Turbine stop valves begin to close 0.1 0.1 Turbine stop valves closed 0.2 0.2 hiain feedwater valves closed 2.0 2.0 Control rods fully inserted 3.1 3.1 Power-operated relief valve initial opening 280. 250.
Steam generator B dry 550. 250.
Steam generator A dry 550. 350.
Primary saturates (hot legs) 1040. 950.
Primary safety relief valves initial opening 1560. 1000.
Natural circulation ends 1620 1 l00.
Fuel cladding heatup begins 2260. 1800.
Fuel cladding temperature reaches 1000 K (1340*F) 2955. 2550.
Fuel cladding ballooning 3425. 2750.
Fuel cladding failure 3490. 2840.
Steam starvation observed in upper core --
3100.
ZrO2rupture .- 3550.
Zr.U-0. relocation -
3550.
Dryout at bottom of fuel 3595. 4100.
Loop seats clear of liquid 3820. -
Calculation terminated 10567. 4600.
26
melting temperature of steel, about 1700 K The lack of unoxidized material in the melt dur-(2600 F), is highly doubtful. Since the primary ing the core / concrete interaction would also affect coolant pressure boundary would fail well before the fission product behavior. With less material t'o severe damage of the core began, the pressure dur- oxidize, not as much heat would be generated, so ,
ing the relocation of the core would be much lower that fewer fission products would be released from '
than during the classic TNILB' sequence. The the molten material. However, less competition for lower RCS pressure would hase a significant impact the oxygen in the concrete may allow some fission on the extent of the direct containment heating dur- products to oxidize that normally would not, ing the melt expulsion from the breached reactor increasing their volatility.
vessel. The lower pressure would significantly This analysis demonstrated the importance of l reduce the amount of direct heating, leading to a performing an integral system analysis of the tran-
- smaller containment loading at the time of vessel sient. Using separate codes in a " piggyback" failure. The slow heatup rate, that leads to exten- approach, the core was predicted to melt within '
sise, if not total, oxidation of the zircaloy, would 4600 s. By performing an integral analysis of the l also affect the containment loading in that little core damage and system thermal-hydraulics, it was unoxidized zirealoy would exist in the melt to inter- seen that the loop seals cleared of liquid, leading to act with the concrete in the containment to generate a completely different accident timing. The melting hydrogen. Rather, all of the hydrogen would have of the core tzircaloy cladding imd fuel) would not been produced while the core was still intact. This be expected to occur until some 10 or 1I h after the ,
hydrogen, together with Ihe fission products transient had begun. The amount of oxidation also '
released from the fuel prior to the reactor vessel changed significantly, with all of the zirealoy failure, would have been released through the failed expected to b0 oxidized at the time of core melting RCS pressure boundary. If the failure occurred in in the integrn! analysis, compared to only 4% oxi-the steam generator tubes first, there may be direct dation at the tirile of fuel relocation in the SCDAP release to the atmosphere through the atmospheric analysis. .
dump valves. The reactor sessel nozzles will also fail, esen if the steam generator tubes fail first. In Three-Channel Analysis. A SCDAP/RELAP5 this case, the failure may be caused by melting of model with three parallel channels in the core and the piping rather than by weakening of the piping at upper plenum, approximating a two-dimensional high pressure by the time at temperature (creep _ model of those regions, was used for the analysis rupture failure). On the other hand, if the nozzles (Case F-le). The purpose of the three-charinel analysis fail first, the extent of the failure would affect the was to investigate the effects ofin-wswi natural circula-possibility of the steam generator tubes failing. A tion on the transient progression. '
failure of the reactor vessel nozzles causing loop in-vessel natural circulation is a single-phase flow to cease may result in no further heating of the steam flow in which hot steam leaving the center of steam generator tubes, so that they may remain the core flows into the upper plenum, is cooled by intact. If they are able to remain intact, the compli- the structural material there, flow s back down into cations caused by containment bypass will be the core along the core periphery (where the spe-avoided. If the failure of the nozzles is not suffi- cific p wer of the fuel assemblics is lower, and ciently large to disrupt the loop flow, however, the hence the temperatures cooler), then back toward steam generator tubes may fail. the center of the core w here it is heated again. This Analysis of the containment response resulting fl w w uld tend to redistribute the core heat, mov-from the two scenarios is beyond the scope of this ing the heat from the fuel to the upper plenum.
analysis. The Bellefonte plant has hydrogen recom- Compared to a one-dimensional (single-channel) biners but no igniters, so that there would be no m del of the core and upper plenum, the tempera-system available during the ThlLB' transient to tures calculated with this natural circulation flow decrease the amount of hydrogen in the contain- should be relatively higher in the upper plenum and ment. The location and timing of the release would I wer m the core. A slower heatup of the core would affect the concentrations of hydrogen in various be the result. There may also be an effect of thd parts of the containment, which in turn would fl w p tiern on the oxidwion of the cladding, in ,
determine whether ignition of the hydrogen would that steam and hydrogen are being recycled into the be possible after the vessel is breached by the mol- c re fmm the uppgr plenum, and on the fission ten core. pr duct retention m the upper plenum, in that higher structure temperatures may reduce the
, 1 27 _
T i I
fs i
amount of released fission products that are con- the upper plenum to the lower plenum, driven by densed in the reactor vessel. the density difference between the fluid in the core The plant transient response using the three- and the core bypass. After the two inner channels channel model was similar to that of the single- ballooned, flow was diverted around the ballooned channel calculation. The sequence of eveets for the region toIhe outer channel. Figure 29b presents the two cases is presented in Table 4. As in the single- core and upper plenum flow pattern after the bal-channel calculation, the loop seals deared, looning in the inner channels. The change in the although only three ofIhem, one in loop A tad Iwo Mow rates was significant. At 3600 s, before the in loop B. ballooning, the mass flow rate near the top of the Two natural circulation flows were establi;hed in core was about 40 kg/s (88 lbm/s) in the inner the core and upper plenum. Figure 29a illntrates channels, with a return flow from the upper the natural circulation now in the core and upper plenum tbrough the outer channel of about 30 kg/s plenum. Vapor flowed up through the two inner (66 lbm/s). At 3904 s, after the ballooning in the channels of the core into the upper plenum, accel- inner channels, the mass flow rate through the bal-erating as it was heated further. Heat transfer to the tooned region (top 4 nodes) was about 2 kg/s upper plenum structures cooled the flow as it con- (4 lbm/s) in the inner channels, and about 7 kg/s tinued to rise. A's the now approached the top of (15 lbm/s)in the outer channel. The flow pattern the upper plenum, it turned out toward the outer changed again when the cladding ballooned in the channel. T he lemperature continued to decrease as outer channel, as illustrated in Figure 29e. As the the vapor nowed dodu the outer part of the upper flow approached the balloon in the outer channel, plenum. Some of this flow continued into the outer it was diverted to Ihe two inner channels. Some core channel, and some flowed toward the middle flow was diverted back to the outer channel as the of the upper plenum again. The flow entering the ballooned region in the inner channels was top of the core in the outer channel was heated as it approached. .\ fore vapor flowed toward the outer flowed downward. This flow eventually turned channel near the top of the core, above the bal-toward the center of the core, where it was then tooned region in the outer channel, carried back to the upper plenum. So there was a The cladding temperature response reflected the natural circulation tiow from the core to the top of How behavior. Before the cladding ballooned, the the upper plenum and back, with a circulating How cladding temperatures in the inner channels within the upper plenum superimposed on it. This increased from the bottom to the top of the core, pattern was established when the liquid level The temperatures in the outer channel increased dropped into the second node from the top of the from the top of the core toward the bottom, as the core. The now returning from the upper plenum vapor was heated as it flowed down from the upper extended further into the core as the liquid level plenum. The peak cladding temperature in the decreased. Natural circulation How in the core outer channel was near the center of the core. After existed um;l the fuel rod clad: ling ballooned. the cladding ballooned, the temperatures in all The fuel rod ciadding ballooned :.t two different three channels increased from the bottom of the times. The top 40To of the cladding in the center core to the top, and middle channels ballooned near 3750 s, caus- Figure 30 show s the cladding temperatures at the ing an 81Fo flow area reduction in each of those top of the core for each of the three channels in the channels in the top 4 nodes (nodes 7 through 10). three-channel calculation and for the single-An 80r oflow area reduction resulted from the bal- channel calculation. The heatup was slower in the looning in the outer channel near 4140 s. The clad- three-channel case, resulting in a delay of about ding in the outer channel ballooned later because 400 s in the onset of cladding oxidation. While nat-the fuel rods were cookr than those in the Iwo inner ural circulation in the core existed, the heatup rate channels. The balloon again extended mer 400o of was about two-thirds that in the single-channel cal-the fud rod length, but from nodes 5 through 8. culation. After the cladding ballooned, disrupting s' These nodes were the hottest nodes in the outer channel, the natural circulation, the peak cladding tempera-ture increased about ilTo faster in the three-The ballooning changed the flow pattern in the channel calculation than in the single-channel core. While a circulatory flow still existed in the calculation until the partial core quench. This dif-upper plenum, no vapor from the upper plenum ference is close to the difference in power between was returning to the core through the outer chan- t he two inner channels in the t hree-channel analysis nel. There was flow through tiie core bypass from and the average core power in the single-channel L
, 28 l
Table 4. Sequence of events for the single-and three-channelintegral analyses of the TMLB' sequence Time (s)
Three- Single-Channel Channel Event Analysis Analpis Scram signal 0.0 0.0 Reactor coolant pump trip 0.0 0.0 Main feedwater salves begin to close 0.0 0.0 Turbine stop vahes begin to close 0.1 0.1 Turbine stop valves closed 0.2 0.2 N!ain feedwater valves closed 2.0 2.0 Control rods fully inserted 3.1 3.1 Power-operated relief salte initial opening 200. 280.
Steam generator il dry 570. 550.
Steam generator A dry 570 550.
Primary saturates (hot legs) I145. 1040.
Primary safety relief sahes initial opening 1595. 1560.
Loop natural circulation ends 1645. 1620.
Fuel cladding heatup begins 2240. 2260.
Fuel cladding temperature reaches IthX) K (1340'F) 3295. 2955.
Fuel cladding ballooning 3750-4140 3425.
Fuel cladding failure 3795. 3490.
Dryout at bottom of fuel 3525. 3595.
Control rod material relocation begins 4140. 3760. !
l
[ oop seals begin to clear of liquid 4155. 3820.
Calculation terminated 8000. 10567.
29
Top of Top of
+ upper + + upper +
plenum v plenum
%) \J %) " " \J %)
A " "
\J
"__ " _____ ,7__+ Top of + hh. + Top of. ____
. _ ,7__ _
core _(E__[ \ \ core I
, ) ,, a m ~
n m m a a JL JL 0
ir( ju a a a n pgo ,
blockage a a a a U( ju If f f f f
. Bottom, . Bottom, of core of core
- a. Before ballooning b. After inner channel c. After outer channel ballooning ballooning U98-KM225-05 Figure 29. Core and upper plenum flow patterns before and after ballooning for the three-channel SCDAl*/RELAl'5 TM Lil' calculation.
2000 i i i
-3000 A Center channel Loop seals clear O Middle channel O Outer channel
. X Single c.hannel -2500 M l'-
v v 1500 - -
e e
-2000 L i
3 s .
[ O n l
' L l $ -
-1500 [
l E 1000 -
E
, e ,
e H H
-1000 lL l"l Z
' ' ' ' -500 500 '
0 1000 2000 3000 4000 5000 Time (s)
Figure 30. Fuct rod cladding surface temperatures at the top of the core for the three-channel and single-channel SCDAP/RELAP5 TMLB' calculations.
analysis. Since the flow through the core was now war about 10 kg/s (22 lbm/s) prior to the core one-dimensional, the only difference between the quench. After the loop seals cleared the flow enter-single- and three-channel analyses after the loop ing the core was about four times greater. The loop seals cleared should be caused by Ihe difference in seal clearing did change the flow in the core bypass.
the specific power of the aundles. It is therefore Before the core quench, the flow in the bypass was expected that a slow heatup of the core would from the upper plenum to the lower plenum. After occur, leading to extensisc in-place oxidation of the quench, the flow was from the lower plenum to most of the cladding, with no relocation of fuel rod the upper plenum.
material until the melting temperature of zirconium Although the three-channel calculation contin-dioxide was reached. ued out past 8000 s, the code was not behaving During the first part of the heatup, the natural properly. An energy imbalance was observed circulation flow in the reactor vessel kept the clad. between the SCDAP and RELAPS portions of the ding temperatures much closer together than in the code; this problem was not seen in the single-single-channel calculation. When core dryout channelcalculation. Although the heatup rates for occurred in the three-channel calculation, the max. this portion of the transient are too low, the tran- ,
imum cladding temperature was i112 K (1542 F); sient should behave much as did the single-channel j the maximum cladding temperature was 1625 K c icut tion, with a slow heatup, nearly complete j (2465 F) when the core dried out in the single- xidation, and failure of the RCS pressure bound- ;
channel calculation. The recirculation of the vapor ry I ng before any core relocation occurs. )
in the core also resulted in a faster dryout of the )
core in the three-channel calculation than in the TMLB' Sequence with RCP Shaft !
single-channel calculation. The more uniform clad- Seal Leakage l ding temperatures will generally result in longer i balloons of the cladding.
The analyses presented in the previous section Reestablishment of loop natural circulation demonstrated the importance of the liquid in the w hen the loop seals cleared may have disrupted the loop seals. If the loop seals clear, the entire RCS m-vessel natural circulation flow if the ballooning heats up, leading to the likely failure of the RCS had not. The flow entering the bottom of the core pressure boundary before significant core damage 31
5 occurs. If the loop seals do not clear, the core mary system and increased steam generation in the retains much of the decay energy, resulting in a primary system, the pressure began to rise in the rapid heatup to temperatures at which the core pump seal leak calculation. The steam generator material will relocate. secondaries boiled dry earlier in the pump seal leak
, To further investigate the possibility of clearing case than in the base case because the temperature the loop seals, an analysis was performed using the of the primary coolant was higher. The higher tem-single-channel RELAP5 model in which leakage perature was the result of the sat urated liquid in the through failed RCP shaft seals was modeled pressurizer being drawn into the hot leg, where it (Case F-2). The location of the leak (near the loop mixed with cooler liquid.
seals) may tend to aid in the clearing of the loop The pressure increase continued until the PORV seals. No data were found that showed a best- opening pressure of 15.93 hlPa (2310 psia) was estimate value for the flow rate of primary coolant reached at about 750 s. The pressure decrease through failed seals on the pump motor shaft or for resulting from the initial cycling of the PORY was the time until failure of such a seal for the sufficient to rescat the valve once, but it opened a Bingham-Willamette pumps in the Bellefonte second time and remained open for most of the plant. Ilowever, data and analyses were found for transient, in the base case calculation, the initial Westinghouse-designed pumps. The analyses indi- PORV opening occurred much earlier (at about cated that the maximum liquid flow rate through a 250 s)and the valve cycled many more times than in failed shaft seal would be 30 IJs (475 gpm) at the pump seal leak case. The cycling of the PORY nominal cold leg conditions of 15.5 NIPa in the base case was governed by the expansion of (2250 psia) and 564 K (557 F). This flow corres- the liquid and continued until the saturation tem-ponds to a hole about 1.03 cm (0.405 in.)in diame- perature was reached at 950 s.
ter at each pump. Although the leak rate for a At about 900 s in the pump seal leak case, the Ilingham-Willamette pump is believed to be less pressure again began a rapid increase caused by two 5
than this value,each of the four RCPs was modeled factors. First, the pressurizer became full ofliquid; with this size leak. In order to assure conservatism second, natural circulation in the primary loops
- in the calculation, that leak rate was assumed to ceased as the cold legs reached the saturation tem-begin at transient initiation. While this is an unreal- perature.The first factor meant that only liquid (of
! istic assumption, it does establish a bounding cal- lower specific energy than steam) was available to culation for determining w hether the loop scals will exit through the PORV. The second factor meant i clear or not. that no movement of heat from the core to the Table 5 presents the sequence of esents for both steam generator tubes was possible. As a result of
.the pump seat leak case and the single-channel the loss of steam removal capability in the pressur-RELAP5 (base case) calculations. Figure 31 show s izer and the loss of heat remosal capability m the the pressurizer pressure for both the pump sealleak steam generator tubes, the PORV was no longer case and the base case. The reactor scrammed at able to control the RCS pressure. The primary sys-O s, and the RCP coastdown started. The p:imary tem then pressurized to the SRV opening pressure system depressurized during Ihe first few seconds in of 17.24 hlPa (2500 psia) at about i100 s in the both cases because the reactor coolant flow was pump seal leak case. The SRVs cycled between the large enough to remove more heat from the RCS opening and closing pressure seven times in the ;
through the steam generator tubes than was being base case, but only twice in the pump seal leak case added in the core. At the same time, the steam gen- because mass and energy were being lost through 4
erator secondaries were boiling dry because there the pump seals.
was no auxiliary feedwater. The pressure decrease The pressure began to decrease shortly after continued in the pump seal leak case about 250 s 1800 s in the pump seal leak case (about 800 s longer than in the base case because of the loss of sooner than in the base case). At that time, most of fluid through the pump seals, t he liquid in t he core had been converted to steam in At about 250 s in the pump seat leak case, the the pump seal leak case, while in the base case the saturation temperature was reached in the hot legs, core still contained enough water to allow boiling.
causing steam to be generated in the primary sys- Figure 32 presents the pressurizer collapsed lig-4 tem. Shortly afterward, the steam generator see- uid lesel from the pump seal leak case and the ondaries had boiled dry, resulting in greatly sinFle-channel RELAPS case. The lesel decreased decreased primary-to-secondary heat transfer. As a farther at the beginning of the transient because of result of the decreased heat transfer from the pri- the mass loss through the RCP shaft seats. The 32 l
_ _ _ _ _ _ _ _ _ _ _ _ _ - - . _ _ _ , _ - . _ _ , , _ _ _ _ _ . _ . _ _ . - _ _ . , _ . . _ . _ _ _ . . _ , ..__s _ _ . _ - .
l Table 5. Sequence of events for the pump sealleak and base cases
! Time l (s)
Esent Leak Case llase Case Transient initiation 0 0 Hot legs reach saturation temperature 250 950 Steam generator secondaries empty 270 350 Power-operated relief valve first opens 750 250 Natural circulation ends 900 1100 4
Safety relief valves first open 1000 1000 Fuel cladding heatup begins 1600 1800 Alaximum fuel cladding temperature 2050 2550 exceeds 1000 K (1340 F)
Calculation terminated 2729 3000
]
t l
18 . , ,
2600 17 -
-2400 m ) n 2 16 -
3#
2 O
-2200 ,
15 3 "
e 3 m a e a L 14 -
- 1 2000 I 13 -
A Pump seal look -
O Base case
\ ot leg H soturallon -1800 12 O 1000 2000 3000 Time (s) l'igure 31. prewuriter preuure for the pump sealleak and base case Rill.AP5 T.\ll.II' calculations, i
33 i
12 . .
Top of pressurizer A Pump seal leak
\ ^
A^ O Base case to -
-30 m E C
- s - -
e e 2 s-- -
20 2 33 I l 33 J J H0 l
2 -
0 0 0 1000 2000 3000 Time (s)
Figure 32, pressuriier collapsed liquid lesel for the pump seat leak and base case RELAP5 TMLil' calculations, pressuriier filled with liquid about 200 s tater in the The total leak flow through the four RCP shaft pump seal leak case than in the base case. seals is presented in Figure 35. The flow rate Figure 33 shows the cladding surface temperature at decreased when the pressure decreased and when the top core elesation for both the pump scal leak and the quality of the upstream fluid increased. The base cases. Core heatup began at about 1600 s (about large changes in flow between 1000 and 1200 s 200 s earlier than in the base case calculation), relleet. reflected liquid draining from the steam generators ing an earlier core uncovering resulting from the loss of to the reactor sessel.
Iluid through Ihe pump seals. The temperat ure increase Despite t his leakage, at the end of Ihe calculation continued until the calculation was terminated by the cach of the four loop seals was covered with about Rl!!.APS code when steam temperatures reached 0.8 m (2.5 ft) and 1.4 m (4.5 ft) of water on the almut 1500 K (220tPl% The cladding surface tempera, steam generator and pump sides, respectively. The ture reached 1000 K (13FF) in the pump seal leak major difference between the pump seat leak and base case TMI.ll' calculations affecting the loop case at 2050 s (about 500 s earlier than in the base se is w s the lack of relief valse cycling in the pump case). lleginning at about that temperature, significant seal leak case.
fircaloy oxidation occurs in the presence of steam, with accompanying heat generation. Since no oxidation cal.
culations are performed by the Riii.AP5 cale, clad. LOFW with One HPI Pump ding temperatures after 2050 s for the pump sealleak Available case and 2500 s for the base case are too low.
ihe A loop hot leg mass flow rate is shown in The results presented in the presions sections Figure 34. A steady natural circulation flow wa' showed that the TMI.ll' sequence leads to sescre established after the RCPs coasted dow n. The nat* core damage. The heatup results from the loss of ural circulation flow stopped near 900 s uhen the primary coolant through the primary relief valves cold legs reached saturation. The temporary flow with no llPI source asailable, llellefonte has llPl inercase after 1000 s occurred when the SRVs pumps capable of injecting at pressures above the opened, drawing more fluid through the hot leg to primary SRV setpoint. A series of analyses was per-the pressuriier. formed using a single-channel Rill.AP5 model to l
34 l
p-- ,-. -- -,-.,,,w,.--,---------m~-,------.m. , - - ----,.w.-r 4 , , - - - . ------m---,,-s , ,-,. ,., . -,.g-, ,,,_e ,,a - , , . - . , . - ,
1500 i i A Pump seal leak
. O Base cose 2000 3
^
n 1250 M b v
v
- -1500
. u i 3 3
- } 1000
$L j L e e
, Q. Ch.
E i
. - 1000 E.
! H -
750 -
1 gQ q ~ O a C A- O O O-
' ' 500
- 500 l 0 1000 2000 3000
! Time (s)
{ l'igure 33, l'uel rod eladding surface temperature at the top of the core for the pump seal leak and base case RI l.Al'$ TNill!' cateulations.
J i 10000 , ,
20000 i
A Q I
( -
15000 N j cn E W .c v - - -
5000 ' 10000 v
.j ,
- - o 0 e-L End of natural circulotton U 5000 L D
' o
~
-0 0- - -
M i M M i O m 2 -
-5000 0 2
' -10000
-5000 ,
0 1000 2000 3000 ikure 3t i cop a hoi se, mass ihm raie n,r ine ni i.ai>3 pump seas ieat eaguia, ion, i l l
l I
i l
}
! 35 1
- . - . - ~ , - - - . - . , . , . . . . - . - - - . - - . . . - - . - , - - . . -- -
10 0 i .
. -200 9
7 80 -
~
N N E
.o ai i .x . 150 _
v v
,? 60 -
Liquid draining from steam generators e o 8 '
(
6 3 3 O
L l .
f -10 0 l
! D - D i
o 40 -
o l w w
. ) M W
m t
O . -g ~""'"-- ~
50 m U
2 20 - -
2 Vap/ or flow
' ' 0 l
0 l 0 1000 2000 3000 l Time (s)
Iigure 35. lotalleak flow through the RCP shalt seah for the RiiLAPS pump sealleak calculation.
determine if the llPI system operating at half of adequate long-term core cooling. The core rated capacity (one llPI pump operational) would remained covered with liquid throughout both provide adequate core cooling for a 1.0FW transients, and core cooling was expected to con-sequence, tinue until the borated water storage tank (IlWST) -
Two sariations on this sequence were esaluated. that supplies the HPI system empties 20 to 25 h The lirst invoked adding the injection capacity of a after transient initiation.
single llP! pump injecting equally into the four The sequence of esents for the two transients is cold legs to the TNILil' scuuence (Case 12 3). ~1he -
presented in Table 6. The esents occurring during ilPI injection was initiated when the primary syv the initial 3.1 s are the same boundary conditions tem preuure initially reached the PORV opening assumed for the TNILII' sequence. Core heatup ,
prewure of 15.93 NiPa (2310 psia). The second sar- above the coolant saturation temperature did not iation (Case F-4) used the same conditions as the occur in either LOFW transient with IIPl. The first, but with the additional prosision that the analyses showed that a single llPI pump can pro-PORV was unasailable (either failed closed or the side adequate core cooling by 3500 s with both the upstream block sahe was closed). The only pres- PORV and the SRVs, and by 5000 s with only the sure relief for the primary system in the second var- SRVs prosiding primary system prewure relief.
lation was through the SRVs.- The primary system pressure response to the The tramients were evaluated until either a sue- sequence with PORV and SRVs is shown in een or a failure criterion had been met. Succew was Iigure 36. The pressure response is similar to that defined as the establishment of adequate long term obsened in the ThlLil' transient, with the time of core cooling, l'ailure was defined to occur if fuel the initial opening of the SRVs delayed by the cool-cladding surface temperatures eseceded the initia- ing provided by the llPI system. The SRV opening tion temperature for potentially significant clad. pressure was reached 2050 s after transient initia-ding midation. The temperature used for the lion and the salves cycled seven times, l'ollowing failure criterion was 1000 K (1340'12). 3400 s, pressure relief was prosided by the PORV, lloth transients were terminated after analyses which was full open from 900 s until the end of the indicated one llPI pump was sufficient to proside calculation.
36 i i
{
l Table 6. Sequence of events for the LOFW with HPI transients Time N
Esent PORV and SRVs SRVs Only Scram signal 0.0 0.0 Reactor coolant pump trip 0.0 0.0 Niain feedwater sahes begin to close 0.0 0.0 Turbine control sahes begin to close 0.1 0.1 Turbine control sahes closed 0.2 0.2 Niain feedwater takes closed 2.0 2.0 Control rods fully inserted 3.1 3.1 Power-operated relief salve initial opening 230 -
liigh. pressure injection initiation 230. 230.
Primary saturates (hot legs) 2050, 2200.
Primary safety relief sahes initial opening 2050. 330.
Natural circulation ends 2100. 2300.
liigh-pressure injection flow exceeds 3500. 5000.
relief sabe flow Transient terminated 4696. 11278.
Figure 37 presents ihe primary system pressure ibrough the relief vahes for Ihe tuo transients. The response to the sequence with no PORV. Ihe RCS combination of PORV and SRVs produced a more pressurized to the SRV opening pressure 330 s after rapid decrease in system mass in the transient with tramient initiation. The SRVs then cycled for the both asailablc until the SRV cycling ended at remainder of the transient. The transient produced 3500 s. IIPI flow was then able to exceed PORV more challenges to the SRVs than the sequence flow, and the primary system began to refill. In the where the PORY was available. At the end of the sequence where only the SRVs were asailable, the calculation, the SRVs had cycled i10 times and RCS began refilling after 5000 s w hen t he 11PI Ilow were cycling the rate of about 10 per hour. The exceeded the flow being expelled through the SRVs are designed for 280 cycles oser the 40-year cycling SRVs. The higher aserage RCS pressure life of the plant.I4 Thus,if the transient continues without the PORVs reduced the IIPI flow, causing 20 to 24 h until the llWST empties, the large num- the 1500-s difference in refill times between the two ber of SRV cycles could result in SRV failure. t ransients.
Figure 38 presents the change in the primary sys- The analyses show that asailability of a single tem mass imentory for the two LOFW with IIPI IIPI pump is sufficient for core cooling regardless tramients. The figure show s the difference between of PORV availability until the llWST empties the mass injected by the llPI and the maw expelled between 20 and 24 h after transient initiation. The 37
18 - i e i i -2600 17 -
HPl on L l g 1 -2400 g n- 7 l d 16 - -
3 e e L L
-2200 g E.
e is - - -
e L j L O. f Q.
14
- A- -2000 i i g i i 0 1000 2000 3000 4000 5000 Time (s)
I-igure 36. Prewurizer prewure for the 1Of W transient uith one llPI pump asailable.
18 - i i i i i -2600 I I 2400 g d 16 - -
S o l i e i l L L l 3 -2200 g e is -
l
- e i L L Q. Q.
~
14 =
. 2000 13 O 2000 4000 6000 8000 10000 12000 Time (s)
Iigure 37 Piewuri/cr prewuse for the i Of W tiansient with one llPI pump and no PORY asailable.
I 38
50000 - i i ' -
1.0 0'10'
- b. PORY and SRVs n l
9.x O SRVs only E
-Q
- O Og -- 0.00
_ O
- C
$ HPl on e c -- -1.00'10, m E -5000 E I 'o
- 2.0 0*10' as -100000 -
i
-150000 i ~M tPI flow exceeds relief' volve flow
- -3.00 10 I
0 2000 4000 6000 8000 Time (s)
Figure 38. Reactor coolant system mass imentory change for the !OFW transients with one flPI pump asailable.
comparison between the LOFW sequences with (Case F-le) codes separatcl), core damage was pre-and without the PORY available showed that the dicted to begin near 2550 s, with zirealoy relocation PORY asailability resulted in considerably fewer occurring near 3550 s. Use of a three-channel core i challenges to the SRVs and, therefore, less proba- model in a RELAP5 calculation (Case F-lb) bility of their failure. slowed the heatup of the cladding and indicated that the loop seals may clear of liquid, causing a Summary of Feedwater Transient Partial quench of the core. Integral analyses using Results the SCDAP/RELAP5 code also showed that the i loop seals would clear of liquid. A single-channel
'. core analysis (Case F-Id) showed that the natural Four i OFW sequences were analyzed. T.he tran-circulation flow through the loops was sufficient to j sients were continued until the plant was in a stable ansfer the decay and oxidation energy from the long-term coolmg mode or until core damage occurred. rhe analyses demonstrated that the core throughout the RCS, significan ty reducing
! availability of a single llPI pump was necewary e com a p rate angtenWng inc nannnt.
L, ore damage began near 2955 s, but relocation of
- and sufficient to present core damage. In the absence of IIPI, there were no actions that the the fuel rod cladding was not expcered until about operators could take to present core damage. 10 h afur the nansient began. The three-channel Table 7 summarizes the feedwater transients. analysis (Case F-le) showed a slower core heatup Injection from a single flPI pump was able to prior to the core quench, but the heatup afterwards present core damage with a total LOFW. The RCS was expected to be about the same as for the single-liquid inventory was increasing by 3500 s with the channel calculation. A RELAPS analysis of the l PORV and SRVs reliesing pressure (Case F-3), and TalLil' sequence with RCP sealleakage(Case F-2)
! by 5000 s with just the SRVs (Case F-4). Failure of resulted in an earlier core heatup but no clearing of the SRVs may result if the PORV is unavailable dur- Ihe liquid from the loop seals.
ing the transient because the number of eyeles of Seseral severe accident issues are addressed by the SRVs would be close to the total number of the analyses. These include hydrogen generation, cycles,280, expected oser their 40-yr design life. in-sessel and loop natural circulation, the retention The TNILil' sequence was the base case tran. of fission products within the RCS, and contain-sient. Using the RELAP5 (Case F-la) and SCDAp ment loading.
39
' l, s s
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The SCDAP analysis of the TMI.ll* sequence higher than thh, so that estensise melting of the l showed that only 4ro of the rirealoy in the core was fuel may occur in the reactor sewel. Thh melting
- I osidized before fue: relocation began. The would then release many of the fiwion products SCDAP/Rl!!.APS analpes indicated that nearly contained in the fuel to the RCS. The estent and (
complete osidation may occur before any reloca. timing of the fiwion product release in the sewel tion begins T he difference between the Iwo calcu. will affect the ultimate source term of the accident.
lations demonstrates Ihe importance of integral The piping temperatures were high enough at Ihe i analyses. The upper bound of nearly complete osi. end of the transient that few fiwion products will be l dation h reasonable, ghen the calculned heatup retained in the RCS. I:iwion pmduch that were !
rate and the resulh of esperiments performed with released w hile the structure temperatures were high comparable heatup rates The lower bound of 4%
would not condense on those surfaces, while rh-osidation is too low because the calculation did not sion product 5 that were released carher and had condensed on the surfaces could be resapori/cd. If account for the feedback between the core and the the steam generator tubes fail, a path through the rest of the reactor veuel and RCS. This feedback would result in more steam being asailable to the MADVs and SRVs eshh for containment bypan.
i core, increasing the amount of osidation. Sicam With thh path, and with significam release of fh-4
' sion pmducts in the reactor sewel from mohen stanation with a low total oxidation would not be fuel, Ihe source term could be s cry large. ifIhe RCS i espected. I failure h in the hot leg piping, the fiwion prodoch
'Ihe amount of osidation aho affects the contain.
may be releaud to the containment befon. the seu ment response atter the reactor sewel faih. If there is a sel faH% Thh would allow time for the released fis-large amount of unosidited materialin the meh, there uon pmduch Io be condensed or otherwhe will be a much more sigorous interaction letween the retained on surfaces in the containment, reduemg mohen core and the concrete, releadng hydrogen and the amount of airborne radioacthity at the time of fission product to the containment atmosphere. If conta nment failure and thus reducing the source
{ thewhlinlematerialtoivosidized fewer fiwion pmd.
i uets in the molten core will tv releami imm Ihe melt e ch faHureof the RCS. prior to sewel melt-
! and dispersed as aemsoh to the containment. (hida-I ihmugh wonM alu) affect the cuent of the direct tion of some fiwion products, honeser, may increase contaimnent heah,ng. Direct containment heating a their miatihty.
w heat ran er nun the ejected eme material to l Inaewei natural eirculation slowed the heatup of emu nnwnt atnanphere during the timc ol sev l the core. I nergy from the core was tramferred to set breach, if the R( S prewure h high when the j the upper plenum siructures I he onset of dadding n,g, g g.gg gg ggi
- oudation waulela>ed in Ihe tbree channel calcula- tuomi/cd as it h ejected from the sewel, prm iding a I nom by 130 s ny the Riii.APS cateulation and by g,gry ,gggggg. g ,
340 s in the St' DAP/Riil.AP5 calculation com. mem mmo@cm ud eMng a gM NM ud !
pared to the ungle channel calculations. Italloon. prewuritation ofIhe containment. If the RCS prev ,
ing of the fuel rod cladding, and the awoelated we h low, We m udd wHI m om of N W reduen,on in Ihe coolant flow area, dkrupted the into the containment, with a minimum heat tranu nalural circulatmn flow m Ihe core. Subsequently. fer wrface area and a unaller immediate loading of I the heatup would be sery similar to the single. We comainmem. The piping temperatures arc high I J
j channel analph. enough during the TMI.ll' analyses that failure of I
% kh high amount of midation in the core, the the piping h likely, w that the RLS preuure could !
fuel md matenal will not reh>eate until the tempera- be near the containment preuure w hen the sewel h l j ture reaches the melting point of tirconium diod breached by the mohen vote materialin the lower [
]
ide, about 2973 K (4M@l:). t he melting plenum. 'the direct containment heating wouki l temperature of the fuel h only about 40 K (72*l') therefore be espected to be small I I
I I
I 1
4 41
SMALL BREAK LOSS OF COOLANT ACCIDENTS 1 hree sequences initiated by a small break IDCA one 11PI pump is the only I!CC mailable and there were imestigated. The sequences awumed that sar- h no Al:W. The S,1) sequence with no operator iom combinations of ECC sptems and the Al W actiom is described in the t hird section. The opera-sptem were asailable. The first tran icnt awumed for mitigating action imestigations for the S,1) that no !!CC or Al W syst ms were mailable. sequence are then presented. I inally, the results of Another tramient awumed that only (me llPI the small break IDCA analpes are summarized.
pump was mailable, and that all the other liCC splems and the Al W sptem wcre unmailable. I he Small Break LOCA with LOFW tramient of primary interest was the S,thequence, which awumes that only the llPI sptem is unasail- and No ECC able. All ofIhese tramients awumed that the oper-ators took no actiom. Two additional analpes were in addithm to the small break initiating the tran-performed for the S.l) sequence that comidered sient. this tramient further awumed ibat all I!CC operator actiom taken to mitigate the tramient. sptems had failed or weie unmailable and that the l he break was a 5.1 em (2.0-in.) diameter hole in Al:W splem was unable to prmide water to the one of the RCP diwharge lines. Ihe break wm steam generators (Case S-1). ~1 hh was a bounding located on the bottom of the pipe. This particular analph that prosided the minimum time to eme 1 site break was chmen becau e, with no high prev damage for the small break IDCAs. Ihe calcula-sure injection,it resulk in the RCS preuure staying tion performed to imestigate this tramient used the abme t he CIT prewure w hile t he core liquid imen- Riii.AP5/Mol)2 computer code and a single.
tory b boiled off and the core heatup begim, channel core modd. f he tramient was terminated Ihe tramients were initiated f rom full power by w hen either a suceeu or a failure criterion was met.
the opening of the break. When the RCS prewme Succew was achiesed if adequate lonpterm core dropped below 13.80 MPa (2002 psia),Ihe reactor cooling was established. I ailure was met ifIhe peak protection sptem initiated reactor seram. 'I be inte- dadding surface temperature reached 1000 K grated control splem Ihen reduced main feedwater (1340^17),Ihe initiation temperature for potentially llow to near reto to present merfill of the steam sign;ficant eladding midation, generators m the reactor power dropped. I or these Ihe sequence of esents for Ihe tramient h analpes, the main feedwater wm ramped Imm full show n in table 8. I he wram signal was generated steady 4 tate flow to /cro llow mer 2 s, the stroke 29.6 s after opening the break. 'I hh signal initi-time of the main feedwater sahes. When the RCS ated dmure of the main feedwater and turbine prewure dropped to 11.82 MPa (17t$ psia), the control sabes as udl as initiating control rod emer gency core cooh ng in jection (1:CCl) sig nal w as imertion. I on RCS ptenute at 47.6 s generated generated, aethating the i CC and Al'W spterm an !!CCl signal. 'Ihe hot Icg subcooling margin and nolating the steam generators by dming the was bdow 6.7 K (12'11 at the time of ITCl, so ma n steam and feedwater i olation sabes lhe the RCPs were automatically tripped. Since llPI Al W sprem lesd demand was wt to Mode C was not mailable to make up Imt imemorb the bmalllucak indicated) by the emergene) safety lea- core dried out; eladding heat up abme Ihe splem tures actuation sptem #1 SI AS) thannel, initiating saturation temperatute began at 2l00s. Ihe a 0.Ol$ m4 0 It/ min) refill of the steam genera. dadding surfaec temperature reached 1000 K tors to 12.2 m (40 It)(More of the operating range (1340'l ) 2750 s alter tramient initiation.
lesel). Ihe RCPs were tripped by the automatic Ihe primar) splem prewure respothe h shown RCP trip protection function when ihe hot leg sub- in Iigure 39. Ihe prenure initially decreased as cooling margin dropped below 6.7 K (12"l-) t he energy remmed by the steam generators and the ,
tramient would normally terminate with core cool- break eseceded the core heat pencration. Ihe RCS ing prmided by Al Winduced ratural drculation prewure then increawd m the coolant espanded and Ir0 splem makeup of the RCS man knt following steam penerator dr)out and bulk eoolant through the break, boiling in the core. Without I:CC injection, the the first analph prewnted h the tramient in core esentuall) boiled dry, as shown by the core which no ifC or Al W h mailable. Ihe nest see- liquid leselin Iigme 40. Ihe primary splem prev tion prewnts the results of the tramient in which sure then decreawd m lew liquid was mailable in 42 l
l - _ . - - - - -- - -., . - - . - - . - - - . . _ - -
l te , , , I i
-2200 I
~
I4 - -2000 g S ~
2 c.
i
-1800 e 12
. e
! u u 3
m a
a g -
-1600 g i L L A A go _' -
-1400' i
1 -
-1200 0 1000 2000 3000 4000 Time (s) ligure 39. Pressurifer pressure for the small break 10CA with LOFW and no ECC.
1 I
5 , , ,
15 l
4 -
n m v
E O 1
v
-i e 3- . 10 -
e 0 *
- i T T j - 2 - -
3 3 7 7 4
-5 -
J J 1 .
i 0 ' ' '
l 0
l 0 1000 2000 3000 4000 1
Time (s)
I igure 40. Collapsed liquid lesel in the core for the small break I OCA with IDi'W and no !!CC.
l 43
Table 8. Sequence of events for the small the ECCI signal was generated. No other ECC sys-break transient with LOFW and tems or feedwater systems were available.
no ECC The initial RELAP5 analysis of this transient (Case S-2) produced some numericalinstabilities in the region of the break. The instabilities were the Time result of subcooled flPI liquid and steam from the Event (s) core mixing in the volume upstream of the break and producing rapid changes in the flow regime.
Transient initiated 0.0 The flow regime switching changed the interphase j mass transfer and void fraction and caused the Scram signal 29.6 break flow to oscillate. The oscillations in the break fl w pr duced peaks in the flow three to four times hiain feedwater valves begin to close 29.6 ,
higher than would be expected (based on the homo- 1 Turbine control valves begin to close 29.7 geneous equilibrium model and upstream stagna-tion conditions from the RELAPS calculation).
Turbine control valves closed 29.8 Since all the llPI injected into the broken loop was spilled to containment, ilPI was turned off in the hiain feedwater valves closed 31.6 broken loop. The result was a smooth calculation of break flow with a RCS depressurization rate 4%
Control rods fully inserted 32.7 slower [to a pressure of 6.9 hlPa (I000 psia)] than with flPI in the broken loop.
Reactor coolant pump trip 47.6 The sequence of events for this transient is ident,-
i Fuel cladding heatup begins c I to that given in Table 8 up to the time of the 2100.
ECCL signal at 47.6 s. IIPI was initiated at this Fuel cladding oxidation beginsa 2750. time with a single HPI pump injecting (1/4 of the flow dumped to containment, the balance injected Transient terminated 3450. equally into the three intact cold legs). The calcula-l tion showed 11PI injection exceeded the flow out the break by 10,800 s. The llPI system with only
- a. Iwentially significant cladd.ag oxidation is awumed I the borated water storage tank as a source can pro-begin w hen the peak cladding surface temperature reaches Hxc x (190 t). vide 15 to 20 h of injection at RCS pressures of approximately 5.5 to 6.9 hlPa (800 to 1000 psia).
I Figure 43 presents the RCS pressure. The pres-the core to be boiled. The break flow, shown in Fig- sure increased when the system reached saturation l
l ure 41, initially became single-phase steam at and boiling of the coolant began, then decreased 500 s, with short increases at 750 and i150 s result- steadily. The pressure decreased more rapidly fol-l l
ing from liquid draining out of the lower portions lowing break flow transition from two-phase to of the steam generators into the cold legs. The pri' single-phase steam flow near 4500 s. Ily 15,000 s, mary source of mass to the break was from the core the RCS pressure was 4.69 hlPa (680 psia) and through the reactor sessel internal sent valves. As decreasing at 0.59 hlPa/hr (85 psia /hr). Extrapo-shown by Figure 42, the transient produced a top- lation showed that the CFT injection setpoint pres-down core dryout and fuel rod cladding heatup sure of 4.24 N1Pa (615 psia) would be achieved by with the peak cladding temperature reaching the 18,000 s. The corr. liquid level (Figure 44) began to failure criterion of 1000 K (1340*F) 2750 s after increase 10,800 s after transient initiation, indica-t ransient initiation. ting that adequate core cooling had been estab-lished. Fuel rod cladding surface temperatures Small Break LOCA With LOFW "'*"i"ed at r below the saturation temperature And Partial HPl throughout the transient. The llWST supply at an ;
The small break transient in this case differed RCS pressure of 4.24 StPa (615 psia)is about 15 h J from that described in the presious section only in and, with accumulator injection predicted to occur the assumed ECC s> stem failure. One-half the approximately 5 h after transient initiation, ade- !
rated ilPI injection capacity of the plant was quate core cooling should be achiesable for that 1 assumed available, with injection initiated when time frame.
l 44
i 250 i . i
- -500 m
9 200 p
(
- -400 E x e Liquid draining from stearn generators v e 150 l 1 3' - I I -300 *-
h 100 -
A -
E D
-200 50 -
30 0 Vopor flow 0 O O 1000 2000 3000 4000 Time (s) l'igure 41. .\ lass flow rate through the break for the sinall break IDCA with IOl'W and no !!CC.
1600 . . .
A Node 1 O Node 2 1400 ,- O Node 3 ~2000 X Node 4 g x 0 Node 5 v 1200 -
+ Node 6 -
L e
- g -
1500 u
] 1000 3
L U e 5-
- c. *
- -1000 "
E. 800 -- E
- O l
600 H :: -
0 - = m a 3 -
-500 400 O 1000 2000 3000 4000 Time (s) l'igure 42. l'uel cladding surface temperatures for the small break IOCA with IDI W and no I:CC.
45
20 i i
. -2500 15 -
m
^ .
HPI on 2000 0 S I M 2 P Q.
v v e
6 10 [ 1500 e u
1 3 3 I En M M M
- . -1000
- u 6
- 0. 0-5 - 2
. -500
' ' O 0
O 5000 10000 15000 I igure 43. Pressurinr pressure for the small break I OCA with LOI W and one lil'1 pump asailable. ;
1 4.2 e i
. - 13.5 1
4
. - 13.0
^ C*
E v v
~~ I2*'
! q 3.8 -- e e HPl flow exceeds brook flow ,
e
~
. 12.0
! _- - . m -
- - - -- ^-
- 11. 5 a a
~
3.4 -
. 11.0
' ' 10.5 3.2 0 5000 10000 15000 Time (s) l'igure 44. Collapsed liquid lesel in ihe core for Ihe small break i OCA with I 01 W and one itPI pump asailable.
46
- . . _ . - . - - . _ _ .. _, _ _ _ _ - _ _ _ _ _ _ _ ~ _ . . _ _ _ . _ . _ . _ , - - - _ . _ - . - . _ . . , - . . . . . - - -
S2D Sequence Analysis sure when the pressure dropped below 13.80 N!Pa (2002 psia) near 25 s. The RCPs tripped at $$ s on low RCS prewure coincident with a lack of ade-1 he S,D sequence (Case S-3) was considered the quate subcooling in the hot leg. Ihe emergency base case transient for the small break 1.OCA anal- safety features were actuated near 55 s. The steam
> ses it is a risk, dominant sequence for prewurized genehators were fed by AI:W to attain and maintain water reactors.- The S,D sequence assumes that a liquid lesel of 80% on the operating range inea-there is no IIPI during the transient, either from t he surement [about 12 m (40 ft) abose the bottom IIPI system or the makeup sprem, and that there tube sheet). Ileat remosal from the RCS continucd are no operator actions.
in the steam generators until the RCS prewure fell
'I"" "* "#"* generatm pNwuN at a ut m s REI l' ari . 'I I / !! . 'S o i ut r d-at u ral m.ulan.on dow duvugh theloops was not 4 9: cc the RELAP5 calculation was scry similar to
'## hi d heause de loop seats dM not clear of the first part of the SCDAP/REl.APS calculation, only the latter uill be presented.
h.'!a quid during the transient. The core began to heat up near 2(MM) s. lleginning at 3150 s, water draining Table 9 presents the sequence of esents for the transient. The reactor scrammed on low RCS pres- from the loops held the core temperatures constant Table 9. Sequence of events for the S2 D transient Time thent (s)
Iransient initiated 0 Reactor scram 25 Nfain feedwater pumps tripped 27 Reactor coolant pumps tripped 55 Emergency safety feature actuation system signal generated 55 Auxiliary feedwater flow initiated l15 flot legs reached saturation 120 Core heatup began 1950 Reactor coolant system prewure below steam 5190 generator prewure l'uct rod cladding oxidation began $650 Cladding ballooned $900 Cladding ruptured 6020 llottom of core dried out 6390 Core flood tank injection began 6630 Control rod relocation began 6800 Calculation terminated i1900 47
until about 5000 s, when the core temperatures steam generators to the reactor s ewel and eore. l he began to inercase again. Oxidation of the cladJing lesel decreased steadily from 1500 to 2300 5. Oser began when the temperature reached 1000 K the rest 27(X) s (until about 5(XX) s), liquid draining (1340al). The cladding ballooned and ruptured, from Ihe loops to Ihe reactor sewel caused the core initiating the release of finion products from the liquid lesel to remain fairly constant. When the fuel rod gap. Shortly after 6600 s, the RCS prewure draining was completed, the core lesel began to had dropped far enough that the core flood tanks decrease again as the liquid was boiled by heat began to inject liquid into the sptem. The stainless transferred from the fuel rods. The core was com-steelin the control rods began to melt and relocate pletely dry by approximately 6400 s. Small incur-near 6S00 s. The core heatup continued to the end sions ofliquid into the core after that time were the of the calculation at 11,900 s. result of CFT injection.
Iigure 45 shows the RCS pressure during the The fuel rod cladding temperatures in four of the transient. The pressure decreased rapidly to the sat- ten asial nodes are presented la 1igure 47. A top.
uration prenure corresponding to the hot leg tem- down core uncosering led to heatup of the core perature at the beginning of the transient. As the abose the saturation temperature beginning near liquid began to boil, the break flow was not able to 1950 s. This initial heatup was slowed and then compensate for the espansion of the coolant stopped by liquid draining back into the reactor caused by the boiling, and the pressure inercased. sessel from the loops, the temperature behmior T he three rapid depreuuri/ations between 500 and paralleled the core liquid lesel behas ior seen in I ig-1200 s were caused by large increases in the break ure 46. When thelesel began to decrease again near flow. The increases occurred when the fluid 5000 s, the temperatures began to inercase. The upstream of the Srcak changed from mostly sapor heatup of the top half of the core slowed when the to mostly liquid as liquid drained from the steam eladding ballooned near 5900 s. The ballooning generator to the reactor sewel. The RCS prewure resulted in inerrased heat transfer from the clad-then remained fairly constant near H.2 Mi a ding. the increased surface area of the cladding 11190 psia) from 1200 to $200 5 w hile much of the after the ballooning occurred increased the heat liquid in the RCS was boiled. Afier $200 5, the sol-transfer rate. The coolant flow area reduction awo.
umetric flow out the break was greater than the ciated with the ballooning caused the steam seloe.
coolant espamion caused by the boiling, and the ity to increase, which in turn increased the heat prewure began to decrease again. The steady transfer coefficient between the cladding and the decrease in the preuure stopped after the Cih coolant l'he initial injection of CIT liquid near began injecting liquid to the RCS near 6630 s. 6630 $ interrupted the heatup ofIhe bottom half of When the injected liquid entered t he core, it boiled, the core. The heatup of the top half of the core raising the RCS pressure. ihe inerceed prewure slowed near 7000 s because the thickness of the osi-terminated the injection, and the prenure di/ed eladding uas increasing. I he osidation rate is decremed again. When the prewure decreased low imersely proportional to the thickness of the oxide enough again, more liquid wu injected. This la>cr through which the os> gen must pass to get to e> cling of the Ci f injection slowed the depreunri-the unosidi/ed /irealoy. The cladding temperature
/ation. Molten control rod material dropping at the bottom of the core returned to the saturation below the core aho slowed the depreuuri/ation. temperature seseral times between 6000 and Ihis material boiled liquid in the lower plenum, 10,000 s when liquid entered the core as a result of causing the preuure to increase. The preuure CIT injection. Molten control rod material falling increase twar 10,500 s was caused by material relo-into the lower plenum caused a teinperature eation to the lower plenum. T he combination of Cl:1 injettion and a low break flow rate resulted in deetcase in the bottom 40% of the core near 10,500 s. T he steam enerated by the interaction of the RCS picume still needing to decreme another the molten material with the lower plenum liquid 1.4 MPa (200 psi) at the end of the calculation wm cooler than the steam that had been entering before ! PI could begin.
the core, so the fuel rods were cooled. Ihk steam Iigure 46 presents the collapsed liquid lesel in al.o swept out some of the hydrogen that had been ihe core. Steam we generated in the core wit him the building up in the core. At the end of the calcula.
first 200 s of the tramient. The increases in lesel tion, the peak cladding temperature was 2963 K between 700 and 1200 s corresponded to the decreases in preuure during that time bec (4M74T). firrors in the subroutine that calculates finion product releme from the fuel presetited the Iigure 45). Iloth rellected water draining from the calculation from proceeding further. The melting 48
. - . _ . . - ~ . . - - .-. . - - . - - . -. - .- -. . . _ - _ . . - . __ -
20 i i i i
. 2500 15 n
, n ~ 2000 O i O =
i n. #
r 3
v v c.
u 3
10 .[ . 1500 e L
3 m #
M #
e - 1000
- u u i n. G-
~
l 5 -
. 500 CFT Injection beelns 0 0
, 0 2500 5000 7500 10000 12500 i
Time (s)
I'igure 45. Prenuriier prenure for the S2D tran ient.
f 1
1 5 i i i i 33 l 4 -
n e i
1 E
((
- i 1 10 -
e 3-. .
> e e >
- .e i
m 3
2 - -
.m U"
3
- U" 5 -
i .J J 1 -
1 l ' ' '
0 ' ' '
0 1 0 2500 5000 7500 10000 12500 i
- Time (s)
Iigure 46. Collapsed liquid leselin the . ore for the S D transient.
2 1
i
- 49 I
~ , , . - - , _ . - - . - , - - - - , - , - , , . _ - - - - . - - - -_ ..--,-.,---,..-_-n-,,. - . , - . - --,_ . ~ - - - . - - - - . . - , - , , - . - - ~ . . .
3000 , , , ,
l A Node 1 O Node 5 2500 - -
O Node 8 X Node 10 N 000 l A A I M S-v v l
l
- 2000 -
4 -
- I 3000 g
' *- Conitof rod +-
O relocalion U
! ( 1500 -
2000
{ i E E
- CFT Injection
- SaltconIng b 1000 - -
1000 fHI = m . . !
=" "' " " " ' ^
500 "~ ..
0 2500 5000 7500 10000 12500 Time (s) l'igure 47. I uct rod cladding surface temperatures at nodes I,5,8, and to for the S2 D transient.
temperature of zirconium dioxide is 2973 K tures at four of the ten axial nodes. The radial tempera-(4892*F), and that of uranium dioxide is 3013 K ture profile is scry flat across the control rod (4964'F); these temperatures would base been component so that the rircaloy guide tube temperature attained shortly after ihe end of the calculation. is nearly the same as that of the stainless steel cladding Despite these high temperatures, no fuel rod mate- and the control material. When the stainless steel rial had rekrated at the end of the cafeulation. This melted at a temperature of 1700 K (260(PF),it flowed was the result of the slow heatup, which alkmed the down the control nxis until it cooled enough to cladding to oxidi/c completely, and is consistent with refreeze. Stost of the Ag in-Cd control material that esperiments. Experiments have shown that with slow was releasal when the stainless steel cladding melted heatup rates [less than 0.8 K/s (1.4'l%)] the cladding flowed down irto the kmer plenum because it has a w ill tend Io oxidite in place, w it h no rekrat ion until Ihe much kmer melting temperature than the steel, around iirconium dioside melting temperature is reached.33 1000 K (134(FF). That is,it will flow farther than the When the melting temperature of firealoy 12123 K steel before enough energy is remosed m that it will (3365'F)] is reached, t!.e tirealoy on the imide of the refree/e. As mentioned abose, this molten material cladding begins to melt. Iloweser, the oxide shell on the then interacted with the liquid in the kmer plenum. As outside of the auld;ng is thick enough that the molten the heatup continued, the stainless sicci at kmer cloa-zirealoy is unable to break through. As more oxygen tiom began to melt. This included material that had diffuses through the oxide layer, the molten /ircaloy presiomly llowed dow n and refroren. At the end of the oxidites, changing back to a solid, calculation, the top 80% of the control rods had The estemiw midation of the cladding generated a melted and rekwated d(mnward; none of the stainless significant amount of hydrogen. Figure 48 shows the steel had rekrated to the lower plenum.
integra! hydrogen generation. At the end of thecalcula- Stost of the control rods in licilefonte contain tion, the top 70% of the fuel nxi cladding was com- II Ct only the asial power shaping rods contain pletely midized. This contributed to t he total hy drogen Ag in Cd.I4 lloweser,Ihe SCDAl'/REl.AP3 code generationof 867 Lg(1910 lbm),whichcorrespondsto cannot currenfly model ll,C in Ihe cont rol rod com-midation of about 79% of the total tirealoy mailable ponent. Changing the control material from Ag in-in the core. Cd to II,C would base a small effect on the Relocation of control rod material dkl occur, calculation, lloron carbide melts at about 2620 K Figure 49 presents the control rod guide tube tempera- (4256'F), well abme it.c melting temperature of 50
1000 i i i i
- 2000 800 -
l . -1500 m
I i
9.x 600 -
f l
V O l
- - 1000
=
j 400 -
y
{
500 200 . -
0 O O 2500 5000 7500 10000 12500 Time (s)
Figure 48. Total hydrogen productio's for the S2 D transient.
l l
3000 , , , ,
1 A Node 1 O Node 3 2500 -
O Node 8 -
X Node 10 4000 m m M
v b
v
- 2000, - - *
$ -3000 g 0
L o
L
[ 1500 g
E 2000 E H F-1000 - -
-1000 500 " ' "
0 2500 5000 7500 10000 12500 Time (s)
Figure 49. Control rod guide tube temperatures at nodes I,3,8, and to for the S;D transient.
51
stainless steel. When the stainless steel relocates revaporization of deposited Csl and CsOli begins (the timing should be unaffected by the different to become significant at surface temperatures control materia!), the unsupported Il4C should fall. above about 800 K (980*F) and increases exponen-That entering the lower plenum would boil liquid. tially w ith increasing temperature. [The vapor pres-The amount boiled, compared to Ag-in-Cd control sure of Csl is about 1.2 x 10-4 Pa (1.7 x 10 8 psia) material, would depend on the relative heat capaci- at a temperature of 600 K (620'F), 1.1 Pa ties. Some oxidation of the ll 4C may also occur, (1.6 x 10-4 psia) at 800 K (980 F), and 173 Pa generating additional hof rogen.15 (0.025 psia)at 1000 K (1340 F).16The vapor pres-The fuel rod cladam7 ballooned near $900 s and sure of CsOf f is about 35 pa (0.0051 psia)at 800 K ruptured near 6020 s. The top half of the cladding (980 F) and 1660 Pa (0.24 psia) at 1000 K ballooned to reduce the core flow area by 89re . The (1340 F).37] The temperatures near the end of the lower half of the core ballooned to a much lesser calculation are in the range that most of the cesium extent, with the bottom node not ballooning at all and iodine that had been deposited on the piping and the fifth axial node ballooning to a flow area near the reactor vessel would have reentered the reduction of about 19fo. When the cladding rup- coolant flow, to be deposited on the cooler piping tured, a path was opened for fission products downstream or carried out the break. The steam released from the fuel to reach the coolant. velocity in the cold leg p, ping at the end of the cal-Figure 50 shows the cumulative release of soluble culation was about 5.2 m/s(17 ft/s).Similarly, the fission products (cesium and iodine) from the fuel fission products deposited in the upper plenum rods to the coolant. At the end of the calculation, would have been revaporized.
about 14.2% of Ihe cesium, iodine, xenon, and The loops seats did not clehr ofliquid. Ilot vapor krypton that was initially in the fuel had been leaving the core flowed through the reactor vessel released. Initial fission product imentories are pro- valves to the downcomer, then out the cold leg to sided in Appendix 11. the break. The hot legs remained near the satura-After the fuel rod cladding ruptured, a brief 01- tion temperature. Con equently, the potential fail-culation (about 400 s) of Ihe fission product trans- ure of the steam generator tubes is not a concern for port was performed. The species transported were (Sis tran ient.
Csl and CsOH. This calculation showed that the Figure 52 presents the break flow during the fission products were transported from the core to transient. The flow rate decreased as the upstream the upper plenum, througl. the vent sahes to the fluid changed from subcooled to saturated liquid, downcomer inlet annulus (the vent vai es were then to a two-phase mixture. The large increases in open through most of the transient), and out the the flow between 700 and 1200 s were caused by cold leg to the break. Of the 5.6 kg (12.4 lbm) of liquid draining from the steam generators; the flow cesium and iodine released to the coolant by increased as the upstream quality decreased.
6400 s, about 54% was retained on the upper Figure 53 presents the void fraction in the solume plenum structures,20% was released to the con- upstream of the break. After about 1500 s, the flow tainment through the break (99.5% as aerosols), out the break was nearly pure vapor, so the flow 18% was retained on the downeomer and cold leg rate decreased as the pressure decreased.
piping,8ro was retained on other structures in the The liquid volume in the Iwo CITs is shown in RCS, and 0.3% remained in the coolant. The upper Figure 54. The RCS pressure dropped below the plenum structure temperatures at this time ranged CFT pressure of 4.24 N1Pa (615 psia) near 6630 s.
from 640 to 1000 K (692 to 1340*F), and the cold initiating the injection of the CFT liquid. As this leg piping temperatures were between 631 and liquid boiled, the RCS pressure increased above the 573 K (676 and $72*F). The steam velocities were CIT pressure, stopping the injection. The pressure about o.12 to 0.60 m/s (0.4 to 2.0 ft/s)in the upper then decreased until injection began again. This plenum and 2.5 m/s (8.2 ft/s)in the cold leg. The intermittent injection continued through the rest of calculation was not carried any further because the the transient. At the end of the calculation, about bulk of the fission product release will occur when 21% of the initialliquid volume had been injected fuel liquefaction begins, into the RCS.
The piping temperatures in the broken cold leg Figure 55 shows the noncondemible quality in are show n in Figure S t. The two solumes closest to four of the ccre volumes. The noncondensible the reactor sessel were near the same temperature, quality is defined as the mass fraction of noncon-with the pipe temperature then decreasing with densibles in the vapor phase. The quality increased ,
increasing distance from the sessel. The rate of steadily until about 8000 s, after w hich it c
$2
.- ------- - -.- - - . . . -- -_ - --_, ._ .-. . - - ~ _ - _ . - _ - _ .
30 i i i i
. 60
. 50 20 - -
^ <
^ -
-40 E c:n ,o b O M
m - -30 m O m O
2 2 ,
10 ,- -20 l
. 10 i
Burst release :
0 0 0 2500 5000 7500 10000 12500 Time (s)
.f i I~igure 50. Total release of cesium and iodine for the S,D transient. 'i l 110 0 , i i i i
2400 A Vessel nozzle O Near vessel 1000-- O At break --2200 X Near pump
^ ^
M k'-
v goo - . . -2000 v e e L L.
3 - "3 800 - -
1800 *-
D D L L e e Q- -
c.
1600 E 700 e
E
> e
-1400 600 - -
IW_-- O , , - n X -X X X- -X- N l 1200 1
l 500 O 2500 5000 7500 10000 12500 !
Time (s) '
I I igure 51. Volume-aserage piping temperatures in the broken cold leg for the S2 D transient.
53
200 , , , ,
Liquid drotnlrig from steam generators 400
^
9 150 -
cn E 6 -
-300 v3 e 7
-+- e D *-
' 10 0 . O y -200 o 3 G "
M n 50 -
m O --100 m M
mi c..m k l_2. m .. _ __
0 ' ' ' '
O O 2500 5000 7500 10000 12500 Time (s)
Iigurc $2. Slass flow rate through the break for the S,I) tran ient.
1 r ,
j 7 y
.,,, y ,, . , , , , ,
0.8 -
C O
4 0.6 -
o O
M
.3 0.4 -
o 0.2 -
0 O 2500 5000 7500 10000 12500 Time (s)
I~igure 51. Void fraction in the solume upstream of the break for the S;I) transient.
L
$4
m '
80 ' ' ' '
. 2800 75 -
. 2600 p p C a v g70 -
[
. 2400 65 -
- -2200 60 O 2500 5000 7500 10000 12500 Time (s)
Figure 54. Liquid solume in the CFTs for the S 2D transient.
03 , , , ,
A Node 1 O Node 4 x o,4 _
O Node 8 ,
G X Node 10 ;
O tr e 0.3 - Control .
- material
.O relocotton S
j 0.2 - -
g 1 s E
$ 0.1 - -
0.0 3 0
3 2500
'= = =
5000 x$
7500 Il --M 10000 12500 Time (s)
Figure 55. Noncondensible quality in core nodes I,4,8, and 10 for the S D transient.
2
Ucereased. The decreasc occurred because less clad- sient were analyzed. Two cases based on the
^\ ding was being osidized. The sharp drop near Abnormal Transient Operating Guidelines ( ATOG) 10,500 s rene;ted the increase in the steam flow for llellefontel8 were investigated. Each case was through the core caused by relocated control rod analyzed until the core was covered with liquid with material boiling liquid in the lower plenum. the I.Pl pumps injecting or until core damage had Figure 56 presents the pressure in the two steam occurred. The first case (Case S-4) followed the generators. Tne large pressure deercase from 300 to operator response set forth in the ATOG, and the 600 s was caused by condensation in the steam gen- second did the same while further assuming ihat j, crators resulting from the AFW injection. When the CFTs were not available (Case S-5). As a pre-the injection stopped, the pressure increased. The lude to presentation of the analyses, the ATOG p; essure esentually increased to w here the .\1ADVs response will be described.
were used to controlIhe pressure. The RCS pressure
, decreased below the steam generator pressure near
$200 s, so the nearly constant steam generator pres. ATOG for the S2D Sequence. The Abnormal sure after Ihat time was the result of no further heat Transient Operating Guidelines are a symptom-transfer from the RCS. based set of procedures developed by Babcock and The AFW flow was initiated near 115 s. The lig- Wilcox and designed to lead the operators from the uid level in the steam generators reached the set- first indication of an abnormal condition through
, , point of 80% on the operating range measurement recovery of the plant. This discussion of the ATOG at about 700 s and was maintained at that level for will be restricted to the S 2D sequence. A more the rest of the transient. detailed description of the ATOG for this sequence is contained in Appendix 11.
~
S!D*
Operator Action Following the initiation of the cold leg break, the InVOSt.igat.lOnS ATOG is entered when the reactor scrams. The operators are instructed to take certain immediate actions which essentially verify and manually Since the S D sequence led to severe core dam- duplicate the plant automatic response. The age, operator actions that may mitigate the tran- response in the first few minutes of the transient 9 e i i
i w AFW on f C m m_ _m ,,
1200 8 * " u -- - O-n n D D e D a' .--
e d7 -
~
-1000 O e 4 u e 3 0 '
- 2 E. 6 ~
E L .
- o. -
-800 L A Steam generator A g 5 -
O Steam generator B 4- i . . . -600 0 2500 5000 7500 10000 12500 Time (s)
Figure 56 pressure in steam generators A and 11 for the S;D transient.
56
includes the reactor scram, the generation of an the LPI pumps injecting or core damage had ESFAS (with the associated closing of val es and occurred.
initiation of the ECC systems), and the trip of the RCPs on loss of subcooling margin. These actions Operator Action Case 1. The first operator are assumed to be completed 3 min after scram, action imestigation (Case S-4) followed the ATOG.
The operators then proceed following a proce- Table 10 presents the sequence of esents for the dure for lack of adequate subcooling margin. The transient.
only actions that are taken that are not automatic The first 500 s of the transient were nearly identi-are the attempts to isolate the break; the one that cal to the base case S D calculation. At 500 s, the affects this analysis is the closing of the block sabe pressurizer spray line block vahe was closed. At in the pressurizer spray line. After checking the sta- 806 s, the controlled cooldown of the RCS began.
tus of the plant, the procedures indicate to the The coupling between the primary and secondary operators that there is a small break LOCA and systems was sery good, so that the RCS responded that they should begin a controlled cooldow n of the quickly to any changes in the steam generator pres-RCS using the steam generators. It was assumed sure. Ily 1000 s, the RCS pressure had been reduced that this cooldown would begin about 13 min after to approximately 5.5 SIPa (800 psia). The RCS the reactor scram. pressure and temperature remained fairly constant A cooldown of 28 K/h (50'F/h)is initiated and between 1000 and 4000 s. The liquid lesel in the controlled by opening the NIADVs on both steam core had dropped low enough by 4000 s that the generators. It was assened that the operators core began to heat up. The ICC procedure was would reduce the hot leg temperature 28 K (50 F) entered at 4020 s. As the steam generator pressure as quickly as possible, and then hold the tempera- was reduced, the RCS pressure and temperature ture constant for the remainder of the hour. The followed. The pressure reduction increased the target hot leg temperatute was reduced another amount of liquid boiling in the core, causing the 28 K (50'F) for each addit.:onal hour the transient two-phase les el to swell. The higher level cooled t he continued. The operators are instructed that, fuel rods back dow n to the saturation temperature.
should the core exit thermocouples show super- After one minute with the core outlet temperature heat, the inadequate Core Cooling (ICC) proce- still at saturation, the controlled cooldown proce-dure should be entered. It was assumed that dure was reentered. It was assumed that the 28 K/h superheat was indicated w hen the core outlet sapor (50 F/h) cooldown rate was still based on the time temperature was 2.8 K (5*F) above the saturation that the procedure was first entered (806 s).
temperature. The ICC procedure was entered one When the controlled cooldown procedure was minute later. reentered, the NI ADVs were closed. The steam gen.
The ICC procedure calls for a more rapid depres- erator and RCS pressures began to increase imme-suritation of the RCS. The steam generator pres- diately. The core began to heat up again near sure is reduced until there is about a 56 K (100'F) 5080 s, and the ICC procedure was entered again difference between primary and secondary 60 s later. As with the first temperature excursion, coolant system saturation temperatures. The core the reduction of the steam generator pressure in the outlet thermocouplu are monitored to provide an ICC procedure quickly brought the core tempera-indication of the fuel rod temperatures and the tures back to saturation and the controlled proximity to sarious stages of core damage, cooldow n procedure was entered again. This alter.
Depending on the core exit temperature, the high nating between procedures occurred eight times point sents on the loops, reactor vessel, and pres- during the transient.
suriier may be opened. The RCPs may be turned injection of liquid from the CFTs began near on, and the hot leg dump-to-sump valves and the 4540 s. This injection helped to return the core tem-PORY may be opened in an attempt to reduce the peratures to saturation during the ICC procedure.
RCS pressure to allow LPI to reflood the core. If When the RCS pressure was reduced, the CFTs these actions are successful in returning the core would inject liquid to the reactor sessel and core.
exit temperature to saturation, the controlled This liquid partially compensated for the mass lost cooldown procedure is reentered. through the break. The injected liquid then had to in either case, the cooldown is continued until be boiled before the core could heat up again.
the core exit fluid becomes subcooled. For these The pressure continued to decrease over time as analyses, the calculations were defined to be over the ICC procedure was entered repeatedly. The when either the core was covered with liquid with pressure decreased below 1.48 NIPa (215 psia) at 57
l l
Table *0 Sequence of events for the two S 2D operator action cases Time
- (s) l I Event Case I Case 2 Transient initiated 0 0 Reactor scram 24 24 RCPs tripped 53 53 ESFAS signal generated 53 53 l Ilot legs reached saturation 130 130 l Lack of adequate subcooling procedure entered 206 206 Pressurizer spray block valve closed 500 500 Controlled cooldow n procedure entered 806 806 ICC procedure entered 4020 4020 Controlled cooldown procedure entered 4140 4140 CFT injection began 4540 -
ICC procedure entered 5140 5060 Controlled cooldow n procedure entered $260 5180 ICC procedure entered 5520 5390 Controlled cooldown procedure entered 5660 5520 ICC procedure entered 5760 5640 Controlled cooldow n procedure entered 5880 5770 iCC procedure entered 5960 5870 Controlled cooldown procedure entered 6120 6030 ICC procedure entered 6490 6100 Controlled cooldow n procedure entered 6600 7480 ICC procedure entered 6810 --
Controlled cooldown procedure entered 6960 -
ICC procedure entered 7170 -
Controlled cooldown procedure entered 7340 -
LPI injection began 7886 7420 Core covered 7950 7540 Calculation terminated 8012 7673 7886 s, initiating the flow from the LPI pumps. of the liquid could not be compensated by the The core was quickly covered, and a stable long- break flow. The pressure decreased rapidly near term decay heat removal configuration was 600 s when the break flow increased as the attained by 8000 s. No core damage occurred dur- upstream quality decreased. The pressure decrease l
i ing the transient; the peak fuel rod cladding tem- near 800 s was the result of the depressurization of perature during the heatups was 611 K (640'F), the steam generators. Since both the RCS and the w hich is below the steady-state operating tempera- steam generators were at saturation, the two pres-l ture of the ciudding, sures were nearly equal through the rest of the tran-Figure 57 shows the RCS pressure during the sient. The RCS pressure oscillated from 1000 to t ransient. l'he pressure decreased rapidly at the 4000 s, reflecting the cycling of the MADVs. The l
beginning of the transient, leveled off when the pressure decreased when the ICC procedure was RCS reached saturation, then increased as boiling entered, then increased when the controlled 58
20 , , e i
- -2500 15 -
n g -
-2000 D Q- p 2 ch.
v v e 10 -, _-1500 e u u 3 3 m V m m ICC procedure entered a o -1000
- EN tN 5 -
LPI on N, vu i e -500 0 O O 2000 4000 6000 8000 10000 Time (s)
Figure 57. Pressurizer pressure for S 2D operator action case 1.
cooldow n was resumed. The pressure decrease near Figure 61 shows the break flow. As in the base 4400 s occurred while the plant was in the con- case S2 D transient, the increases in break flow near trolled cooldown; one hour had elapsed since the 600 and 800 s were caused by a decrease in the qual-procedure was first entered, so the hot leg tempera- ity of the fluid upstream of the break. The flow ture could be reduced another 28 K (50'F). remained fairly constant after 1000 s. The large The steam generator pressures are shown in Fig- increase in flow near the end of the transient ure 58. Condensation driven by the injection of reflected the LPI llow, some of which flowed out AFW caused the pressure to decrease at 200 s. The the cold leg instead of down the downcomer.
pressure oscillated between 1000 and 4000 s as the The core collapsed liquid level is presented in Fig-MADVs were opened and closed. The valves were ure 62. The level reflected the cycling of the pres-cycled to n aintain the desired cooldown of the sure, with the level increasing as the pressure RCS. As with the RCS pressure, the pressure decreased. Injection driven by the LPI pumps decreased rapidly when the ICC procedure was quickly covered the core with liquid at the end of entered, and increased when the controlled the 'ransient.
cooldow n was resumed.
Figure 59 presents the fuel cladding surface tem- Operator Action Case 2. The first operator peratures in the top three axial levels. The brief action case demonstrated the effectiveness of the heatups did iwr extend any lower in the core. The ATOG in mitigating the S D 2 sequence. The second figure shows that eight temperature excursions case assumed that the CFTs were not available to occurred during the transient, none of uhich determine their effect on the transient (Case S-5).
resulted in temperatures exceeding the initial The response of the plant was identical to the steady-state operating value. first operator action case for the first 4500 s of the The upper plenum and hot leg vapor tempera- transient, when the CFT injection began. After tures are presented in Figure 60. Iloth temperatures 4500 s, the plant response was similar to the first were close to the cladding temperatures throughout operator action case. When the core oidlet became the transient. The increase near 7000 s would not superheated, the steam generaters were depressur-be as large as was calculated, since the vapor tem- ized. The subsequent depressurization of the RCS perature could not exceed the maximum cladding caused t he two-phase level in the core to swell, cool-temperature, ing Ihe portion of the fuel rods that was heating up.
59 l
1
20 , , , ,
l A Steam generator A
- O steam generator B 2500 15 0
- -2000 -e n.
2 v v S
e go -. _ 1500 e
' L 3 3 h $
- O WADV cycling , 1000
- g Q- n.
5 -
- -500
' ' ' ' O 0
O 2000 4000 6000 8000 10000 Time (s)
Figure 58. Pressure in steam generators A and B for S,D operator action case 1.
900 , , , ,
A Node 8 O Node 9 O Node 10 1000 800'- .
m ^
v M l'-
v 8 700-- -
800 8 3 3 O O l
E600l .
600 $
E j E e
.- g _ ,-_;- g -
l
)) o 500 - -
- 400 400 O 2000 4000 6000 8000 10000 Time (s)
Figure 59. Fuel cladding surface temperatures in nodes 8,9, and 10 for S,D operator action case 1.
60
900 i i 8 '
A Core outlet O Hot leg
_-1000 800 ~ -
m 2
v
--800 e e 700 --
L 3 3 0 0 L L g . e e 600 -
- o. 600 n.
E
- = : - -_2-_- -_ Q :- f 500 - 3 -
400 g i I I 400 0 2000 4000 6000 8000 10000 Time (s)
Figure 60. Upper plenum and hot leg sapor temperatures for S D operator action case 1.
600 i i e i
-1200
? 7 g N
-1000 cn E
.o 6 400 - -
O
-800 ,
-+-
O 6 O 6
-600
- o
)
o 200,- ~
400 C
M M sn O in I
o
( -200 2 0
- h.
- ia..
. m m. . 2 mm
': LPI on O
O 2000 4000 6000 8000 10000 Time (s) l'igure 61. Mass flow rate through the break for S 2D operator action case 1.
61
4.5 i i i i
-14 4 ~
^
E 12 -O v 3.5 -} -
v m \ -
O m 3-- h _ 10 0 f -
, hf ht ,
._ l- *-
k, S- 2.5 -
LPI on -
-8 3 a 2 -
1 i
' f 6
15 O 2000 4000 6000 8000 10000 Time (s)
Figure 62. Collapsed liquid leselin the core for %D operator action case 1.
Iloweser, in this case there was no water being case. This led to a faster depressurization of the RCS injected to compensate for the mass lost through and an earlier recoscry of the plant.
the break. As the transient progressed, the depres. The effect of the CFTs on the transient was to surization of the RCS was less effectise in cooling keep the cladding temperatures lower w hile slow ing the core; it required a larger pressure decrease to the depressuritation of the RCS. In neither case return the cladding to the saturation temperature. was the core damaged. The later reflood of the core This behasior is shown in Figure 63, which (about 400 s) with the CFTs injecting put no presents the fuel rod cladding temperatures in the extraordinary demands on the operator; the only top four axial nodes. The duration of the tempera. difference in operator actions was that the ICC pro-ture excursions increased as the transient contin. cedure was entered two more times w hen the CFli ued. The temperature response also shows that werc as ailable.
after the temperature returned Io saturation and the controlled cooldown was reinstated, the heatup began again almost immediately. When the pres. Summary of Small Break LOCA sure began to increase, the level dropped, uncoser. Results ing the top of the fuel again. The last two times the temperature returned to saturation, near 6200 and 6400 s, the cladding did not stay at the saturation F ve small break LOCAs were analyicd. Two of temperature long enough (60 s) for the controlled these transients resulted in core damage; three did cooldown procedure to be reentered. The LPI not. Table 1I summarizes the small break LOCAs.
The two transients without ilPI or operator pumps began injecting liquid at 7420 s, reflooding the core. The peak fuel rod cladding temperature of actions (Cases S-1 and S-3) resulted in core dam-ace within 2 h. The S,D sequence (Case S-3) led to 655 K (719'F) was attained just before the LPI sNere core damage a'nd fuel melting after about began. As in the first operator action case, this tem-3-1/2 h. The transient in w hich IIPI flow was avail-perature was low enough that no core damage occurred during the transient- able(Case S-2) led to a stable feed-and-bleed of the RCS with no core damage. IIPI flow exceeded the Table 10 presents the sequence of esents for the two break flow after about 3 h. Without operator operator action analyses. The second case spent more actions, the asailability of one 11PI pump was nee-time in an inadequate cooling mode than did the first essary and sufficient to present core damage.
62
The two transients in which operator actions resulting loading of the containment during the were considered (Cases S-4 and S-5) both led to a course of the transient.
successful recovery of the plant with no HPI The amount of unoxidized material in the mol-shortly after 2 h. These actions were based on the ten core material released from the vessel will affect Bellefonte ATOG. The procedure followed to miti- the core / concrete ir.teraction. If there is little unox-gate the transient was straightforward. Since it was idized material, fewer fission products will be basically an extension of a normal cooldown, the released from the pool to the containment as aero-operators would not be expected to encounter large sols. The amount of aerosols in the containment difficulties in executing the procedure. atmosphere at the time of containment failure will The results of the core damage analysis for the affect the total release of fission products to the S2 D sequence provided information related to sev- atmosphere, eral severe accident issues. The cladding was exten- The high structure temperatures along the flow sively oxidized (79%) prior to relocation of any of path indicate that few fission products will be the fuel rod material. Most of the hydrogen pro- retained in the RCS. Fission products that were duced will be released to the containment prior to deposited on the surfaces would be revaporized and breach of the reactor vessel. The operable contain- released through the break to the containment, ment hydrogen recombiners will be able to remove Containment bypass caused by failure of the steam some of the hydrogen during the course of the tran- generator tubes is not a concern. The loop seals sient, although their capacity of 0.05 m /s 3 remained covered with liquid throughout the transient.
(100 cfm) will cause little change in the amount of Consequently, there was no loop flow through the hydrogen in the containment. The timing of the steam generators, so the hot lluid in the reactor vessel release, while the reactor vessel is still intact, did not enter the hot legs or steam generator tubes. The together with the location of the release will affect flow from the core outlet was into the upper plenum, the concentrations of hydrogen m the various parts through the reactor vessel vent valves to the down-of the containment. These concentrations will comer, and out the cold leg to the break.
determine the potential for combustion and the 700 , , , 800 600 il -
n -
n g -600 v p v
e ==, =4=3 ,
5 500 Mk g g -
3 -400 g a
E a
- E 400 -
A Node 7 ~
. O Node 8 0 Node 9 -200 X Node 10 300 ' ' '
O 2000 4000 6000 8000 Time (s)
Iigure 63. Fuel rod cladding surface temperatures in nodes 7,8,9, and 10 for S D operator action case 2.
63
k i
l .
l Table 11. Summary of the small break LOCA analyses l
! Calculation Transient Case Code Core Model Ended Result ,
LOFW, no ECC S-1 RELAPS Single channel 3450 s Core damage began at 2750 s LOFW, one HPI S-2 RELAPS Single channel 15,000 s No core damage, RCS liquid inventory increasing by pump available 10,800 s l
S2D sequence S-3 SCDAP/RELAP5 Single channel 11,900 s Core damage began at 5650 s l
! (no llPI) 1
] g S2D operator S-4 SCDAP/RELAPS Single channel 8012 s No core damage, liquid level above core by 7950 s action case i S2D operator S-5 SCDAP/RELAPS Single channel 7673 s No core damage, liquid level above core by 7540 s action case 2 (no CFTs) 4 4
t I
CALCULATIONAL UNCERTAINTIES This section presents a general, qualitative discus- from the steam generator side of the h30p seals to the sion of the uncertainties associated with the analyses pump side, where it can be dmwn into the cold legs presented in the presious sections. These uncertainties more easily. The cycling of the relief valves causes the will be those beyond the inherent limitations of the water in the loop seals to mose back and forth between code; that is, phenomena that the codes were not the pump and steam generator sides of the loop seal designed to calculate uill not be discussed. This can allow some of the water to now into the cold The RELAP5 and SCDAP computer codes are leg, from u here it can flow into the reactor vessel. (The fairly mature. The SCDAP/RELAP5 code is still cold legs are sloped downward to the reactor vessel.)
being developed and has undergone sirtually no The cycling also changes the temperature of the liquid assessment at this time. flowever, as the analyses as the saturation temperature of the RCS changes.
have shown, this code has capabilities that are When the RCS is pressurizing, the temperature needed to perform sesere accident analyses, such as increases. When the salves open and the pressure the feedback between the loop behavior and the decreases, some of the liquid will Hash to steam. If core, and between the core deformation (balloon- enough water is removed from the loop seals during ing) and the core hydraulies. these cycles, a path will be opened for steam to How l The clearing of the kiop seals was a significant event through the loop. Once the clearing begins, condensa-in the Ull.II' sequence analyses. The nature of the tion in the cold legs may help it to continue. Condensa-transient changed after the liquid was cleared, slowing tion of steam in the cold kgs would cause the pressure the heatup of the core and leading to the probable fail- to decrease, which in turn would tend to draw more ure of the RCS pressure boundary long before the ves- liquid from the loop seah into the cold leg. The lack of set would be breached by molten core material. relief valve cycling in the pump seal leakage analysis Although the loop seah cleared of liquid in three ofIhe was the probable reason that the k)op seals did not four system calculations, a common esent did not initi- clear during that transient. Other changes in the RCS ate the clearing. In the three-channel RELAP5 and pressure can also help t he k>op scals to clear, such as t he single-channel SCDAP/RELAP5 calculations, the pressure w ave generated by molten cont roi rm! material k>op seah cleared uhile the RCS pressure was decreas- falling into the liquid in the lower plenum.
ing after the SRV cycling but before the PORY had The modeling of the once-through steam generators begun to cycle again. In the three-channel SCDAP/ is another area of uncertainty. In most of these tran-RELAP5 calculation, the k)op seals cleared after the sients, however, relatively simple How conditions PORVs had begun to cycle again. In each case, the existed on both sides of the steam gene ator tubes. In clearing began in the A k)op; after one or both k)op the LOFW transients, the secondary side boiled until seals in the A k)op cleared, both of the loop seals in the there was no water left. The heat t,ansfer was then to 11 k)op cleared. The reason for this is that one A-k>op single-phase steam. On the primary side of the tubes, cold kg is connected to the top of the pressurizer the heat tmnsfer was also to a single-phase Guid through the pressuriict spray bypass line. The top of through most of the transient. In the analyses in which the pressuri/er is the low pressure point of the system no core damage occurred, the steam generator tubes because the relief salves are open either constantly or had liquid in them throughout the transients. In the intermittently. Once one loop seal clears, the pressure core damage transients, the liquid drained, leaving difference between the hot and cold legs changes, caus- steam inside the tubes. This was also the case in the ing the 11 loop seals to clear. Some factors affecting the small break transients with a LOFW (Cases S-1 and kx)p seal clearing are the density difference between the S-2). When feedwater was available, the secondary side Huid in core and the steam generator tubes, the cycling had water that was being boiled until the RCS pressure of the relief valves, and condensation in the cold legs, dropped below the steam generator pressure, effee.
The density difference between the Huid in the core and rively terminating the heat transfer across the steam that in the steam generator tubes determines the rela- generator tubes. During this time, the Guid in the tubes tive heights of the columns of water on either side of changed from single-phase liquid to single-phase the k>op seal. The water level is higher on the pump vapor. llecause the steam generators were decoupled side than on the steam generator side because the Guid from the RCS relatively early in these transients, their in the steam generator tubes i3 denser than that in the exact behasior was not critical to the transient progres-core (saturated steam instead of superheated steam). sion. The transients that would be most affected by As the core continues to heat up, more liquid moses differences in the steam generator behavior were the 65
S2 D operator action cases (Cases S-4 and S-5). The part of the core to be damaged while the bottom AFW injection affected the RCS behavior because the remains cosered with liquid.
two systems were coupled throughout the transient. In The modeling of the cmssuows in the core is not an effort to better model the AFW it.jection, a split straightforward. No data exist on what the crossuows steam generator model was used. This model sepamted are during various stages of the transient. Since there is the tubes into two regions. The larger region contained no way to check the flows, one must rely on their gen-the tubes that would not be directly affected by the eral characteristics. The Dow patterns that were estab-AFW. The smaller region modeled those tubes that the lished with the three-channel models were physically AFW would wet as it Howed down the tube bundle. reasonab!c, based on the densities of the Guids in the For these tuo transients, the heat transfer would be three channels. The resistance to now between the from boiling liquid on the shell side to condensing channels is not known. The SCDAP/RELAP5 code steam on the tube side. Th:re was no liquid level in ihe does not change the flow area or resistance for the tubes. If the coupling were not as strong as was cateu- crossuow when the cladding balloons; only the axial lated, Ihe RCS would base depressurized more slowly, now area is changed. This limitation of Ihe code, how-resuhing in a longer heatup of the fuel rod cladding. ever, had a very small effect on the calculated results.
However, the heatup that was calculated was so short The cladding ballooning in the three-channel that extending the heatup would still not lead to core SCDAP/RELAPS TMLB' analysis (Case T-le) damage before the LPI began and recovered the core. disrupted the in-vessel natural circulation flow that The two-phase level in the core also affected the had been established. Less extensive ballooning S2 D operator action cases. The level swell associ. may not have disrupted the now. A limitation of the ated with the depressurizations of the RCS cooled calculation was that a coplanar flow blockage was the fuel rods, terminating the heatups. If the two- assumed to occur in each region of the core in phase level were lower, the heatups w ould continue. which the cladding ballooned. The actual balloon-This, in turn, would lead to further depressuriza- ing would vary from rod to rod, depending on the tion of the RCS. In the first operator action case, local conditions. However, a similar sit uation exists this would allow more CFT liquid to be injected, in any calculation in which more than one compo-which would raise the liquid lesel in the core and nent is modeled by a single representative or aver-cool the fuel rods. In both operator action cases, age component. The ciuestion then is whether the the RCS would depressurize to allow LPI earlier calculated balloon is representative of the average than was calculated. Again, because of the large balloon in that region of the core. The sausage-type difference between the calculated peak cladding ballooning that was calculated is probably accu-temperatures and the temperature at which core rate, because the relatisely uniform temperatures damage would begin, core damage would not be (axially) of the cladding would tend to promote the expected if the two-phase level were lower than was formatian of longer balloons. The radial extent of calculated. In the other analyses in w hich the core the ballooning determined the flow area reduction.
uncoscred, a lower two-phase level would cause a Further study is needed to determine how sensitive slightly earlier heatup of each of the core nodes, the flow disruption is to the size of the balloon; it with no noticeable effect on the transient results. may be that the now can be disrupted by fairly The balance between the break How and the HPI small regions of ballooning, so that the calculation How in the small break LOCA with HPI (Case S-2) is not sensitive to the extent of the ballooning.
could affect the results of that transient. The HPl How The timing of the operator actions for the S2 D exceeded the break now before the core began to analyses is also uncertain. The response of the uncoser. If the break Dow were someu hat larger, or the plant, however, indicated that differences in the H PI How smaller, the core may hase begun to uncoser timing of the actions would not have a significant and heatup. If it had, the depressurization rate would effect on the transient progression. Higher core hase increased because less liquid was being boiled. temperatures would result from longer delays in The lower RCS pressure would allo,' more How from beginning the ICC procedures, but core damage the HPI and would result in earlier CFT injection. It would still not be expected because of the low peak may also be possible to establish a balance between the temperatures that were calculated. If the controlled flows with a larl in the core such that the top of the cooldown procedure entry were delayed w hen exit-core is heating up. The break How would just offset the ing the ICC procedure, the depressurization of the boiling of the injected liquid, keeping a constant toel RCS would continue, leading to an earlier initia-in the core. Such a heatup of the core may allow the top tion of LPI How and recosery from the transient.
66
The now through the relief vahrs is another area of those calculated. This would not be expected to have a uncertainty. The relief mhes were modeled to provide noticeable effect on the transient behavior, since the the design How at the design pressure with pure steam. How brough the relief valves was generally single-Two-phase and single-phase liquid How rates through phase vapor.
the valve are not known and may be different than 1
f 67
CONCLUSIONS Seseral LOFW and small break LOCA transients zirconium dioxide was reached. The smaller were analyzed using the RELAPS, SCDAP, and amount of unoxidized material in the molten SCDAP/RELAP5 computer codes. Conclusions core will also affect the core / concrete interac-drawn from the analyses are presented below. tion after the reactor vessel fails. Less hydro-gen will be generated in the containment, w ith
- 1. The availability of one fiPI pump was neces- fewer fission products tricased from the mol-sary and sufficient to present core damage.
ten pool into the containment atmosphere.
With no HPI or operator actions, the small 4. Nateral circulation in the core reduced the break LOCA and LOFW transients all corc heatup rate compared to once-resulted in core damage. With one H PI pump tbrough calculations.
injecting, the plant was able to attain a stable decay heat remosal configuration with the cladding at the saturation temperature Three-channel models of the core and unper plenum were used in the analysis of (Cases F-3, F-4, and S-2). This configuration the TNILB' sequence (Cases F-lb and could be maintained until the borated water storage tank emptied (after 15 to 25 h), w hich F-le). The in-sessel natural circulation transferred heat from the core to the upper shou!J proside enough time to restore power plenum and allowed cooler vapor to return or to repair failed systems and present core damage, to the core from the upper plenum. This resulted in a slower heatup of the core,
- 2. Operator actions could present core dam- with delays in the timing of the onset of age for the S,D sequence. cladding oxidation, ballooning, and ses ere core damage. The higher structure temper-dy following the Abnormal Transient atures in the upper plenum will also reduce Operating Guidelines, the plant was the retention of fission products in the "PP"PC""*' I depressurized to allow LPI both with and without the CFTs asailable. Seseral fuel rod cladding temperature excursions 5. Cladding ballooning disrupted the natural occurred, but the peak temperature was circulation in the core.
less than 40 K G2 F) abose the steady-state operating temperature without il$e The three-channel analysis of the TNII 11' CFTs (Case S-5) and was below the sequence (Case F-le) had a well-deseloped steady-state temperature with the CFTs natural circulation flow between the core asailable (Case S-4). and upper plenum. When the fuel rod cladding in the two inner channels bal-
- 3. The transients that led to sesere core dam, looned, the downward flow in the outer age generall) had slow heatup rates that channel was reversed, so that the flow resulted in extensise in place oxidation of throughout the core was from the lower the fuel rod cladding. plenum to Ihe upper plenum. The flow was then primarily one-dimensional, so that The TNILB' and S,D transients were ana. the heatup should be very similar to that lyzed through the' core damage phase calcu'ated using a single-channel model of (Cases F-Id and S-3). Heatup rates below the core. Further investigation is needed to 0.8 K/s (1.4^F/s) were calculated in the inte. determine if less extensise ballooning, gral code (SCDaP/ RELAPS) analy ses. With both in fench and in flow area reduction, heatup rates that low, mosi of the cladding would disrupt the natural circulation.
will oxidize completely before any fuel rod relocation begins. Oxidation of 79% of the 6. The TNILB ' sequence has a high probabil-zirealoy in the core was calculated for the S,D ity of RCS failure prior to the reactor ves-sequence before the melting temperature of sel melt-through.
68
- .-_ __ -. =_. - .- . . . - .-
The loop seals were calculated to clear of with the resultant ecre quench, significantly liquid in three of the four analyses of the altered the core damage scenario. The core ThtLil' sequence. With this liquid heatup rate was slowed to the point that fuel remosed, a strong sapor natural circula- rod relocation would not be expected until tion flow was established in the loops. The after the cladding was completely oxidized, piping was heating up at nearly the same compared to relocation with 4% oxidation in rate as the core. This could lead to failure the SCDAP calculation.
of the piping or steam generator tubes up to 7 h before the core begins to relocate, if 9. Excessive cycling of the SRVs occurred the steam generator tubes fail, there is also w hen the PORV was unavailable.
the possibili.y of containment bypass. The failure of the RCS will allow the system to The LOFW transient with one llPI pump l
depressurire, so that the melt ejection at available but the PORV unavailable the time of vessel failure may occur at low (Case F-3) avoided core damage, flowever, pressure. This will minimize the extent of without the PORV (Case F-4), the feed l
the direct containment heating. and bleed of the plant was through the SRVs. This resulted in approximately the
- 7. No operator actions were available to r re- number of cycles of the SRVs that are vent core damage for the ThtLil' expected in their lifetime. Consequently, sequence, failure of the SRVs may occur.
The only action that the operators could take 10. The clearing of the loop seals in the I
would be to open the PORV, given the sys- TAILII' sequence significantly lengthened tems available during the transient, llowever, the transient.
the PORV was open through most of the transient until core damage occurred. The loop seal clearing resulted in a natural circulation flow through the loops. The
- 8. Integral analyses are needed to investigate additional structural material available as the core damage portion of the transients. a heat sink slowed the heatup of the core, so that relocation of the fuel rod material The interactions of the core with the rest of was not expected until about 10 h after the the RCS are important in determining the transient began. This is much longer than damage progression and timing. The clearing the 1- to 3-h loss of ac power that is of the loop seals in the TNILil' sequence, assumed for the transient.
l i
a 69
REFERENCES i
1
- 1. N1. Silberberg, J. A. N1itchell, R. O. N!c>er, and C. P. Ryder, Reatsesunent ofthe 7&chnical#asesfor E3timating Source Tennt, NUREG-0956, July 1986.
- 2. A. N1. NolactLos ski et al., Interhn Report on Accident Sequence Likelihood Reassessment (Accident i
Sequence Evaluation Program). February 1983, t
- 3. V. II. Ransom et al., RELAPS/3/ODJ Code 3/anual, Ibhunes I and 2, NUREG/CR-4312. EGG-23%,
August 1985.
- 4. C. NI. Allison, E. R. Carlson, and R.11. Smith, "SCDAP: A Computer Code for Analyzing Light Water Reactor Sescre Core Damage," Proceedings of the ANS/ ENS International Aleeting on iight 7
lihier Reactor Seven' Accident Evahtation, Cambridge, Ala, August 23-September 1,1983.
l i
- 5. T. C. Cheng et al., "REl.APS/SCDAP - An Integrated Code for Sesere Accident Anab sis," Pro-l ceedings of the Thirteenth ilhter Reactor Safety Researth inharmation Aleetinx. Gaithersburg, AfD, Octoher 22-23, l985, NUREG/CP-0072, pp. 347-355.
- 6. A. D. Knipe, S. A. Ploger, and D. J. Osetek, PBF Severe Fuel Damace Scoping kst-IPst Results Report, NUREG/CR-4683. EGG-2413, August 1986.
- 7. Z. R. N1artinson, D. A. Petti, and 11. A. Cook, PBFSevere FuelDamage kst I-I Test Re3 nits Report, lois. / and 2, NUREG/CR-4684. EGG-2463, October 1986.
- 8. S. llagen and S. O. Peck, " Temperature Escalation of Zirealoy-Clad Fuel Rods and flundles Under Sesere Fuel Damage Conditions," Paper No. TS-1.9, International Aleeting on Light ilhter Reactor Severe Accident Evaluation, Cambridge, AIA, August 23-September I,1983.
l
- 9. W. E Driskell et al., Developmental Assessment of the Severe Core Damage Analysis IbcAuec: SCDAP/AIODO, EGG-NTAP-6212, N! arch 1983.
- 10. R. Chambers, Analysis of Core Hehavior During a Station Blackout Transient (TAILH') Jar the nellefonte Pnmuri cd linter Reactor, Draft, NUREG/CR-3979, EGG 2342, August 1984.
I1. S. Kawasaki et al., "Recent NIRllT Tests in JAER1," Sixth American-Japanese-Gennan-French fuel Hehavior ilbrkshop Ibkui-Alura, Japan Alay 20.1981.
- 12. C. N1. Allison, D. L. llagtman, and G. A. lierna, "The influence of Zirealoy Oxidation and Nielt-ing llehavior on Core llehasior During a Sesere Accident," Fifth International Aleeting on Thennal Nuclear Reactor Salety, A'arlsruhe, ikdend Republic of Gennany, September 9-13,1984.
l3. S. llagen and S. O. Peck, Out-of-pile Dperiments on the High Ihmperature Behavior of /r-4 Clad f uelRods, Kernforschung/entrum Karlsruhe Report KfK 3567,1983.
l 14. Tennessee Valley Authority, HelleJonte Nuclear Plant, Final Safety Analysis Report, DoeLet 50-438, I
resised through Amendment 22 December 1982.
i
- 15. E. C. Ileahm, R. P. Wichner, and C. F. Weber,"themleall actors Affecting Fission Product Trans.
port in llWR Sescre Accidents," Proceedinus of the Tivelfth ilhter Reactor Safety Researth InJonna-tion Alceting, Gaithersburg, AID,0etober 22-26,1984, NURl G/CP4K)$8 pp.109-l21.
70
- 16. J. A. GiescLe, P. Baybutt, II. Jordan, R. S. Denning, and R. O. Wooten, Analysis of fission Prod-uct Transport Under Tenninated LOCA Conditions,1851l-NUREG-1990, December 1977.
l7. J. C. Cummings et al., Status Report on the Fission-Product Research Program, NUREG/CR-1820 and SAND 80-2662, N1 arch 1982.
l8. llabcock & Wilcox, Bellefimte Nuclear Plant Units 1 & 2 Abnonnal Transient Operating Guidelines
/brt I, Document No. 74-1135402, Revision 5, January 1985.
I l
i
)
1 1
1 i 71
i l
APPENDIX A COMPUTER CODE DESCRIPTIONS A-1
l l
APPENDIX A COMPUTER CODE DESCRIPTIONS The three primary computer codes used in the drical and slab heat structure geometries. Thus, transient analyses were REI AP5/NIOD2,A-I fuel rods, control rods, instrument tubes, and flow SCDAP/NIODI,^-2and SCDAP/ REl.AP5. A-3 A shrouds can be represented. All structures of the brief general description of each of the codes, same type, geometry, and power are grouped together with information on the specifie sersions together; one set of input parameters is used for used, is presented below. each of these groupings or components. Code input identifies the number of rods or tubes in each
'*" P "e"' "d '.heinelatiw positions for the puis RELAP5/ MOD 2 pose of radiation heat transl.er calculations.
Alodels in SCDAP calculate fuel and cladding tem-The RELAP5/NIOD2 computer code was desel-peratures, zirealoy and stainless steel oxidation, oped for best-estimate transient simulation of pres-hydrogen generation, cladding ballooning and rup-surized water reactors and associated systems. It is ture, fuel and cladding liquefaction, flow and a one-dimensional, two-fluid, nonequilibrium of the liquefied materials, release of fis-f i thermal-hydraube code utilizing a six-equation , sion products, fragmentation during reflood, and hydrodynamic model. This model provides conti-subsequent debris behavior.
nuity, momentum, and energy equations for both SCDAP also contains tua models for calculating the licuid and the sapor phases withm a control the thermal-hydraulie conditions in the vessel. One s olume. The energy equation contains source terms of these models is the CHAN component from the which couple the hydrodynamic model to the heat T R AC-BDIA-4 computer code. The second is a structure conduction model by a convectise heat coolant boiloff model which, when used, replaces transfer formulation. The code contains special Ihe TR AC portion of SCDAP with a simplified cal-process models for cr,ticali flow, abrupt area .
m mmp nent to mo1-changes, branching, crossflow junctions, pumps, eu on at imn accumulators, vahes, core neutronies, and control ".nt and kmn mmponent to component.1lu. s umplified model was used for the calculat,on i systems. Crossflow junctions can be used to described i,n this report.
approximate two-dimensional modeling.
The SCDAP code is be,mg used to help plan and The cycles of the code used in the analyses were n lyze the results of severe core damage experi-17,21,22,36.01, and 36.02, w hich are stored under ments, to aid in the determmation of probabilities Code Configuration Control Number F01581, and and tmeertainties in risk assessment analyses, to tape numbers A41094, A33853, B40630, and identify the major contributors to sessel behavior A01511, respectisely, at the Computer Science Lab-
" "E .c r une vering accidents, and to help oratory at the Idaho National Engineering Labora-determme the order and timmg of the events tory (INEL). Cycles 17 and 22 were used to bsened by a plant operator during core heatup.
calculate the TNILB' transient. Cycle 21 was used The sersion of SCDAP/NIODI used to perform for the pump seal leakage calculation. Cycle 36.01 ,
was used for the remaining feedwater transients,
- '"" #"I '"IC"' " ""'" " U' w hich .is stored on tape number B41974 at the Com-and Cycle 36.02 was used for the small break tran-puter Science Laboratory at the INEL sients.
SCDAP/ MOD 1 SCDAP/RELAP5 The SCDAP/NIODI computer code calculates The SCDAP/RELAPS computer code is an the behasior in a light water reactor (13VR) vessel 13VR system transient analysis code that is cur-during extended periods of severe oserheating. rently being developed. It can be used for simula-SCDAP simulates core and vessel plena disruption tion of a wide variety of system transients of by modeling heatup, geometry changes, material interest in LWR safety but is designed especially to relocation, and debris formation, heatup, and calculate the behasior ofIhe reactor coolant system melting. SCDAP allow s detailed modeling of eylin- during sescre accident transients. The core, A-3
i primary sy stem, secondary system, feedwater train, user. The chemical form of the fission products is and system controls can be simulated. The code fixed. All of theiodineis assumed to bein the form models have been designed to permit simulation er of Cst, with the remaining cesium being trans-postulated accidents ranging from small break loss ported as CsOH. Fission products do not interact of coolant accidents to severe accidents. Transicut with the surfaces of SCDAP components (fuel conditions can be modeled up to the point that the rods, control rods, control blades, and shrouds).
assumption of a predominantly rodlike geometry in The sersion of the code used for these calcula-the core is no longer valid. tions was Cycle 7, with updates. The updates i SCDAP/RELAP5 was produced by incorporating included error corrections that hase been added to models from the SCDAP and TRAP-NIELTA-5,6 subsequent sersions of the code, changing ihe min-codes into the RELAP5/ MOD 2 code. The SCDAP imum liquid fraction in a volume from 1.0 s 104' components model the structures in the reactor core.
to 1.0 x 10 8, and up rading to a newer sersion of Ihe PARAGRASSA code ocrsion 50531), which The TRAP-MELF models were used as a basis for the e Icut tes the fission product release from the fuel, fission product transport and deposition models. The This new sersion also calculated the release of tellu-feedbacks between the various parts of the code were
"?"'Y "'# "E #~
1 deseloped ro pros ide an integral analysis capability. For channel TMLB' calculation.
>! example, the changes m. coolant flow area associated The core damage part of the code does not con-wuh fuel cladding ballooning are taken into consider- sider the interaction of certain structural materials, ation in the hydrodynamics. such as the formation of a Zr-Fe eutectic by zirealoy The RELAPS/ MOD 2 and SCDAP codes were and stainless steel. Interactions between molten described presiously. The fission product behasfor material and the Guid below the core are not explie-includes aerosol agglomeration, aerosol deposi- itly modeled, although enough information is tion, evaporation and condensation, and chemi- available to use a control system to dissipate the sorption of vapors by stainless steel. Fission energy of the molten material. The oxidation of products are assumed to be released equally over stainless steel is not modeled. The control material the entire length of the fuel rods. The released fis- in the control rods is assumed to be Ag-In-Cd; il4 C sion products enter the coolant as aerosols, bein3 control material, such as is used in the full-length put into the smallest size bin and allowed to control rods in Bellefonte, cannot be modeled.
agglomerate or esaporate as conditions dictate. Most of these capabilities either have been added to l The number of aerosol size bins used, as well as the the code since the calculations were performed or fission product species tracked, is selected by the will be added in the near future.
I l
e A-4
References A-1. V. II. Ransom et al., RELAPS/AIOD2 Code 3/anual, lo/umes / aml 2 NUREG/CR-4312.
EGG-2396, August 1985.
A-2. C. N1. Allison, E. R. Carlson, and R.11. Smith, "SCDAP: A Computer Code for Analyzing 1.ight Water Reactor Sescre Core Damage," Pn>ceedings of the International 3 feeling on Light liiner Reac-tor Sen ere Accident Evaluation, Cambridge, Sla, August 23-September I, I933.
A-3. T. C. Cheng et al., "RELAP5/SCDAP - An integrated Code for Sesere Accident Analysis," Pm-ceedings of the Thirteenth librer Reactor Safety Research Information Sleeting, Gaithersburg, AfD, October 12-25,1935, NUREG/CP-0072, pp. 347-355.
A-4. J. W. Spore et al., TRAC-BDI: An Advanced Best Estimate Computer Program for Boiling if bter Reactor Loss-of-Coo / ant Accident Analysis, NUREG/CR-2178, EGG-2109, October 1981.
A-5. 11. Jordan and N1. R. Kuhlman, TRAlt31ELT2 User's Alanual, NUREG/CR-4205, ilN11-2124, Ntay 1985.
A-6. 11. Jordan, J. A. Gieseke, and P. Ilaybutt, TRAft3/ELT User's 3/anual, NUREG/CR-0632, ilNil-2017, February 1979.
A-7. Argonne National L.aboratory, Light-likter-Reactor Safety Research Pmgram: Quarterly Pmgress Refrort, July-September 1981, NUREG/CR-2437 Vol.111, ANL-81-77 Vol. Ill, February 1982.
A-5
APPENDIX B BELLEFONTE PLANT DESCRIPTION .
11-1
APPENDlX B BELLEFONTE PLANT DESCRIPTION llellefonte is a 3600.N!Wtt) llahcock and Wilcox Initial Conditions pressurized water reactor being built by the Tennessee Valley Authority. Figure 11-1 shows a The initial conditions for the RELAP5 and schematie of Ihe reactor coolant system (RCS). SCDAP/RELAP5 calculations were representathe The llellefonte RCS has two hot legs and four of the anticipated operating conditions of the cold legs. The loops are raised to enhance the llellefonte plant at 100% of rated core power.
potential for natural circulation cooling. At the top Table 11-2 compares the values of selected plant of each of the hot legs, before they enter the steam parameters obtained from the llellefonte Final generators, are sent lines. Vents are also located at Safety Analysis Report (FSAR)ll-I for full-power, the top of the reactor sessel and pressurizer. These steady-state operation with those calculated by the vents are normally used to remose noncondensible two codes. The initial conditions were obtained from gases from the RCS during startup. steady-state calculations that used control systems to The two steam generators are of the once- drhe the reactor coolant pump speed and main feed-through design, so that they produce superheated water vahe area to produce the desired loop mass steam during normal operation. Pressure control flow rates and steam generator secondary side mass, for the steam generators is prosided by spring- respectively. The secondary side steam flow was loaded safety relief vahes (SRVs), modulating adjusted to produce the desired cold leg tempera-atmospherie dump sahes (NI ADVs), nonmodulat- ture. The feedwater temperature, pressurizer pres-ing atmospheric dump sahes, and condenser dump sure, and core power were held constant. The vahes. Auxiliary feedwater is provided by three comparison in Table B-2 shows that the calculated pumps, two of whieh are motor-drisen and one of initial conditions were in good agreement with the w hich is steam-drhen. published plant operating parameters. Differences A pressurizer is connected to one of the hot legs, between the desired and computed values are well prosid ng pressure control for the reactor coolant within the uncertainty in the anticipated plant oper-system . The pressurizer spray line is attached to one ating parameters.
of the cold legs. One power-operated relief sahe The initial radial temperature profile used for the (PORV), and two SRVs are connected to the top of fuel in t he RELAP5 t ransient calculations was calcu-the pressurizer. lated using the FRAPCON-211-2 computer code.
The core contains 205 fuel assemblies. Each This radial temperature profile and the core axial assembly is a 17 x 17 array of 264 fuel rods, power profile were based on the Bellefonte FSAR 24 control rod guide tubes, and one instrument published values for the end of the first cycle (460 tube. There are 72 full-length control assemblies, effective full power days) to proside the maximum whose control material is ll4 C. The eight part- fission product inventory. The FRAPCON-2 code length axial power shaping rods use Ag-In-Cd con- was also used to provide temperature profiles for the trol material. three regions in the REI AP5 three-channel core Emergency core cooling is prosided by two high- model at the end of the first cycle.
pressure injection (llPI) pumps, two low-pressure The initial fuel rod temperatures, core soid frac-injection (LPI) pumps, and two core flood tanks tion, and core inlet fluid enthalpy for the SCDAP (accumulators). The core flood tanks and LPI pro- TNILB' calculation were obtained from the single-side liquid directly to the top of the reactor vessel channel RELAP5 transient calculation. The downeomer. The HPI is prosided to each of the SCDAP calculation was started shortly before the four cold legs. RELAPS-calculated maximum fuel cladding sur-Table 11-1 presents salues for sarious automatic face temperature reached 1000 K (1340 F). At that trips and setpoints for the Bellefonte plant. temperature, SCDAP begins to calculate cladding 11-3
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1 Table B-1. Component and control system setpoints for Bellefonte Setpoint Value I on RCS pressure reactor scram, N!Pa (psia) 13.80 (2002)
Low RCS pressure ESFAS, N1Pa (psia) I1.82 (1715)
Reactor coolant pump trip (coincidence of:)
Low RCS pressure, N!Pa (psia) 11.82 (1715)
Low subcooling margin, K ( F) 6.7 (12)
Low pressure injection begins, N!Pa (psia) 1.48 (215)
Pressurizer PORY Opening pressure, N1Pa (psia) 15.93 (2310)
Closing pressare, NIPa (psia) 15.76 (2285)
Pressurizer SRVs Opening pressure, .\lPa (psia) 17.24 (2500)
Closing pressure, N1Pa (psia) 16.48 (2390)
Steam generator NIADVs Opening pressure N1Pa (psia) 8.41 (1220)
Closiag pressure, N1Pa (psia) 8.07 (1170)
Steam generator SRV bank la Opening pressure, N1Pa (psia) 8.87 (1287)
Closing pressure, N!Pa (psia) 8.18 (1187)
- a. Steam cenerator SRV setgwnh me based on rated pressure + 3% accumulation to open and rated pressure -5% bkmann to dose.
oxidation. The initial fission gas imentory was cal- 11-3 show a flow chart of the plant and operator culated by SCDAP. The internal gas pressure at tran- response to the small break transient with no HPl.
sient initiation was calculated by FR APCON-2. The The assumptions and details of the flow path asso-perfect pas law was then used to calculate the pres- ciated with the expected scenario are described sure at the beginning of the SCDAP ealculations.
below. Redundant actions are included in the response because they are present in the sarious ATOG for the S,D Sequence procedures.
The flow chart is entered w hen ihe reactor scrams The Anticipated Transient Operating Guidelines n I w RCS pressure.Section I of the ATOG, (ATOG) are a symptom-based set of procedures "Immediate Actions," then specifies that the reac-designed to lead the operator from the first indica- t r and turbine be tripped. Section 11. " Vital Sys-tion of an abnormal condition through recovery of tem Status Verification," is then entered. The the plant. This discussion of the ATOG will be operators arc here instructed to check that all auto-restricted to their application to the S,D sequence. matic plant responses are functioning and to take I ollowing the initiation of the cold leg break, the appropriate actions if they are not. l'or this see-ATOG is entered when the reactor scrams. The nario, it is assumed that the plant responds accord-operators are instructed to take certain immediate ing to design; that is, the reactor power is actions which essentially serify that the plant auto. decreasing, the control rods are all fully inserted, matic response is as it should be. the main turbine stop salves are closed, the letdow n The response of the operators was based on Resi- flow is minimized, the feedwater has run back, the sion 5 of the lleliefonte ATOG.Il-3 :igures i 11-2 and instruments and control systems base power, and 11- 5
Table B-2. Comparison of computed and desired steady-state parameters Parameter RELAPS SCDAP/RELAPS Desired Core thermal power (N1W)a 3600. 3600. 3600.
Pressurizer pressure (NIPa) 15.2 15.2 15.2 (psia) 2210. 2210. 2210.
Pressurizer level (m) 4.95 4.75 4.95 (in.) 195. I87. I95.
Hot leg temperature (K) 603.9 603.6 6(M.0
(* F) 627.4 626.8 627.5 Cold leg temperature (K) 574.0 573.6 574.1
(*F) 573.6 572.8 573.7 Total loop flow (kg/s) 19832. 19832. 19832.
(lbm/s) 43722. 43722. 43722.
Reactor coolant pump head (NIPa) 0.84 0.89 0.82 (psi) 121. 129. I19.
Steam generator pressure (NIPa) 7.31 7.31 7.31 (psia) 1060. 1060. 1060.
Steam generator liquid mass (kg) 14061. 15314. 15268.
(Ibm) 31000. 33761. 33660.
Steam generator feed (kg/s)b 1034 1034. 1014.
(lbm/s) 2280 2280. 2236.
Feedwater temperature (K) 520. 520. 520.
(O F) 477. 477. 477.
Cesium imentory (kg) 187.
(lbm) 413.
lodine imentory (kg) 10.5 (lbm) 23.1 Tellurium inventory (kg) 24.1 (lbm) 53.1 Xenon imentory (kg) 259.
(Ibm) 572.
Kr)pton inventory (kg) 29.2 (ibm) 64.4
- a. Core adal power shape based on 46n effectise full power day.
- b. Per steam generator. Steam flow set to same salue.
Il-6
Sections I, II Section III 6
Verif y auto trip of RCPs l Trip reactor l or manually trip 4 6 l Trip turbine l l Initiate HPl l 6 4 l Verif y power decreasing l l Maintain proper SG level l 6 6 Verif y all control isolate possible RCS leaks rods on bottom (close pressurizer spray 4 block volve)
Verif y oil main turbine stop volves closed dequate Yes 4 subcooling Verif letdown through I9'" If superheated at bl ck orifice only any time go to ICC No l Verify f eedwater runbock l CFTs Yes 4 empty l Verif y ICS/NNI power on l 6 No Verify station auxiliaries Lock of Yes powered from unit station SG heat service transformer ransf er i
Low pressu e ESFAS No actuation Tube Yes I rupture Verif y HPI, LPI, AFW initiation and MS, MFW No Procedure CP-103 i
k fio vo[v c ed l Verify HPl flow and SG levels l l
No 6 CFTs Yes Verif y AFW initiated emptying and maintainin proper SGleve No p
< Ensure CFT isolation dequate No Verif y RCP outo volves open subcooling -
trip and SG g morgen I evel incre as e
, Open COVs or MADVs Yes g
=
6 l Con trol HPI l Verify heat sink l 6
Continue coofdown dequate No 3 28 K/hr s ubc o olin g RCS No Dum cools down RCP(s)
Yes p Vent RV head for every 28 K of cooldown h f rom ICC L193-KM225-06 l'igure B-2. Diagram of the ATOG response to the S,D transient.
B-7
ICC l Initiate HPI and LPI l Region 2 procedure e
l Increase SG tevels to 50% l Reduce SG pressure to aT..,
of 50-61 K and maintain AT...
6 l Ensure CFT isolotlen velves open l whenever Yes O en RCSg.ressure
>t g PgRV Close PORY when AP = 0.69 MPa No or before entering ICC region 3 2 , 3 Region 3 procedure
~*
t or sat 6- l Stort one RCP per loop l l Stort att RCPs l 6 Reduce SG pressure to AT...
Depressurire SCs as quickly os possible '
6 Open RV and hot leg high Open or verif y open all high " "*"
point vents 0
y Yes E No T=T"' Cycle PORV to maintain V AP of 0.17-0.34 MPo L
,o Yes g
[ Oecrease running RCPs to one per loop 4
Determine 2 Of 3 e
l Waintain Arw and los SC pressure l l Close hat leg dump-to-sump l l or soturo!ed I:
4 Close RV high point vents and pressurizer vents SG heat No 2,
transfer 1 CP-104 l Yes Adequate Yes 2,
subcoolin 1 CP-105 l No b_eP-103
l Close PORV i
l L193-K M221-3) '
Figure 11-3. Diagram of the ICC portion of the ATOG response to the S D tratisient.
2 B-8
i the power has been shifted to the sersice trans- in Section lli, so that CP-103 is entered l3 min after former. At this time, a low pressure emergency the reactor scram.
safety features actuation system (ESFAS) signal The first action in CP-103 is to serify llPI flow will base been generated. The plant control system and the maintenance of the proper steam generator then automatically turns on 11Pi (which doesn't lesels. No action is specified for the absence of 11Pl inject in this sequence), l.Pl and auxiliary feed- flow. The lesels in the CFTs are checked (they base water (AFW), and closes the main feedwater and not changed). The isolation valves for the CFTs are main steam isolation vah es (NISIVs). These actions then verified to be open. A depressurization of the are verified by the operator. If there is flow through steam generators is then begun to cool and depres-the PORV, the block valve is closed. The operators surize the RCS. The N1 A DVs are opened to cool t he then check that AFW flow has been initiated and is RCS at a rate not to exceed 28 K/h (50 F/h). (With maintaining the proper liquid lesel in the steam the N1SIVs closed, only one N1ADV is available to generators [0.61 m (24 in.) on the operating range each steam generator.) The procedure calls for measurement]. If at this time the plant is not sub- opening the NIADVs and/or the condenser dump cooled (it will not be), the reactor coolant pumps vai es (CDVs).11oweser, the CDVs are unavailable (RCPs) are tripped automatically (this actually because the NISIVs are closed. If the RCS does not happens shortly after seram) and the steam genera- cool down as the steam generator pressure for level setpoint is increased to 80% of f ull range. decreases, Ine RCPs are bumped to promote natu-The operator verifies these actions and is instructed ral circulation. The bumping procedure is assumed to control HPI flow according to a procedure to be turning on a pump for 10 s, then turning it (moot point in this sequence). The subcooling mar- off, with one pump being bumped each 15 min. At gin is then checked again to determine the next step the end of each hour of cooldown, the reactor ses-in the procedure. The ATOG indicates that the set head is vented, as defined by another procedure.
operator actions up to this point should be accom- When superheat is detected, the ICC Section is plished in 2 to 3 min. For this analysis,3 min is entered. The operators are instructed to initiate assumed. IIPI and LPI and to increase the steam generator Since the plant is saturated, Section Ill A, " Lack liquid lesels to 80%. The steam generators are then of Adequate Subcooling klargin," is entered. The depressurized until t he difference in saturation tem-operators are instructed to verify that the RCPs hase peratures between the primary and secondary sys-tripped and to manually trip them if they have not tems is between 50 and 61 K (90 and 110 F); this been. H PI is then manually initiated, and the steam temperature difference is then maintained. The iso-generator level is checked and maintained at the lation valves for the CFTs are then verified to be appropriate level (80%). Actions are then taken to open. The operators are then instructed to use the attempt to isolate the break; the only one that affects PORV to depressurize the plant whenever the RCS the plam modelis the closing of the pressurizer spray pressure increases to 15.9 N1Pa (2310 psia). (This block valve. Ihe subcooling margin is then checked did not happen in the transient.) The core exit tem-again; the system is still saturated. The operators are perature is then used to determine what actions are then instructed to go to the inadequate Core Cool-to be taken, based on Figure 4 of the ATOG, w hich ing (ICC) Section if the core exit thermocouples at is included here as Figure B-4. If the plant is sub-any time indiente superheat. For this analysis, the cooled (region 1) or saturated, CP-103 is reentered.
core exit thermocouples were judged to show super-If the temperature indicates that the plant is super-heat when the upper plenum temperature was 2.8 K heated, but the fuel rod temperature is still below
($*F) abose the saturation temperature. There was 1033 K (1400'F)(region 2), the established depres-no superheat at this time. The operators then check surization of the steam generators is continued. If to see if the core flood tanks (CFTs) aie empty (they the temperature indicates that the fuel rod tempera-are not), if there is heat transfer from the primary system to the secondary system 1n the steam genera- tures are between 1033 and 1255 K (1400 and
, , 1800 F)(region 3), what will be referred to as the I tors (there is), and whether there is an mdication of a
" region 3 procedure" is entered. If the core exit l steam generator tube rupture (there is not). The temperature indicates that the fuel rod tempera-operators are then mstructed to continue with procc-dure CP-103, " Transient Termmation Following an tures are above 1255 K (1800'F) (region 4), the Occurrence that L. caves the RCS Saturated with the ]egbn 4 pmcMud, .n entered M. du.s anaW. ,
it was assumed that the ICC region is being contin-SGs (Steam Generators) Removing Heat." It is assumed that it takes 10 min to perform the actions uously m n t red. % hen the region changes, a B-9
Core Exit Thermocouple Temperature ('F) 500 750 1000 1250 18 , , , ,
2500 Region 1 Region 2 Region 3 m
O- 14 - e E -
2000 *iiii S
e e a
a h 10 -
- - superheated - T. . > 1033 K 1500 g g 04004) g M
O M g O 8 -
1000 m
- - T > 1255 K 08007)
Region 4 500 2 ' ' ' ' '
450 550 850 750 850 950 1050 Core Exit Thermocouple Temperature (K) un-mus-os Hgure 11-4. Definition of the ICC regions as determined by the core esit temperature.
1-min delay in entering the appropriate part of the begins. If the plant is in region 2 or 3, the " region 3 procedure was assumed.
procedure"is continued. If the plant is in region 4, in the " region 3 procedure," one RCP in each the " region 4 procedure"is entered.
loop is started. The steam generators are depressur- If the " region 4 procedure" is entered, all the ized to establish a difference of 36 K (100 F) RCPs are turned on. The steam generators are between the saturation temperatures of the primary depressurized as quickly as possible (open all and secondary systems, and this temperature dif- 51ADVs). The hot leg and reactor sessel high point ference is maintained. The reactor sessel and hot vents are opened if they are not already open. The leg high point sents are opened. If there is no indi-PORV, pressurizer vents, and the hot leg dump-to-cation of heat transfer in the steam generators, flow sump are opened. The core exit temperature is then through the PORV is established. If there is heat checked; if it is still superheated, the " region 4 pro-transfer, the PORV is cycled to maintain a pressure cedure" is continued. If it is at the saturation tem-difference of 0.17 to 0.34 N1Pa (25 to 50 psi) perature, one RCP in each loop is turned off. The between the primary and secondary systems. This steam generator pressure is maintained as low as requirement is in conflict with the temperature dif- possible, and the steam generator levels are kept at ference requirement. It is therefore assumed that 80ro of full range. The hot leg dump-to-sump is the pressure difference will be maintained while the closed. The reactor vessel, hot leg. and pressurizer steam generator continues to be depressurized. The vents are then closed. The presence of heat transfer core exit temperature is then checked to determine in the steam generators and the subcooling margin which region of Figure 11-4 the plant is in. Ifit is in are then cl}ccked to determine which recoscry pro-region 1 or is saturated, recovery of the plant cedure to use.
11-1 0
References 11-1 . Tennessee Valley Authority, Bellefonte Nuclear Plant, Final Safety Analysis Report, Docket 50-438, resised through Amendment 22, December 1982.
11-2 . G. A. Berna et al., FAAPCON-2: A Cornputer Codefor tl.e Calculation nf Steady S' ate Thernial-
.\lechanical Behavior of Oxide Fuel Rods, NU R EG/CR-1845, lau uary 198 l .
11-3 . llabcock & Wilcox, Bellefonte Nuclear Plani Units I & 2 Abnornial Transient Operating Guidelines Part I, Document No. 74-1135402, Resision 5 Ianuary 1985.
l 11-1 i
APPENDIX C MODEL DESCRIPTIONS c-i
APPENDIX C MODEL DESCRIPTIONS The input models used to represent the Belle- tor vessel and the top of the hot legs in the fonte plant in the RELAP5, SCDAP, and SCDAP/ Bellefonte pl::nt.
- 2. The Bellefonte steam generators are of a RELAP5 calculations are described.
different design than Oconee. Where Oconee uses aspirated once-through steam RELAP5 Input Model generators, Bellefonte uses the non-aspirated integra' economizer once-tbrough steam generators.
The RELAP5 model of the Bellefonte pressur- 3. Bellefonte has larger diameter pump sue-ized water reactor (PWR) was constructed primar- tion piping than does Oconee. This is ily from information prosided to the Idaho related to the larger system flows that are National Engineering Laboratory (INEL) by the required for the higher power level of the Tennessee Valley Authority (TVA). The informa- Bellefonte design.
tion included detailed blueprints, thermal- 4. llellefonte uses the Babcock and Wilcox hydraulic design specifications, descriptions of design (ll&W) hlark C fuel assembly with most plant subsystems, anticipated operating a 17 by 17 fuel rod array, whereas Oconee parameters, and copies of the Bellefonte Final uses the Ntark B design fuel assembly with Safety Analysis Report (FSAR).C-1 The following a 15 by 15 fuel rod array.
discusses the details of the Bellefonte RELAP5 steady-state, full-power model. The two primary coolant loops for Bellefonte The RELAPS Bellefonte model details all of the cach consist of a hot leg, a steam generator, two
, major system components. Features modeled pump suction legs, tao reactor coolant pumps, and include all major primary system coolant loop and two cold legs. The loop containing the pressurizer,
- reactor sessel flow paths, secondary system main referred to as loop A, was nodalized as shown in
- feedwater paths downstream of the main feedwater Figure C-1. The second coolant loop, referred to as i valses, and the secondary main steam paths loop B, was nodalized as shown in Figure C-2.
j upstream of the turbine stop valves, including the Table C-1 summarizes the relationship between the
! main steam isolation valves. The Bellefonte physical components of the Bellefonte coolant RELAP5 steady-state model utilized 183 control loops and the corresponding mathematical compo-volumes,190 junctions, and 185 heat structures to nents of the RELAPS model.
!- simulate the nuclear steam supply system (NSSS). The pressurizer is connected to loop A as show n l The nodalization scheme used for the Bellefonte in Figure C-1. The pressurizer surge line is con-i RELAP5 model is based on the Oconee RELAPS nected to the hot leg at Component 108 utilizing
! model developed at I N E L.C-2 The Oconee the REL AP5 crossflow model. This approach pre-
! RELAPS model has been used to evaluate small vents partitioning the hot leg flow stream momen-breaks, steam line breaks, and plant operational tum into the pressurizer surge line and artificially
{ and osercooling transients. Significant design dif- clevating the pressurizer pressure. The spray line is i ferences do exist between the Oeonee PWR (177 connected between the pressurizer steam dome and fuel assembly plant with lowered loops) and the ene of thc A-loop cold legs through an inline modu-llellefonte PWR GOS fuel assembly plant with lating valve (Component 616). Primary system raised loops). The rimary differences are: pressure relief is provided by safety relief valves I (SRVs) and the power-operated relief valve (PORV)
- l. The steam generators in Bellefonte are connected to the pressurizer steam dome, ,
raised relative to the reactor vessel, which The four reactor coolant pumps in Bellefonte are I climinates the deep loop seal in the pump llingham Willamette pumps. The single-phase
[ suctions of the oconce design. This tesults head and torque data, as well as the pump rated i in shorter piping runs between the steam conditions, were supplied by TVA. Pump friction j generators and the pumps and longer verti- was based on assumed friction torque equal to
- cal runs of hot leg piping between the reac- 2.25% of rated torque with 1% static friction and I
i j C-3 1
i 616 02 04
[615 ;
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360 700 ..... ..... 165 180 t 12 l 51 1
! 365 705 ,,
135 14 0 14 5 1! 2 15 0 12 5 370 827 l l
,,o ,30 A Loop i 2 ; i 2 i LN86009-1 i Figure C-1. NWalization of reactor coolant loop A for the RELAP5 calculations.
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Table C-1. Correspondence between the physical and mathematical components in the primary loops for the RELAPS model of Bellefonte Physical Component RELAPS Component (s)
Loop A flot leg 100,101,105,108,110 Steam generator inlet plenum 115 Steam generator tubes 120 Steam generator outlet plenum 125 A-1 pump suction leg 160 A-2 pump suction leg 130 A-1 reactor coolant pump 165 A-2 reactor coolant pump 135 A-1 cold leg 170,175,180,181 A-2 cold leg 140,145,150,151 Loop B Hot leg 200,201,205,208,210 Steam generator inlet plenum 215 Steam generator tubes 220 Steam generator outlet plenum 225 B-1 pump suction leg 260 B-2 pump suction leg 230 B-1 reactor coolant pump 265 B-2 reactor coolant pump 235 B-1 cold leg 270,275,280,281 B-2 cold leg 240,245,250,251 Pressurizer Surge line 600,601,605 Pressurizer 610 Pressurizer dome 615 Spny line 620 Spray valve 616 Power operated relief sabe 804 Safety relief vahe 802 Containment backpressure for relief vahes 803,805 1.25redynamic friction at rated conditions. These boiler. The additional solumes did i.nprose the sta-friction values are typical for this type of reactor bility of the heat transfer solution for the steam coolant pump.
generatc tubes during steady-state (al;ulations.
The Bellefonte steam generators are the B&W Oserpressure control for the steam generator sec-once-through design. The primary and secondary ondary side is provided by both screty relief salves sides of the boiler region were nodalized with and modulating atmospheric dumr. valves, with twehe axial control solumes, as opposed to the secondary side kolation prosided by tutbine stop eight solunie nodali/ation used for Oconce. The salves and main feedwater vahes. The relationship additional solumes were used to define the calcu- between the steam generator secondary side physi-lated transition region from the subcooled nucleate cal components and the corresponJing mathemati-boiling regime in the lower portion of the boiler to cal components of the REl.AP5 model are the film boiling regime in the upper portion of the presented in Table C-2.
C-6
Table C-2. Correspondence between the physical and mathematical components in the A- and B-loop steam generator secondaries for the RELAP5 model of Bellefonte Physical Component RELAPS Component Steam Generator A Feedwater line 700,827 Niain feedwater valve 705 Downeomer 300 Tube bundle 310 Steam dow neomer 315,320 Alain steam line 325,330,340,350, 360,370 Turbine stop vahe 365 Ntain steam isolation valve 335 Safety relief vahes 811 Modulating atmospheric dump valves 809 Containment backpressure for relief and dump valves 810,812 Steam Generator B .
Feedwater line 750,927 Ntain feedwater valve 755 Downcomer 400 Tube bundle 410 Steam downcomer 415,420 Alain steam line 425,430,440,450, 460,470 Turbine stop valve 465 Ntain steam isolation salve 435 Safety relief valves 911 Niodulating atmospheric dump vahes 909 Containment backpressure for relief and dump valves 910,912 i
i The nodalization scheme used for the reactor matical compenents is given in Table C-3.
vesselis show n in Figure C-3. The flow between the The auxiliary feedwater system was deseloped to region of the upper plenum inside the plenum cylin- provide level control during the small break tran-der and the region between the plenum cylinder and sients for which the system is assumed available.
the core support cylinder was modeled with two The model includes control variables to determine juactions. The upper head is modeled as a flow- secondary level, time-dependent volumes and june-through solume with a junction representing the tions to simulate auxiliary feedwater injection, and flow from inside the plenum cylinder to the upper trips to initiate the system with best-estimate delays head via guide weldment penetrations. A second for pump turbine power.up and instrumentation junction represents the drain holes that allow flow response times.
from the upper head to the region between the The low-pressure injection (LPI) system was plenum cylinder and the core support cylinder. As developed to simulate best-estimate injection of is common to Il&W design vessels, pressure relief LPI as a function of RCS pressure for the small from the upper plenum to the dcwncomer is pro- break transients. lloth LPI trains were simulated by 5ided by eight reactor sent vahes that are fully open a single time-dependent volume and junction. ,
at a differential pressure of 0.125 psia (862 Pa). The two core flood tanks (accumulators) in j The correspondence between the physical compo- llellefonte were modeled for the small break loss of
- nents ofIhe reactor sessel and the RELAPS mathe- coolant accidents. The ACCUNI component in C.7 4
/////////////
550 u
@ 555 @ 538 @ 535 557, , , g;,,, ,
151 @ 560 @ 540 @ 530 181 . . . . , _ _ _ _ ,
251 y 562 o Ei * "I "---"-"i +!u + l 281
@ 565 @ 525 567, , 545
@ 1 520 @
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3 4
5 h 5 4 515
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f I
570 5 3 3 6 2 2 7 1 1 b o o i
l
@ 8 505
@ 9 a
Ir 575 f//////////
LN86009-3 Iigure C-3. Nodali/ation of the reactor sessel for the single-channci RIit AP5 calculations.
C-8 l
Table C-3. Correspondence between the physical and mathematical components in the reactor vessel for the RELAPS base case model of Bellefonte Physical Component RELAP5 Component (s)
Inlet annulus 555,557,560,562,565,567 Dow neomer $70 Lower plenum 505,575 Core 515 Core bypass 510 Upper plenum 520,525,530,535,538,540,545 Upper head 550 Vent valve 536 RELAP5 was used to simulate the accumulator assemblies in Region I (center channel),
I tank, surge line, and check valse. The system is 19.3 kW/m (5.9 kW/ft) for the 100 fuel assem-i passise and does not require trips or control varia- blies in Region 2 (middle channel), and i
bles for initiation. 12.9 kW/m (33 kW/ft) for the 44 fuel assemblies
- Heat structures were used throughout the model in Region 3 (outer channelt to characterite the heat capacity of and resulting A nodalization study was performed to deter-
! heat transfer from major structural masses in the mine if three channels in the core were sufficient to j Bellefonte PWR. Included were structures to model model the expected aatural circulation flow. A j loop piping; steam generator tubes, shrouds and model of the reactor sessel was used, and a high i vessel; pressurizer walls and heaters; reactor vessel pressure boiloff transient was simulated. Compari-walls, core barrel, core baffles and former plates, sons were made between models with three and five plenum cylinder, internal supports and guide strue- parallel channels in the core and upper plenum.
tures; and the nuclear fuel. The outside of the pip- For the five-channel model, the two inner chan-ing and sessel surfaces were treated as adiabatic nels in the three-channel model were cach divided boundaries, eliminating environmental heat losses. into two channels. The resulting average linear heat The core fuel heat source was modeled using the generation rates were 20.0 kW/m (6.1 kW/ft) for RELAP5 point kinetics package with scram signal- the 13 fuel assemblies in Region 1 (center channel),
t initiated control rod insertion simulated. 19.7 kW/m (6.0 kW/ft) for the 48 assemblies in The nodalization of the three-channel core is Region 2,19.7 kW/m (6.0 kW/ft) for the 50 shown in Figure C-4. Three parallel flow paths assemblies in Region 3,18.9 kW/m (5.8 kW/ft) 3' were modeled in the core with volumes at the same for the 50 fuel assemblies in Region 4, and elesation connected using crossflow junctions. Fig- 12.9 kW/m (3.9 kW/ft) for the 44 assemblies in ure C-5 shows the regions of the core hat the chan- Region 5 (outer channel). The results of the boiloff nels represent and Table C-4 gives the calculation showed little difference between the correspondence between the physical components three- and five-channel results. The calculated tem-and the RELAPS mathematical components in the peratures during the entire transient were nearly core. The core average linear heat generation rate identical, as were the flow patterns and magni-2 for the single-channel model was 18.0 kW/m tudes. Based on these results, it was decided that (5.5 kW/ft). The average linear heat generation three channels were sufficient to approximate a ratc was 19.8 kW/m (6.0 kW/ft) for the 61 fuel two-dimensional model.
, C-9
l l
l 520 l l l 507 514 521 l l l ... .. ...
506 613 N h 619 l l l ........ ..
504 % 512 % 518 510 l l l .....
603 511 517 l l l ..............
k 502 609 516 l l l ... .........
501 k 508 515 l l l l 605 L198-KM2:5-Og I'igure C-4. Nodalization of the core for the three-channel REl.Al*5 calculation.
C-10
l l l l l l l Region 1 _I I I I N s Region 2 -
l Region 3 l I I
! I L193 KM226 08 l'igure C.5. Crou-sectional siew of t!'e regions in the three-channel core modch.
SCDAP Input Model square lattice to apprmimate the shape of a e>lin-der. A single fuel awembly was modeled along with a portion of the sewel upper plenum. Figure C-6 1 he reactor sessel was modeled for the SCDAP shows the nodalization used for the SCDAl' model code primarily utili/ing specifications outlined in of the llellefonte reactor sewel. The model of the the llellefonte 1:SAR 0-1 Other information used fuel awembly will be discuwed first, followed by a to construct th
- model included blueprints, design description of the upper plenum model, specifications, and anticipated operating parame. The il&W .Niark C fuel awembly consists of ters that were prosided to INEl. by TVA. 264 fuel rods, 24 control rods, one instiument The core of the llellefonte NSSS consists of tube, spacer grids,and end fittings. A 17 x 17 array 205 il&W Niark C fuel assemblics arranged on a of fuel rods and tubes is formed within a structure Table C-4. Correspondence between the physical and mathematical components in the core for the RELAP5 three channel core model of Bellefonte Ph> sical Component REl.Al'5 Componenth)
Core Region I (61 fuel awemblic4 501,502,503,504,506,507 Core Region 2 (100 fuel awemblic4 508,509,StI,512,513,514 Core Region 3 (44 fuel awemblies) 515,516,517,518,519,521 C-I l
l Upper planum zone
- - Plenum cover 6
4 3
Elevation above core 2 Core bottom, ~~ Uppor core plate nodo m {ft) 3
- *- Top of cladding 10 3.6 (11.7) 9 3.2 (10,5) 8 2.0(9.2) 7 2,4 (8.0) 6 2.1 (0.8)
- & Instrumentation tubo 6 1,7(b.6) 4 1.3 (4.3) f - M Fuolrods 3 0.9 (3.1) $- l
.5 .
2 0.6(1.0)
- .- @ Controf rods Coro intot L193.KM 226 10 f Ilyuic C-6. Nodalitation of the reactor scucl for the SClul' calculatuut 1
C 12
of spacer grids, end fittings, and guide tubes. A awembly. T he region between the upper core plate crow section of a t>picalll&W Mark C fuel awem. and the plenum soser was disided into four asial bly is show n in I igure C 7. Three components were /ones, used to model the fuel awembly: a fuel rod compo-nent representing the 264 fuel rods, a control rod component representmg the 24 control rods, and a SCDAP/RELAP5 Input Model third component representing ihe single instrument tube. All of the 264 fuel rods were modeled in a single component, caeh with core-aserage power, The input model for the SCDAl'/RI!I.AP5 cal-because only an aserage power rod was modeled in culations was a combination of the Riil.AP5 and the single-channel RI!!. Al'$ calculation and SCDAP models described presiously. The heat SCDAP codeinput was dependent upon that caleu- structure for the fuel rods in the R[it.AP5 model lation. I he fuel rods, control rods, and inst rument was replaced by SCDAP components representing tube were disided into ten equal asial segments the fuel rods, control rods, and instrument tubes, called nodes. For the fuel rods, sis radial nodes The nodali/ation in the core was changed from were used; for the control rods, two radial nodes; sis to ten asial nodes. T his change was aho made in and for the instrument tube, four radial nodes, t he cor responding elevations in t he dow neomer and Table C-5 presents the relation between the asial core bypaw. lor the three channel model of the nodes in the core and the height abose the bottom core and upper plenum, the three regions in the of the fuel for the sarious modeh. core for the three-channel Riii.AP5 model were for the upper pienum. only the region inside the estended into the upper plenum. In each of the plenum e>linder abose a single fuel awembly was channel , SCDAP components were used to model modeled. 't his region was disided into fise asial the fuel rods, control rods, and instrument tubes.
iones as shown in Figure C 6. The bottom ione Iigure C H shows the nodalitation of the reaetor included the upper end fittings and upper grid sewel for the three-channel model.
O O O O O contrai red guide tube O O O O O O O O O O
/
/
O-'O O O O
/
/ O O in.trum.nt.'I O O O - - - M iy u... ...n Iigme C 7. Cromestional siew of a typical Alark C fuct awembly.
C-13
Table C-5. Relation between the core axial nodes and the height above the bottom of the fuel for the three code models Height Above 130t:om of Fuel, m (ft)
Core Axial Node RELAP5 SCDAP and SCDAP/RELAPS I 0.3 (0.9) 0.2 (0.6) 2 0.9 (3.1) 0.6 (1.8) 3 1.6 (5.3) 0.9 (3. I) 4 2.3 (7.6) 1.3 (4.3)
$ 3.0 (9.8) 1.7 (5.5) 6 3.5 (l 1.5) 2.1 (6.8) 7 2.4 (8.0) 8 2.8 (9.2) 9 3.2 (10.5) 10 3.6 (l 1.7)
Since interactions hetween molten core material ty pes and numbers in the hot legs had to be changed and fluid in the lower plenum are not yet explicitly to allow the connection of these lines; the geometry modeled in the code, a control sptem was used to of the hot leg piping was not changed. The piping simulate them. A heat structure was placed in the dow mtream of the PORV and the SRVs on the pres-lower plenum bolume number 505 on Iigure C-3) suri/cr was also modeled to the reactor coolant drain that had the surface-loaolume ratio of a 2.54 em tank.
(l-in.) sphere. Ihe heat source p osided to Ihe cen- 1he steam generator nodalization was also ter of the structme was Ihe amount of energ> abose changed for the S D calculatiom. The tubes were split the saturation temperature contained in the mate-between two parallel channels on both the primary tial that had relocated below the bottom of the and the secondary sides in order to better model the core. This energ) was Ihen tramferred to the cool-ausiliary feedwater injection. One channel repre-ant by the Rlil AP5 heat conJuetion and comec- seated the 10ro of the tubes that are affected by the tion models. injection of auxiliary feedwater. These tubes were Sescralchanges were made to the Rlil AP5 model awumed to be the outer loro of the tubes in the bun-f or the S;D operator action imestigatiom. The high die. The other channel represented the 90re of the point sents from Ihe hot legs,Ihe reactor sessel sent, tubes on the imide of the bundle. The auxiliary feed-and the sent line f rom the prewuriier were all mod. water was injected into the smaller riser section. Ihe eled, t he ef fluent hom these sents flowed to the two parallel channeh were connected by crowllow reactor coolant drain tank, w hish was aho modeled. juactiom on the riser side. On the primary side, the t he ruptme disk on the drain tank was modeled two channeh were both connected to the inlet and ming a trip sahe, connecting the drain tank to the outlet plena of the steam generators. Iigure C-10 containment. Ihc Il-loop hot Icg domp to-sump was shows this mxlati/ation for steam generator A.
modeled. I hese components w e c modeled in antici-Ihe p(mer input to the code is in table form. Decay pation of their me during the inadetpiate core cool. heat was calculated by the code at 20 points in the ing pnwedure.1 igure C.9 show s the nodali/ation of transient Ohe masimum number of points alhmed),
the sent lines and dump toaump. Ihe component with knear interpolation between the points.
550 l l l l 524 -
528 -
532 538 I I I I 523 -
527 -
531 540 i i I I
522 -
526 -
530 i i i 545 520 -
525 -
529 I
650 -
670 -
690 649 -
669 -
689 648 -
668 -
688 i
647 -
667 -
687
- i , .............
3
- 646 -
666 -
686 510 645 -
665 -
685 j 644 -
664 -
684
- i i i .............
i 643 -
663 -
683 j i , , .............
i 642 -
662 -
682
- i , , .............
- 641 -
661 -
681 l l l l l 505 L193 KM225 01 I
ligure C-H. Nodalization of the core and upper plenum for the three-channel SCDAP/RELAP5 calculation.
j
! C-15
4 0
a 2
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I 2 389 2 399 3 388 -
398 3 4 387 -
397 4 320 ------------ ' ' ------
5 386 -
396 5 6 6 385 -
395 .
121 122 2 ------------ i i ------
7 384 -
394 7 8 383 -
393 8 3 ------------ o I -----
9 382 -
392 9 4 10 381 -
391 10 330 11 380 -
390 11 700 , .. . . . . . . . . . . . , , .
300 12 305 12 t I 12 5 L193 KM225 02 l'igure C.10. Nodalization of the split steam generator for the S 2D operator action calculations.
C-I7
References C- l .
Tennessee Valley Authority, Be//efonte Nuc/ car Plant, FinalSafety Analysis Report, Docket 50-438, resised through Amendment 22, December 1982.
C-2. C. D. Fletcher et al., RELA PS Therntal-Hydmu/ic Analysis of Pressurized Thennal Shock Sequences for the Oconee-I Pressuri:ed likter Reactor, NUREGICR-376l, EGG-23l0, June 1984 C-18
- At PO"I 'out8Bt R dd ** f 80C *es per me , .t ears u 3. aswCLE.A st0ukATORv ;
leAC FOmas L' "l. . ..,'
NUREG/CR- 41 3mi 3m BIBLIOGRAPHIC DATA SHEET EGG-247 116 st$'avC? ION '% 78*E #EvtRSE J LE Avt SLA
- T8T Lt 4%O sve t T s Feedwater ansient and Small Break Loss of Coolant Acci 'nt Analyses for the Bellefonte # ' ' " ' " " "
Nuc1 ear P1 ant ,0,,, ' ., "" " "I , . . ,
[ March,,,,,,,,,0,,,,,,1987 Pa"uT.D. Bayless ,
041. .t.
Charles A. Dobbe g j Rosanna Chambers G ADO *15S itarmes te Come
} March e PROJECitT Asa woma whit mvwSta 1987 i Ptm*OmwiNG On ..aeet.fiON 4.wg A%D we EG&G Idaho, Inc. ,,,,,,,,,,,,,,_,,,
Idaho Falls, Idaho 834 A6354 ines te Cases steYtPt08*tPont 10 SPO450#i4G OMG.%.d4 TiO% %AWt 4%D W4*LinsG ADOptSS e Division of Accident Evaluati ' .
Research Office of Nuclear Regulatory Re arch ' " " ' ' '' "~~~~~
U.S. Nuclear Regulatory Commissi Washington, D. C. 20555 4 /
12 WP*L ewtant aa. NOf tl
,..s,..C,,,w..,
k Specific sequences that may lead to re da'ntage were analyzed for the Bellefonte nuclear plant as part of the U.S. Nucilar Regulatory Commission's Severe Accident Sequence Analysis Program. The IWLAPS, SCDAP, and SCDAP/RELAPS com-puter codes were used in the analyse [The two main hitiating events investigated were a loss of all feedwater to the stcah generators and 'n small cold leg break loss of coolant accident. The transients air primary interest withjn these categories were the TMLB' and S D2 sequences. VariMions on systems availability were also investigated.
Possible operator actions that coMd prevent or delay core da'niage were identified, and two were investigated for a small break transient. All of thehansients were analyzed until either core damage begar3yr long-term decay heat remoh was established. The analyses showed that for the sequences considered the injectio(flow from one high-pressure injection pump was flecessary and sufficient to prevent; core damage in the absenceof operatoractions.p* actions were damage S D sequence; no operator available in the to prevent cperator 2
TMLB' sequence.
s -
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feedwater transient, small Nreak LOCA, Bellefonte i
' 't SECunity CLASS +8sC ATGN i
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EG&G Idaho P.O. Box 1625
/daho Falls, Idaho 83415
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