ML20217G217: Difference between revisions
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| document type = CORRESPONDENCE-LETTERS, INCOMING CORRESPONDENCE | | document type = CORRESPONDENCE-LETTERS, INCOMING CORRESPONDENCE | ||
| page count = 77 | | page count = 77 | ||
| project = | | project = TAC:M95144, TAC:M95145 | ||
| stage = Other | | stage = Other | ||
}} | }} | ||
Latest revision as of 01:42, 21 March 2021
| ML20217G217 | |
| Person / Time | |
|---|---|
| Site: | Sequoyah  |
| Issue date: | 04/06/1997 |
| From: | Shell R TENNESSEE VALLEY AUTHORITY |
| To: | NRC OFFICE OF INFORMATION RESOURCES MANAGEMENT (IRM) |
| Shared Package | |
| ML19317C563 | List: |
| References | |
| TAC-M95144, TAC-M95145, NUDOCS 9708070126 | |
| Download: ML20217G217 (77) | |
Text
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Tennessee Vaney Authority, Post Office Box 2000, Soddy-Daisy, Tennessee 37379 2000 [%y .jk..pt j NJ['!f 7 : . $f %. .. p_ t , , b :.;;L g c u [P%e . a e m.- Mis
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April 6,1997 M:, .g .i.?
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U.S. Nuclear Regulatory Commission fg,"jhg ATTN: Document L:,ntrol Desk Washington, D.C. 20555 Y.y5$$j
& c. ben;J; Gentlemen: g(i t . .!:# 3.. ,d+v. - .m In the Matter of 5:MDb.[
) Docket Nos. 50-327 .- y n Tennessee Valley Authority ) 50-328 5.[9/gh4
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SEQUOYAH NUCLEAR PLANT (SON)- RESPONSE TO REQUEST FOR ADDITIONAL . b - 60 INFORMATION - TECHNICAL SPECIFICATION (TS) CHANGE REQUES'B6-01 ON MO CONVERSION TO FRAMATOME COGEMA FUEL (TAC NOSM95144 AND M95145) ?sh..;.,Q= .
References:
- 1. NRC letter to TVA dated January 8,1997, on the above subject h'9 .. . . . ;p)..m): .
re i$fC:%
- 2. TVA letter to NRC dated February 7,1997, on the above subject
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. :, m.;w.. e.
m l 3. TVA letter to NRC dated March 17.1997, on the above subject 's.% $
;s%W.' s::.i. <
- 4. TVA letter to NRC dated March 20,1997, on the above subject
' N. . sNo.
- 5. TVA letter to MC dated March 25,1997, on the above subject A3 sq.-:.g&? -cp
.1s.k Ug.J. r q#
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- 6. TVA letter to NRC dated April 1,1997, on the above subject
!.Q],:g%G e h The purpose of this letter is to provide a single letter containing a complete response to I
+
questions received from NRC in reference 1 and subsequent telephone calls on the above subject. nW , h,f:}.jp.,j
- x. ,
1 I . . .t..h l As requested in reference 1 and telephone calls with NRC, references 2 through 6 were g. ' 3 provided to facilitate a timely review of TS 96-01 which was submitted on [ gl;gf;j%!J April 6,1996. At NRC request, this submittalisbeing provided to replace references 2 4-.t9d 1 through 6, with respect to resolving NRC questions, to provide a single reference. Note I / that reference 6 will continue to be the submittal to use in regard to the TVA commitment for TS 96-01 and the retyped TS pages, ; q . g oe . ,m n n n o. c _I)uf '(pA, 0o . n' 3 0" .A te -' pyggg th lll]gj.g.gggg
- i E'
s_ _om _ 9708070126 DR 970406 I I ' LtNII 4 p ADOCK 05000327
'( l PDR j a
Al _m A W ,Y f -
- sy n mINSI y j. _
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a...,,., . - - , . - - - - - . . - . .. ., U.S. Nuclear Regulatory Commission
. Page 2 _
April 6,1997
. Enclosure 1 contains the proprietary responses to reference 1 and additional NRC-questions. Enclosure 2 contains the non-proprietary responses.
Since Enclosure 1 provides information which is proprietary to Framatome Cogema Fuels (FCF), the application for withholding and affidavit signed by FCF, the owner of
-the informat'on, which were provided in references 2,3, and 5 are still applicable. The application for withholding and the affidavit set forth the basis on which the information may be withheld from public disclosure by NRC and address with specificity the considerations listed in Section 2.790, Paragraph (b)(4) of the NRC regulations.
I '
- Accordingly, it is respectfully requested that information which is proprietary to FCF be withheld from public disclosure in accordance with 10 CFR, Section 2.790, of the NRC L regulations, i
I . Correspondence with respect to the proprietary aspects of Enclosure 1 or the supporting FCF affidavits should be addressed to J. H. Taylor, Manager of Licensing Services, Framatome Cogema Fuels, P. O. Box 10935, Lynchburg, Virginia 24506-0935. Please direct questions concerning this issue to Keith Weller at (423) 843 7527. Sincerely, [)n R. H. Shell
~g
! Site Licensing and industry Affairs Manager Enclosures cc: See page 3 o
m..
. m U.S. Nuclear Regulatory Commission Page 3 April 6,1997 cc (Enclosures):
Mr. R, W. Hernan, Project Manager Nuclear Regulatory Commission One White Flint, North 11555 Rockville Pike Rockville, Maryland 20852 2739 NRC Resident inspector Sequoyah Nuclear Plant 2600 Igou Ferry Road Soddy-Daisy, Tennessee 37379-3624 Regional Administrator U.S. Nuclear Regulatory Commission Region ll 101 Marietta Street, NW, Suite 2900 Atlanta, Georgia 30323 2711 i l i
1 ENCLOSURE 2 RESPONSES TO NRC OUESTIONS NON PROPRIETARY VERSION e
y REOUEST FOR ADDITIONAL INFORMATION l TENNESSEE VALLEY AUTHORITY SEOUOYAH NUCLEAR PLANT, UNITS 1 AND 2 DOCKET NUMBERS 50-327 AND 50-328
- 1. Revision 2 to the BAW-10168 evaluation model was modified during the staff review with regard to the Moody break flow model and discharge coefficient. Please verify that the analysis that was used to support.this Sequoyah Nuclear Plant (SON) fuel change and Technical Specification (TS) amendment was performed using the approved model.
Rosponse-Since the time that the small break LOCA analysis of BAW-10220 l was performed, several modifications were made to FCF's small l . break LOCA evaluation model, BAW-10168, Revision 2. The changes l were made in response.to the NRC review of the document. Chief l f among these modifications - with respect to the SQN small break l . LOCA analyses - are changes to the break discharge coefficient l ) and the nodalization of the pump suction piping. The original l evaluation model implemented a variable 0.7/1.0 discharge l coefficient, the approved model requires the use of a constant l 1.0 discharge coefficient and the upflow side of the pump suction l piping is more finely divided to better predict the occurrence of l loop seal clearing. l
}
As a result of FCF's efforts to bring the small break analysis l into compliance with NRC-approved methodology, additional l problems with leak node modeling were identified and resolved l (see the response to NRC question number 8 in this set). The l small break spectrum was rerun using consistent homogeneous leak l junction inputo, a nonequilibrium fictitious leak control volume, l
-and the containment and leak volume areas set equal to the cold I leg pipe area. l l
Analysis of a spectrum of small breaks predict a modest clad I temperature excursion, approximately 1162 F, for the-limiting, l 2.75-inch, break case. The work fully complies with the l NRC-approved SBLOCA evaluation model, as. documented in BAW-10168, l Volume II. No evaluation model exceptions have been-taken and no l evaluation model difficulties were encountered in the process of l analyzing this event. I l I FCF Non-Proprietary
- 5. 9. 2- Small-Break LOCA Evaluation Model .l l
RELAP511s used to predict the reactor coolant system j: thermal-hydraulic _ responses to a-small break LOCA. The code -l has been approved by the NRC for licensing application and ; is documented in detail in Reference 5-3. 'RCS nodalization is based on the model described-in Volume II of the NRC- l approved BWNT RSG LOCA EM, Reference 5-1. Nodal diagrams of l the 1K)N small break LOCA model are ~ presented in Figures 5.9- l 1 and 5.9-2. The small break LOCA model is similar to that j used in the large break analyses. ! The reactor core is divided radially into two regions l similar to that of the LBLOCA model; one region represents l the hot fuel assembly and the_other represents the remainder l of the core. The core is further divided into twenty axial l segments. Cross-flow junctions connect hot assembly fluid ; nodes with the adjacent " average" assembly nodes. This ! arrangement allows the computation of hot assembly cladding l and-vapor temperatures with limited influence by coolant l
-from the average core and provides resolution of the mixture l level to withinlapproximately 0.5 foot. Initial fuel pin l L parameters are calculated with-the TACO 3 computer code l (Reference 5-6) . The reactor vessel downcomer and upper l plenum regions are represented in finer axial detail than !
those of the large break LOCA model to give a better i representation of the void distribution that affects the ; system hydrostatic balance. l l In the small break LOCA modeL the RCS is subdivided into l i two flow loops. One loop represents the broken loop and the l other represents the three intact loops. The pressurizer is attached to the hot leg of the composite intact loop. -The l nodalization is similar to that of the large break LOCA -l model. ! l The steam generator tube region is divided-into two radial l regions. -One region represents the shortest half of the l tubes and the other region represents the remainder of the l tubes. This provides sufficient modeling accuracy to l simulate tube-draining effects; tube draining can-be l sensitive to tube length. l
. l The reactor coolant pump suction nodalization has been i altered relative to the large break LOCA model to produce-an ;
accurate hydrodynamic representation of loop-seal clearing. l Two additional nodes are added to the downside of the pump l , suction pipe and,three nodes to the riser section. This ! allows finer resolution of the v.cid distributions and l FCF Non-Proprictny ,
-. --- - .- -. - -- . - _ ~ . . - . . . . -
elevation heads that control the occurrence and timing of ;
~1oop seal clearing. l The' bottom alevation of the lowest node of the intact loop ;
pump suction piping is artificially extended one. foot below l the corresponding node in the broken loop. This _ -l preferentially promotes the clearing of only the broken ! loop. An RCS. configuration characterized by a single clear j
. loop and three intact loops conservatively resttlets steam l flow to the break. The added restriction can result in ;
worsened core conditions and the potential uncovering of the l core. ! l Both the broken loop and the intact loop reactor m: slant l pump discharge piping are modeled as four nodes. In the l large brea'k model, the intact loop is modeled as one node. l Using four nodes provides an accurate simulation of the l hydrodynamic effects of the ECCS injection. l I The. computer code options and generic input requirements ! ] used in the small break portion of the BWNT RSG LOCA EM are ! summarized in Volume II, Tables 9-1 and 9-2, of Refet ace 5- l
- 1. Consistent with the RCS loop modeling, the phase non- l
}1 equilibrium option is selected for the artificial leak i volume, node 276. The break path is treated homogeneously- l for both critical and non-critical break flow predictions. l In addition, the volume area of the break mass sink volume !
is set equal to the volume area of the artificial leak node i to preclude the development of any contribution to the momentum flux gradient at the break junction. The small break LOCA model is fully compliant with the NRC- l approved guidelines af the BWNT RSG LOCA EM established in l Reference 5-1. It has been developed utilizing.the RELAP5 : large break LOCA model, described in Section 5.2 of this l report, as a basis. The small break model adheres to the ! requirements of 10CFR50 Appendix K and contains demonstrated ; conservatism for the evaluation of-ECCS mitigation of a l postulated small break at SON. l l 5.9.3 -Inouts and Assumntions l
. l The major-plant operating parameters u. sed in the SQN small l break LOCA analyses follow. These ipputs are similar to l those utilized in the large break analyses. I
- 1. Power Level - The plant is assumed to be operating ;
in steady-state at 3479 MWt (102% of 3411 MWt).
, e FCF Non-Proprietary 1
h
=.-1 ,-.. -m , - _e_ . . ,.-
l
- 2. Total System Flow - The-initial Reactor Coolant ! ;
System (RCS) flow is' 348,000 gpm. ; l l
- 3. Fuel Parameters - The initial-fuel pin parameters: - ; i are.taken from TACO 3 (Re f e rence _ 5-6) runs 'l
_ performed for BOL' fuel conditions. ; l
- 4. ECCS The ECCS-flows are based on the assumption l 1
of a single active failure. A single train of =l ECCS is modeled as described in Volume IT, Section l
'4.3.2.2, of_the BWNT RSG LOCA EM, Reference 5-1. l For the case of a centrifugal charging line break, l charging flow is, assumed to be spilled to the l containment. l l
- 5. . Total Peaking Factor (Fo ) - The maximum total l
peaking factor assumed by this analysis is 2.5. l The hot assembly peaking for small break analysis 1 is illustrated in' Figure 5.9-3.. l I l
- 6. The moderator density reactivity coefficient is' !
based on BOL_ conditions to minimize negative l- l reactivity. l l
- 7. The, cladding. rupture model is based-on NUREG-0630. l l
5.9.4 Analysis Results ! l In the SQN small break LOCA analyses, seven brea'k cases were l considered independently to predict core and system l responses over a spectrum of break sizes. Small break l spectrum results in Volume II, Appendix A, Reference 5-1, j indicate that break areas corresponding to 2- to 6-inch l diameters produce the most severe core depression. Breaks l of 1.5 , 2.5 , 2.75 , 3.0 , 3.25 , and 5-inch' diameters in l the bottom of the reactor coolant pump discharge piping were_ l analyzed for-SQN. In addition, a 1. 34-ir ch diameter l centrifugal charging line break, located in the top =of the l piping,-was analyzed. l l Table.5.9-1 presents time sequence of events for each of the ! small break LOCA cases. Fuel thermal ' responses for the hot l pin are included in Table 5.9-2. Parameters of Interest to l the small break analyses are shown in. Figures 5.9-4 through- l _ 5.9-28. There are seven sets of figures, each set contains l five_ plots. The five figures of each set show (1) the RCS l pressure,_ (2) the break flow rate, (3) loop seal levels in l the pump cuction,downflow and upflow pipes, (4) core l collapsed level, and (5) hot spo.t cladding temperature. l FCF Non-Proprietary
I A_relatively slow depressurization rate occurs in the 2.75- l inch break case. The core does uncover, making the 2.75- l inch break the most limiting case of the small breax l spectrum analysis. The resulting peak cladding temperature l (PCT) is 1162*F. Core metal-water reaction for the 2.75- ! inch break is negligible because the cladding oxidation rate l 1s not significant below about 1500 F. l
\
For breaks smaller than 2.75 inches in diameter, core l cooling is maintained by a combination of steam relief at l the break and reflux cooling in the steam generator. The l core does not uncover for these smaller breaks. For breaks l larger than 3 inches in diameter, the rapid depressurization l rate following loop seal clearing has twc positive effects l on the core level. One is increased ECCS flow, and the l other is increased core level swell. No core heatup was l predicted for these break sizes. l I The centrifugal charging injection line break is postulated l to allow examination of a small break LOCA that is ; characterized by a degradation of high pressure injection. ! The break size is insufficient to allow significant l {' depressurization of the RCS. Coolant addition in the l progression of the transient is, therefore, governed by the l high pressure injection alone. With a broken injectico l line, a large_ portion of the ECCS flow associated with the l centrifugal charging flow is directed to the break. To l ensure a conservative result, all of the charging flow is l assumed lost to the break, and the transient is mitigated by l safety injection pumps only. The results of the charging i line break indicate that the core remains covered by the l 3 mixture level and that no core heatup occurs. l l The small break analyses are terminated when the break flow l rate is exceeded by the ECCS flow rate. Note that the l collapsed liquid level at the end of the transient may still l be below the top of the core. The core mixture levels at l the end of the analyses are, however, above the top of the l active core and the RCS pressure is still falling. A steady l increase in ECCS injection and continued core cooling is l therefore assured. l l 5.9.5 Comoliance to 10CFR50.46 - l l The small break calculations directly demonstrate compliance l to two of the criteria of 10CFR50.46 and serve as the basis l for demonstrating compliance with two others. As seen in l
, the figures and ,in Table 5.9-2, the highest peak cladding l i,
FCF Non-Proprietary
i l temperature, 1162'F, and the highest local oxidation, about l l 0.004%, are well below the 2290*F and 17% criteria. l 4 The whole-core oxidation criterion of 1% cladding reaction l is met as well in the small break LOCA analyses. Whole-core ! oxidation willibe much less tnan the peak-local oxidation l . figure of 0.004%. Whole-core oxidation associated with ! l small break LOCAs, utilizing the assumptions and inputs as l documented above, is negligible. l , i l The fourth acceptance criterion of 10CFR50.46 states that' l calculated changes in core geometry shall be such that the l core remains amenable to cooling. The calculations in j Section 5.9 directly assess the alterations in core geometry l that result from the LOCA, at the most severe location in l l the core. These calculations demonstrate that the fuel pin l cooled successfully. Further, for SQN, no hot assembly l cladding ruptures occurred in any of the small break LOCA l cases. Therefore, the assembly retains its pin-coolant l , channel-pin-coolant-channel arrangement and is capable of f l being cooled. l l t' The fifth acceptance criterion of 10CFR50.46 states that the l calculated core temperature shall be maintained at an ! acceptably low value, and decay heat shall be removed for l the extended period of time required by the long-lived
, radioactivity remaining in the core. Successful initial
- operation of the ECCS is shown by demonstrating that the l core is quenched and the cladding temperature is returned to ,
near saturation temperature. l l Compliance to the long-term cooling criterion is ! demonstrated for the systems and components specific to SQN l in the FSAR and is not related to the fuel design. The l initial phase of core cooling has been shown to result in l low cladding and fuel temperatures. A pumped injection ! 4 system capable of recirculation is available and operated by l the plant to provide extended coolant injection. Therefore, ! compliance with the long-term cooling criterion o 10CFR50.46 has been demonstrated. ' l 4 FCF Non-Proprietary 1 4
Table 5.9-1 Small Break LOCA Time Sequence of Events l l l 1.5-Inch 2.5 luch 2.75- 3.0-inch 3.25- 5. 0-in ch CCI IJne l Events, Seconds Break Break Inch Break inch Break Break l Break Break l Break Initiation 0.0 0.0 0.0 0.0 0.0 0.0 0.0 l Reactor Scram 91.5 33.4 27.6 23.4 20.2 10.2 115.1 l RC Pump Coastdown 91.5 33.4 27.6 23.4 20.2 10.2 115.1 l MSIV Closed 91.5 33 4 27.6 23.4 20.2 10.2 115.1 l l SI Signal 104.6 40.7 "34.0 28.6 24.4 10.1 l29.7 l FFW Isolation 101.5 43.4 37.6 33.4 30.2 20.2 125.1 l Pumped ECCS Injection 14 L .6 77.7 7 l .0 65.6 6 L .4 47.1 166.7 l leop Seal Clearing NA 1267.8 989.5 784.9 636.1 247.3 NA l Top of Core Uncovers NA NA 2370 20'i0 1690 210 NA l Peak Cladding NA NA 2914 2295 NA NA NA l Temperature l l Accumulator injection NA NA 2895 2395 1830 490 NA l
\ 4 l
I ( l Table 5.9-2 Small Break LOCA Results l l l I Results 2.75 Lach 3 Inch Break l Break l Peak Claddmg Temperature."E 1162 828 l Peak Temperature Locanon, ft 10.9 1 l .6 l Rupture Time, sec NA NA l Rupture Iscation, ft NA NA l Maximum L.ocal MtW Reacnon, % ~0.004 ~0 l Total MlW Reaction, % <0.004 ~0 l l l l FCF Non-Proprietary
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Figure 5.9-2 REIAPS Small Break LOCA Model - Sequoyah Noding for Primary Loop with Model 51 SGs
. rem ron m m. .
i 1 i . i
= = l ea I m9 m 1 -9 - l 4 4
- I = 1 -I = '
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Figure 5.9-3 Small Break LOCA Study Hot Channel Power Profile 3.5 l 1 j- i i- : ) i ! l . 3.0 - - - ~ ~ ~ - - 4- d: - - - - - - - - - - 1
,e m.w
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w a
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i 1.0 - - t F-- = -
---t-- e--
3 i i i ! 0.5 0 2 4 6 8 10 12 ELEVATION, FT Figure 5.9-4 1.5-inch Pump Discharge Break Primary System Pressure 2400 I ! ' 1
! r 4
+
2000 ;- -- 4 --- - J --- + , ,
-t =
i i .! I 1600- ,---- i-,.- -- --;- - -- 6 : U l , 3
- ia : ;
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.g 1200- -- -- - - e - -- ~
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*-*=9 e.e gae -- .e. 4,+ie.-es.+.+.g., ._.e.,6++..q.,wy..se4 l
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1 ! .i t.: O y 0 500- 1000 1500 2000' 2500 3000 3500 . 4000 m .. TIME. S
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Figure 5.'9-5 1.5 inch Pump Discharge Break Leak FIOw Rate 1600 i -
, , , 1 3 -! I i )
? ( ! -i >
. -1400- h--- - -
-- h---- -
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0 500 1000 1500 2000 2500 3000 3500 4000 TIME. S Figure 5.9-6 1.5-inch Pump Discharge Break Pump Suction Loop Seal Levels i 20 LEG END 15- -
-+ . - - . . .". INTACT LOOP SEAL DOWN. )
1
, i BROKEN LOOP SEAL DOWN
- O- INTACT LOOP SEAL UP 10- ....-t.-......f..........._ _
------ BROKEN LOOP SEAL UP
. . = - ymmmmmmm.m mmmmmm. __ _ ____ a g ;
} ! ! , \ I i k 5- a- 1
--- 9 4 d .
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-20 0 0 500 1000 1500 2000 2500 3000 3500 4000 W TIME. S t
Figure 5.'9-7 1.5 inch Pump Discharge Break; Core Collapsed Level 16 ii i 1 ( i i i l i I 14- -~
~~~~~*------+-~~~+-----4~~--~}- } !
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i
+ . +- --..- + e- s -.
0 0 500 1000 1500 2000 2500 3000 3500 4000 TIME. S Figure 5.9 8 1.5 inch Pump Discharge Break Hot Rod Clad Temperature
=2400 t 4 i
2000- --
!- 4----- - - - -
---- -~~ - h ---j ---~~ - ^
! t
- i ~i ! ;
i- l i i g 1600-
----- a,- _ i,- 4 ; -
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500 1000 _ 1500 2000 2500 __3000 3500 4000 ___ - m _ TIME. S I
Figure 5.9-9 2.5-inch Pump Discharge Bre k Primary Syst:m Prcssura 2400 t 1 - i
)-
i
! i 2000 L-. ,.-.....-.l. L-- 1. - .-.. . .!...~..-
f - -- --- --. I ! i l l , i !
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i ! : 1600< --!- '
---? - .. 4 - [. -- - .. .,
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f I I 400- -
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r- - 4 - 3-
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O O 500 1000 1500 2000 2500 3000 3500 4000 TIME. S Figure 5.9-10 2.5-inch Pump Discharge Break Leak Flow Rate 1600
- i i $ I 1400- - -- r + ..-4 --- - - + - - - - -
! i : :
i i ; t i I I < i I i i 1200- -a 4- t- 4- L-- -f G sm . J jQQQ- 4..- . . uf ! - r e i i
! ! A f 800 --- 4- - 1. - + -- - - - -
S l l C3 i J i' 6 i. m 600- - -- l- --i -
--j--- i- 4-v3 $ i
*( :
-' i 2, {'
I i i O 400- --- --- --.-- .. + . -- l
- a '
'co.
i
.. ,1 h l t l l *[ @
b '9p - q t A.,
,200 .:. _4......
.. . n 7 4- , &
7 l ! ' o
?
l z a, O y 0 500 *1000 1900 2003 2500 3000 3500 4000 n. TIME. S
Figure 5.9-11 2,5-inch Pump Discharge Break Pump Suction Loop Seal Levels 20 - LEGEND
- j. i, -
- I 15- - -- a-- - -
o INTACT LOOP SEAL DOWN, 1
.! BROKEN LOOP SEAL DOWN -t i . j - 0--- INTACT LOOP SEAL UP 10- b----- - -- ----. 1 ------ BROKEN LOOP SEAL UP '
.caaaaaha. !
k 5- - - - -
.-+ - -~~r 7- = +-- -h-- - j - - - - - -
j i d ! I ! b 0 '^^^^^^^^4Tr~yRd ^^^ c
- -+ h-~~--
@ \ "p e l O 5- - - -
-M- t- " ' " ~ ~ ~
f
, 4 10 7 _ ,,
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.j $ < . . . . . . . .. . . - . . . . ~ , . . . . . . .m,...-.- .
i i * .
, ! ! i i 20 -
0 600 1000 1500 2000 2500 3000 3500 4000 TIME, SECONDS Figure 5.9-12 2.5-inch Pump Discharge Break Core Collapsed Level 16 , i 1
- ~~--
14- +- - - - .
- ~. e - - --- - ' - -- !
I i i 12 k ...
.F- 4 - ..,.;. _ v. { .-
- ; 5 i i : 1 i k 10- '- I A- N-- --*
i h "; ---- 8- --
- ) -
e o i i i I A-
$ 6- h y ,---* -i-i .D i * =-
l t 3 i 4 g 4- ~~~-- L -- - A -- 4 --- *C O i
. o w
j
%c 2- - - - - -
r -t -- O
, i g
. I w
- n. O O O 500 1000 1500 2000' 2500 3000 3500 4000 A TIME, S -
i
Figure 5.5.13 2.5 inch Pump Discharge Break Hot Rod Clad Temperature 2400 i , i i r s > 8 i. 1 i i i a ; i i 1 2QQQ. .-.~....l... ... 4 ..{....-... .4... , . . . . - ~ !.. j .m j i 4 ; 1 : . ! ! 1' . l 1 : i ! j QQ . 4, .. j... ., . .p
..__.-...h....... . ,4.... ...- .. ..j . 1
- u. i.
t N 3 l;
>= ,
1200- --
~~~ - --4---
t t
. -+ 4 *-
1 e----- 1 i ' I i i .
, ! I l
~
l 4 i i . gQQ. .............i... ..j .. , . ,, . . . .j. . ....j..... . . 4., . ... 4 l
> i !
i . i L > i i I i i e ! 400- -
'---+----e..--+ + 1
+ - --4 l ! i .
, i : 1' 0-0 500 1000 1500 2000 2500 3000 3500 4000 TIME. S Figure 5.9-14 2.75-inch Pump Discharge Break Primary System Pressure 2400 ;
1 l i i' i 2000 !- s -- - - t i . i i 2 . 1600 -~~ - L ----------i- - - - - - -
$ . i I { '
2 1 , 3 1200< -
--+-
i * -b, M i
~'
E l-l l i [ ! b 800< --- + -- e - --~ ~ !: +---- t--- ; e l i .it p e
> ; c.
! S 1 i- ! O i I. !
w 4QQ. ~ . . . ~ . . . . ' . . . . .. [ . 4 ........q..... ...j-- - ..- d ~.+ p.--- %C
! i O
i ! g
, t z O U 500 1500 2000 2500 3000 3500 4000
- O 1000 TIME. S t
W Figure 5.9-15 2.75 inch Pump Discharge Break Leak Flow Rate 1600 ' s 3 1 3 1 . 3 t , t 1400- --- -+- p
-+
1 6---.--
..L
- -~~-
4- .----. .b. . - 4 i l ! J l } i i . 3 1200< - - - " - --.i---- L- -- -- L -..J-...-~~ f i i = jogo. i d 4 f 4
--__.-_+_____.. '
-4m_ _ _9__. . ; _ .___ _4 .L.___ _
ul i ! > H ! I { i ! i 3
$QQ-
.-...,..m........!......b-i-
4 t .
.f .
4 l 1 o j i i , 4 a i 4 : g 6QQ. --~,.u. --- ,_ , - . - , n-~ sn ,-~.. 4._ . ~_m_ ~~.
-m < i i
1 2 , l ) 4on- - - - ~ ~ . ..- ,
. - . - . .~
. r- . - -
f goo. h ! . . i 4 f
..4,...._...4....._._..;i i
_ i I ; 1
; . W !
' 4 0 .-
0 500 1000 1500 2000 2500 3000 3500 4000 TIME. S Figure 5.9-16 2.75 inch Pump Discharge Break Pump Suction LOOP Seal Levels . 20 LEGEND 4 i 15 . - - - -
- + - - - - + - - - - .
/. INTACT LOOP SEAL DOWN-BROKEN LOOP SEAL DOWN G- INTACT LOOP SEAL UP 10: ---
- - r-I ------ BROKEN LOOP SEAL UP
." . - >_-= i , ,
i f [ 5- - i,- - - 1- - "-"-E T i T
- i i i !
g;j :
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^
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9
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, i o
w
. t c 15- l I--- -I""'"-~" ' " ~ ~
5
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. z
. a.
-20 -
3000 3500 4000 9 0 500 1000 1500 2000 2500 TIME S
- - . _ - - _..-.- -.. - .. _ -- - . ~ . - . . . - - - . . -
.- iF gure 5.9 2.75-inch Pump Discharge Break -
- Core Collapsed Level j'
- 16 i4 _ . _ , _ . . _ . _ _ . . _ , _ --. ._-..._.,.._. ..+._ , . . _ . . , _ _
-1 .
-i i 3
; i +
i i t . , 1 12- --- -- -----+--~~+-----t-----j--~~T 4- , ! (( -+-- - t----~~ r 1 -t-----+---- [ 4 8- ----------- --4----- b 9 i
-o , !
i O
-r 4
a 6< - - - - - - - + - - p t----* 1*- ;, 3 a l, + t ,i i; <. i }; 4< - e* 4mam+e. m.4 . . mm .n.4+.m --., ..e.-- gm .on .a i > i ! 1 !- .- j t i a g i 8 ! ! d i . 1 I ! 2 . .
.-.. . 4, .. . . . ,._ a 4, . . . F, .-__ _ ; ,
1 I r i j i {- i 8 j !
- 1 i i
+
, . ; l 1- +
0 0
- O 500 1000 1500 2000 2500 3000 3500 4000 4
, TIME, S i Figure 5.9-18 2.75-inch Pump Discharge Break
- Hot Rod Clad Temperature 2 400 -
i , ; 1
- - ! i
- i I ! i.
2000- ------! -- -L i
-t, , ,
i 1 l
- i 1
- ; 4 i i i 1600- 4 -4 - - - 4-
- u. 1 ! ;
; i ;
'4 . }a i i 1 i i, ; :
4 ' +
' I200- -- * - - - + --
- 4, 4, ~
i j -
- i
-800- --+-~~~~- ;-- - y k &
7 3- r -~~- i i
! ti
< ( , ,
?. .
i , ca,
! l !- !
~
,! 8 400 +- a . _a_ . .i-- ; i A so 1 ! ; i
.i
=
1 8 Z i, 1 t 4-w i 0 O
- j. 0 500 2500 3000 3500 4000 W
-1000- 1500 2000
( e, TIME, S 1 f 1
. - , - < g-- r, 9 , ,,-cy., , , , . - . . _ ...__.,#.. . . , . , , - ,- - - , , . - . . _ , , , - - , - ,
Figure 5.9-19 3-inch Pump Discharge Break Primary System Pressure 2400 - j.
. i
- - -' + - - - + - - - -
2000- --
- -+ - --
. i f
i 1600 a- - 4-- - - i- -
-t- - - --
us i i t A i : uI i
; i a:
o 1200-a
. -. .-- - --- r-
~
) !
800-t
, l 400-
+ + - -- -
+- -k-0 0 500 1000 1500 2000 2500 3000 3500 4000 l
l TIME, S i l Figure 5.9-20 3-inch Pump Discharge Break Leak Flow Rate 1600 1400- -- 4- + > - - + 4 1200- - 4 -+- i- 6 -- - -- E ! s 3 1000-> ui
. _ __;___ ; d.-. __ .
H I -$
, 800- --- ---- - ~ ~ - - - - + ---
O p > a 600- -- -+-- -- - - i - + - 2- -? - - - - i - D 2 ,
$o
^00- - - - -
-t--- -+ + P 4- -
t 4
}
2 i o. 6 m. 200- - - - -- --- ?--
, 8 W 7
, W 0 O O 500 1000 1500 2000 2500 3000 3500 4000 "
TIME, S i
~
Figure 5.9 21 3-inch Pump Discharge Break Pump Suction Loop Seal Leve s 20 - j 4 LEGEND 15< - 1. .". INTACT LOOP SEAL DOWN,
-- -i j ;
^
BROKEN LOOP SEAL DOWN 4 - 0-- INTACT LOOP SEAL UP jo. ......4 . .
------ BROKEN LOOP SEAL UP
........ i ,
s . b 5- --
+ - -- t - ;- - - -
--f--+-- I - ---
d : I
! ! i
'^^^^^^^ ^^^-
b 0 -- -- 00- - - l-5- - P - - - - - - -
-10 < -
+1 ' '
i i i !
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e) & . ---*+-~..$.. .. ,. . . . . ...~..g~ --- k... ...g. i i ! i { ' i i .
+ i I
~
20 0 500 1000 1500 2000 2500 3000 3500 4000 TIME, S Figure 5.9-22 SequOyah 3-inch Pump Discharge Break - Core Collapsed Level 16 - - - i ' 14- - -- r- - -- - - , -.- -- -- [ I' + , i , i 12- \ i
--+-- t- - - + - +- - -
4 i i !
! ; i !
? i 5 .
k 10< -- -- t- - I i-----+- d ! ! ! ( , 8< . o : ! ! ; i a 6< t t --- t t- --i . r- t-l t i l D c i i ! - i V 4< - - - - i . - i- 4. !.. - . c;
- ; t i- i
=
o. i, O
+ i ! i w a ; A 2- -4 -- ?- ?- ; ;..- .- .-
o i i ! ! Z 4 A 0 0 0_ 500 1000- 1500 2s 70' 2500 3000 3500 4000- A TIME, S t n . .
. . . . . . . __________J
Figure 5,9-23i 3-inch Pump Discharge Break - Hot Rod Clad Temperature 2400 . 1 p -- l, !. 3- -! i i i- i
'2000- - - - - - -
+l -
- . - - - -;L -- -E-i
-4 ~~- - J -- - - -
- 1
!- l' t -f I
l- - u. 1600< --~--j- +--- y - -i - -& ,
. } --- i
.-- _.. 4 - __._.. _ _ _ . _ _ _ _ _ , _ _ . _ _---__+.m_m-__ _ _ _ _ ,
I 11200<:800-f I
--r--~~
i i
- - ~ -
t i i ! ( i 1 l' I n ,- . 4QQ- . d m... -e. d *. m4. m= -
-y....w,....
i-0 0 500 1000 1500 2000 2500 3000 3500 4000 TIME, S Figure 5,9 24 3.25 inch Pump Discharge Break Primary System Pressure 2400 - , ;
>, j i 4 I i ; i i i i I I 1 .
1- : -1 1 [ 1 -r i 2000- - - - - -
- -r -
j -- - - -t. - -- t----- i : t I 3 l ! l i ! ! 1600- - jc -
--+ t; i
=t. 7---
; i i l
' .1200 -.___.4 __4__.___,
1
. _ _ _ _ _ 4- _ .. -
7e m- i , i I l
.(
-800- +--- d '
= g --
----~~~4- ~~
{-
.zc-
, c1.
t O
. i -w 400< - y- + - - - - - + ,
t
- 4-i 4-
;- r sL..
i -: [ c i 1 - o I ! Z
- i .; i i l w 0 U 0 500 1000 1500- 2000 2500 3000 3500 4000 W TIME. S !
1
Figure 5.9-25 3.25 inch Pump Discharge Break Leak Flow Rate - 1600 i i i ! 4 i ; 1 < i i i ! 1400< -+---~~i----+------- - L-~~ ---+i- - - - - t l 1 ! 1200< - ~ ~~- 4 - - - - - - - - - - - 4- r-- --+--~ ~ I
$ 1000 +.--- -__4- _ - _ ;-- __ __d_ -+- -- , _ _J j -
I k 800< --
-----+-- -+- -i -
+
3 I' - i i 9 w _- .
- i i m 600' +
- t -- F --*
T- *-- r i l Q i ! I s 400-
%---+-- - --
i s
- -+ 4 i
r-- . i 200- - -- . . - - - +- b- . C _i i s ! l ! ,
'N I O
O 500 1000 1500 2000 2500 3000 3500 4000 TIME, S Figure 5.9-26 3.25-inch Pump Discharge Break Pump Suction Loop Seal Levels 20
, LEGEND 15- --
+ - - -, m...., t. INTACT LOOP SEAL DOWN,
( i BROKEN LOOP SEAL DOWN
- O- INTACT LOOP SEAL UP 10 --
i-_ BROKEN LOOP SEAL UP i i ., i j t 5 - 1 -. J ..- 4 7_ ._ _. 7 d i i ! g I qq- j-a 0 ;O0000 -- -d [--- j h.5 }#b' .._j . i ; , o 10< - --~~~a - m' i j
*c o.
l, l 4
- C w
1 5
! i ? b
~
15- --- ! ;- t-- 4 o i Z i .I i i , m
-20 y
0 500 1000- 1500 2000- 2500 3000 3500 4000 A TIME. S t x -
Figure 5.9 27 3.25-inch Pump Discharge Break Core Collapsed Level 16 i e i k< . d. . . ~ . , . . - , . , - -
- e. ,
1 i . i 12- i - -. -q
- 4 - . - - - - -
-- - -- -- f --
! I h a C 10'
--( -h-
--+- i- -- k -- - - - --
d i h 8- --- -- h -- k, -F- -- $- - G i , o i i !
$ 6 ---- .-
i A - - - . - - .
-r - - - - - - - - -
4 4< ---+ -- - .- -e --- - - + - - - - - - - 1 l 2 - -
.. 4.. .. . - - . 4 .. , . . . . . - 4. . _ . . .
i 1 0 , 0 500 1000 1500 2000 2500 3000 3500 4000 TIME, 3 Figure 5.9-28 3.25-inch Pump Discharge Break Hot Rod Clad Temperature 2400 2000- - , +- , 1 1600- - -- r- . - - -
---r - - t '" -
' - ~ " ' * ~ ~ -~
uI i cc i . 1200 . - - . . . 2. .. . s. . - . ... k i e i ! p 800- i -
- -. 4 .- .
z ( :
'C a
i O 400- -- - E s C Z 0
. w y
0 500 1000 1500 2000 2500 3000 3500 4000 W TIME, S 1 1 J
4[' l, I E .
= Fig'ure 5,9-29 5-inch Pump Discharge Break-
[ P:Imary System Pressure 2400 l +
, 1, r I 't i l l 4'i J i i 1
] i -t i
, 2000< -------+-----4.>.--+.- -- --- i- -! -
.. 2 l
e I
> l
-1600- ---- ,- - - - - - v- -
--{ -+- 4--- --~~-
1 t .I f ,- . i i t 1200< -- 4-- -+---+---t'~~~~"~'~+'-*~'-A '
, i e i
t- [ , -pQQ< .- m.. .. . - - [...,...,4.,_ f ... - . ... ..p. _..., 4 i
} l r
j-l 1 4' - 400< -p- -
----,---9.-- b - - +- ,
-F i- i i
. j
- r. ! ,
4- ; , 0 l 1 0 500 ,1000 1500 2000- 2500 3000- . 3500 4000 , TIME S
. Figure 5.9-30 5-inch Pump Discharge Break Leak Flow Rate -
i 1600
, !, I, -
i i' d i !
-j(oQ.
; ... m y m. .m. ... .. 1 _ ... , f... . _
J. _ . . . . . . . . . . ...-.4,,-..,..,.,. i 4 ; t t l -
.i 1 1 L ;-
1200-
-) ;-
- - ~ ~ , - --:----+. e- - --y e
s 5 jooo. . . _ - .. . r--
- - - - - - +- - - ' - * - * - - - '
f g-j j i -l
' 2 i ,
---+-----+
i
--+ ~&------+ + - 3: 800- : ; 3-
- o. I ;
; it ! 1 ! .
; y 800< .- -
- - - -- -. r b ;---i- :j :3 E
~
, }a 1 M
=
I 'l l ' 400 - ..: .. 7 __.4 _ _ _ - j,.-- ..p , -g,. 1
! o w
i . i. k-
. b.
2 -m**== E w M-- - e**w we eau..s++ - o' E =
-p==+m w eem ,m w..a me.um 4 -+meese o ee... 4. -
.v.
=
i i ; 6 ; i-
!- ! Z i b . j i i, !
w j' O 0
~
U 4
, 500 1000 1500 2000' 2500 3000 3500 4000 4
TIME. S a d h i c,--. . ..wc
, . ,.v,- - - - -
Figure-5,9 31 5-inch Pump Discharge Break Pump Suction Loop Seal Levels 20 LEGEND l 15- -----+----+--~~-4--.-- - - INTACT LOOP SEAL DOWN,
-! BROKEN LOOP SEAL DOWN 4 - 0-- INTACT LOOP SEAL UP 10- - - - - - - + ! ~~ ~ ~~ *' BROKEN LOOP SEAL UP l 2 i
- j i l .
t- -5 -__._ _ ! _ _ _ _ _ . _ _ w . - . _ _ _ ! d
~- a y ;-~ <
- . - - . s l 5-. 0 4 .--- . 7._ .
8 h I g 5 - -. _.7 ,. _ f._ ,._ r - - _ 4 -\-- . Ia i t 30 .._.I A m. _4.. i_ l '
.... - 1._. . _. !. ._ .. .. ..._
I bleh I ! i i l I } 5 l
*15 ' -~ ~~~ l -+, - 4, --, - - . }. -4 !- .- ,
I I, i
~
-.20 O 500 1000 1500 2000 2500 3000 3500 4000 TIME. S Figure 5,9-32 5-inch Puma Discharge Break Core Collapsec Level 16
- i .
14 , ..._ - i 4....... .4 .. -p._ ...y . _ . . . i - i : .
! l 3 1 i i i i . i 12- -
Y- f --
-+ i-I I' I i i ! ! !
(E 10' - d - -- .4 --+ d-~ d-- I-h 8- -
----+--- -~-f +.-----4 ~~ -j - - + - - - - - - . - - . .
8-3 6< -- -- - - j, -- { s . s g 4 --_ . .- _- _ a ._ __[ _2. g t o w i l i i- i : 4 2 -- - - ;. - ?- _ 1... .
! _....), d.
a r ,
) Z
. i
~ u.
O g 0 500 1000 1500. .2000' 2500 3000- 3500 4000- A TIME, S i
E Figure 5.'9 33 5 inch Pump Discharge Break Hot R0d Clad Temperatuiu 2400 , , . , I
! 6
, i i f ;
4 1 j i '
! j t 1 1 i ! '
l 2000- - + ' L :L --+i i- L --
! j i :
! i I i ~l i;
, . ,i ;
S 1000- ----!--*- -] - -k- - f -- I - - - i, I i. ! ! ! i l, ., ,i I
! l, i, i i i.
1200- ' -- - li - i J
-+
i 4 -4 . l { - ! ! i i I, i . !. l . t i
' i !
} } ! l SQQ, . 4. - . .. - . 4 .. 4..._ . .... - - ...o ,
. 4, .
- . 4, . . . . . .
i ! I i i i ,i 00- i-4 f- -} k- - - h- ' - 1 i- - 3 1 . i
! ! t I , 4 I, ,
I , 4 , ' i O , 0 500 1000 1500 2000 2500 3000 3500 4000 TIMti. S Figure 5.9 34 CCI Line Break Primary System Pressure 2400
)
j i 2000- - J- -- + 4- -- - . , , , 1 i k . i i i 1 1 1600- - - -
+- - 4- -- -
t-- - - - - , - - l 4 i j i i 1 k l I d 1200- -
, m-.- d -
- -- - I.- . - -- 4 --
l 800- ------!------= t-- -- L ~ -- -- - - D I l j -l 4 0 u
. ; g ,
i
. i o
400- -- y-- - ;. - .. . t - 4 y 4
; e o
1 Z 0
! . t A y
0 100u 2000 3006 4000 5000 6000 W TIME, S a
- Figure 5.9 35 CCI Line Break Leak Flow Rate j ~
1600 ' I i ! i ) } 1400< -------+-------!------d-~~---I-----~~~------ 1 l . - . - -. - - - - . 1200< - - - . - - . - - - - . - - . . , - - . . - . - 4 1 j 4 1 4
- 1000< --
--+------~~t------+------
N g , e00 - - . - . - - + - - . - - . - . - . - - . - . . . . _ . - - . . . -- - . - . . . . - . . . - - . - - - . - 600- ~ ~ ~ ~ - - -- ~ ~ ~ ~ - - - -- - - - - - - - -- ~ ~ - ~ ~ - - -- - - - - - + - - - - - -. 4
.400< ~~~~~ - n-- --- - -~~~~~ ~ ~ ~ ~ - - + - - :-~
4, : i.
-----~~~6-l 200- -- - - - - - - - - -----L----+-- - -- ---- -
W $ !- ! Y, i
- O
,- 0 1000 2000 3000 4(N0 5000- 6000 4 i TIME, S 1 . Figure 5.9 36 CCI Line Break
- - Pump Euction Loop Seal Levels
, 20 - LEGEND e 15- -
-- - - -r -- - A INTACT LOOP SEAL DOWN-
?
i BROKEN LOOP SEAL DOWN i 5- - G- - - INTACT LOOP SEAL UP i - -- , 10 ~~_---_-..-_._-7.._-__
+ BROKEN LOOP SEAL UP ,
.A 11 1-A aa?oaA - =tratP=trMP -
.N .
1, 'is b 5- - -
?
- +-
---;---- +-f-- -
1 ,' . . . I. .. . . . . . . . w w 'w. . w w w w w w w v. w w w 1- vl , 1 i l 5- ------g-- ~ ~ ~ ~ - - - - - - - j -~~~~-~~i---~~- p a ,
-t ;
- - . 3
.j o , -~~~.__,_--_-.uu_.-__~.-_----__.---.-.__-_-- - 1' o
st 15- - -- +- , - - - - - - - - -- d o
----~~-+-------1; '
i' l2; 1 !
- 20-6000 bW 0 .. 1000 2000 3000 4000 5000 TIME, S v
- f 1-e
1 Figure 5.9 37 CCI Une Break Core Collapsed Level 16 . f I ' i i 14< .- - ...,.... . . . , . . _ . . . . . . 4. . . ~ .......- ,, ,s .. . i i I ,
! i ;
jg< ....,.....l...
.......... ..4....., . ....)...
. . . . ..1.......,,..,_...,
! , . I I ,
'. I' 4
- . 1, ,
i , , 10- - l-
?- 3 -!- 4- . . ! - --
! t i i ,
e' i 1 I t f hi e. o ..n.,*.u6...o..oi, . f, . ..,.....p. ,.a... ..,u. e o.g... . . . . . . , , . , , , ,
# : i i i !
C3 l l l i 5 cy i ; i i i i i
- - t - --
4 a 6- .7 - 4
--- j- -
--- 4 - ---+
-----7--.---<
- i I
i I' ' y-- ~ ~~.' ~ 4- '. --+- - - * - - + , -
! ! I 1
f $ i 4 l 2< . ., ..; .. . - .. - .. p. .- .; ... ...'.. . _ . - . . . 4 , ,
; I I
i 0 , 0 2 4 6 8 10 12 14 16
- cC'^ - TIME, S Figure 5.9 38 CCI Line Break Hot Rod Clad Temperatuie 2400 r i 2000< - i --
-- +
l i t
- ,- , ,1 4- --+-- 4- -
- u. 1600 < ; i i :
Y .
- + - - - - - i.- - - - - 4 ' - 1-- - -- L-
; i i1200'
! l ;
i i ! i : i 800< - - --- -a {
! .I
-+----
I'
~-
g
, .m
' , ' 1, o e
A , 'c C. O w i i , i i 400< --
-- i- - - - - +- --- -- - -
A. i c: l i O 2
. Lt O U 0 1000 2000 3000 4000 5000 6000 W TIME, S i
I i i , S.12 References ; l j l i 5-1 BAW-10168P Revision 02, BWNT Loss-of-Coolant Accident l l Evaluation Model for Recirculating Steam Generator l ! 1 Plants, October 1992. j ! l J 5-2 BAW-10158P Revision 03, BWNT Loss-of-Coolant Accident l ! l Evaluation Model for Recirculating Steam Generator l : Plants, November 1993. l l l ? 5-3 BAW-10164P Revision 03, RELAPS/M002-B&W - An l Advanced Computer Program for Light Water Reactor ! i LOCA and-Non-LOCA Transient Analysis, October l- , 1992.- l l 5-4 BAW-10171P Revision 02, REFLOD3B - Model for Multinode l Core Reflooding Analysis, January 1989. j l 5-5 BAW-10166P Revision 04, BEACH - A Computer Program for l t
-Reflood Heat Transfer During LOCA, October 1992, l l
5-6 BAW-10162P, TACO 3 -' Fuel Pin Thermal Analysis ; Code, October 1989. l l 5-7 BAW-10092P, CRAFT 2 - FORTRAN Program for Digital 1 Simulation of a Multinode-Reactor Plant During Loss-of- ! Coolant, April 1997. l ' l 5-8 BAW-10174P Revision 1. Mark-BW Reload LOCA Analysis-for l i the Catawba and McGuire Units, September 1992. l l i 5-9 V. H. Ransom et al., RELAPS/M002 Code Manual, Volumes 1 ; and 2, NUREG/CR-4312, EGG-2396, 8/85. l l 5-10 BAW-10177P, Mark-BW Reload LOCA Analysis for the l Trojan Plant, October 1990, l l : 15-11 BAW-10184P, GDTACO - Urania Gadolinia Fuel Pin Thermal l l~ Analysis Code, February 1995. ! l : 5-12 BAW-10172P, Mark-BW Mechanical Design Report, July l f 1988, , ! g t C' 4
~
FCF Non Proprietry , 1 L s....._--...- - ~.~._._,_........_,_..--...,,_.--a
. . . , . . . - ~ . _ , . . . , . - , ,
The following supplies supplemental information tor the response to Question _1. ! Response ! t T h e amall break LOCA noding diagram submitted with the response to Question l'was incorrect in its depiction of the break and the ECCS injection for the broken cold leg. The attached figure replaces that diagram. The artificial break node is-shown as node 276 and the ECCS lines for the broken cold leg are shown to be entering node 276. f The.ECCS injection for-the broken cold leg is modeled in accordance with Section'4.3.2.2 of FTI(BWNT) small break j LOCA evaluation model, Volume II of BAW-10168P-A (Reference l 1). For the pump discharge break spectrum (PDB spectrum), i node 276, the artificial break node, is located below the _ RCSipiping with a downward orientation. This orientation assures that liquid resident in_276 will not flow into the- ; pump discharge piping. Thus, when emergency core coolant l
- (ECC) is injected into 276, it will flow'out of the break i and not be available to replenish the reactor vessel l inventory.
The emergency core cooling system (ECCS) for Sequoyah > comprises four injection systems: the high pressure, pumped ; centrifugal charging injection system (CCI); the intermediate pressure, pumped safety injection system (SI),
. the passive accumulator tanks; and the low pressure, pumped residual heat removal system (RHR) . The FTI evaluation l
model , Section 4.3.2.2, requires that the broken loop ECCS r flow be modeled as injecting into the artificial break node, node 276, after loop seal clearing. Since the artificial break volume is relativ$ly small, 14.6 ft' for Sequoyah, the a evaluation mocel alloss that some portion of the broken loop-ECCS flow can be placed directly in the RCS piping prior to
-loop seal clearing to prevent potential water hammer.in the :
- artificial break node.
For the Sequoyah analysis,-all of the ECC system flow
' injected into the broken loop piping is modeled as flowing into node.276 after loop seal clearing has occurred. Prior
.to loop seal: clearing, . the CCI flow is modeled as entering the pump discharge piping,_ node 270; the SI, accumulator,
-.-----v,-h- --
,,,,,vu-, ,.r, ,y,,,y-.,yy ., ,-b yy ., ,,e.,,,%m.,my ,-,7,r y, _w y w ,,y.. .n. , ,-,,w- , ,,m c , m y y ,y,--ry ,,,-y----.--vu-+ w,
Qw4r,nems,WWupa x r ? i ( and RHR are modeled as entering node 276 both before and after loop seal clearing. 14 ode 276 is modeled as a nonequilibrium control volume. The break flow path, labeled LEAK on the diagram, is modeled as homogeneous. The containment volume area is set equal to the volume area of the break node to eliminate momentum flux gradients during periods of non-critical flow. The flow area for the path connecting the pump discharge piping to the break node, node 275 to node 276, has an area of one third of the area of the RCS piping. All of these selections are in accordance with the approved FTI small break LOCA evaluation model. To assure that the problems mentioned in the response to Cuestion 8 were no longer occurring and that the break flow during saturated discharge was being controlled by the Moody Critical Flow Model, several hand calculations of the break flow were made for the spectrum cases and compared to the RELAPS/ MOD 2-B&W cutput. At least one check was made for each of the seven cases submitted. The transient time at which the checks were done was varied to examine different portions of the transients. In all cases, the break flow , rate calculated by RELAP5/ MOD 2-B&W matched the independent calculation of the Moody critical mass flux. This confirms that the code, under the homogeneous option for the leak path, conforms to the requirements of the evaluation model and to those of 10CFR50.46 and 10CFR50 Appendix K.
Reference:
- 1. BAW-10168P-A, Revision 3, "RSG LOCA BWNT Loss-of-Coolant Accident Evaluation Model for Recirculating Steam Generator Plants," Volume II - Small Breaks, Framatome Technologies Incorporated, Lynchburg, Virginia, December 1996.
FIGURE 5.9-2. RELAPS/2 SBLOCA LOCA' MODEL - Sequoyah Noding for Primary Loops with Model 51 SGs STEAtt GEndEf4ATost STEmes essEAAToft I =_l-1 = l-i - 1 .1- = Q d
.I = l l l
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= 1 mi = l --J l g =
pI = ._] 4
= l 1 1= l = l I= *= l 5 I
.i l m m i = * = .= =
m _. . Z 1 Z --- - ___ Z.Z[1 : . - - . . Z ZZ Z. ZZ Z _ T ___ __ __
~ ~
..{ T} }
{ Z
..' +1tt ' '
t Li 4
+l >= 11 = 1 1 .= ] l == l i,i 1
l L'Isr-l>> t-l' U- E m
=hM=
= HdH =1- ~ ==
8-W- mee_ l Y. r =^1 -
~
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./p
t REQHEST FOR ADDITIONAL INFORMATION TENNESSEE VALLEY AUTHORITY SEOUOYAll NUCLEAR PLANT. UNITS 1 AND 2 DOCKET NUMBERS 50-327 AND 50-328 4 Question 2: i The proposed TSs deviate from the approved Standard Technical Specifications (STSs) for Westinghouse plants (NUREG-1431, Revision 1) with regard to removing items from the TSs and relocating them to the Core Opetating Limits Report (COLR). Provide justification for this deviation from the Approved STSs for the quadrant power tilt ratio < (QPTR) and the f 3 (6I) and f (AI) inputs to the OTAT and OPAT reactor protection system, considering the Nuclear Steam Supply System (NSSS) and nuclear instrumentation remain Westinghouse designs. Additionally, justify the use of a QPTR limit of 1.03 which is less conservative than the formerly used 1.02. Revised Response to Question 2: The parameters relocated to the Core Operating Limits Report (COLR) associated with f ( AI) and f:( AI) are the positive 3 l and negative breakpoints for the AI deadband (the AI limits) and the slopes for reduction of the OTAT and OPAT trip , setpoints for each percent that the magnitude of AI exceeds the breakpoints. These parameters are determined on a < cycle-specific basis, and their allowable magnitudes vary based on the specific fuel cycle design, as determined by ' the FCF cycle-specific three-dimensional core power distribution analysis. The methodology for determining the f3 ( AI) and f ( A'I) limits is described in the approved topical report BAW-10163P-A. Therefore, approved methodology for calculation of these parameters on a cycle-specific basis is in place. Thus even though the NSSS and nuclear instrumentation remain Westinghouse designs, the cycle specific variation of the f ( AI) limits and use of NRC-approved methodology justify relocation of these parameters to the COLR. The equations for the OTAT and OPAT trips will remain in
-section 2.2 of the Technical Specifications. Although the l f(AI) parameters are incorporated in section 2.0 of the Technical Specifications, they are not safety limits, buc l are used in the trip equations to protect against violating j safety limits.
l The purpose of the f (3 AI) and f ( AI) limits is to adjust the OTAT and OPAT trip setpoints with variations in the axial . power distri' ution, as measured by axial power imbalance (AI). The reason that the f (oI) and f ( AI) breakpoints and t slopes are cycle-specific is that they are sensitive to the axial power distribution. The axial power distribution can i l FCF Non-Proprietary l
l vary significantly from one reload core design to the next, and the limiting values of 4I veroua thermal power level (or AT) are dependent on the peaking factors in each cycle-specific core design. Since a reload core-specific power distribution analysis is performed for each reload core design, the f ( AI) and f:(oI) limit breakpoints and slopes i are determined on a cycle-specific basis. By defining cycle-specific values of the f (AI) t and f,( AI) breakpoints and slopes, use of overly conservative values of these parameters that would unduly restrict plant operation or possibly cause an unnecessary reactor trip is avoided. In addition, if a core with more restrictive (higher) power peaking is designed, then more restrictive f (aI) limits could be necessitated by the reload safety evaluation analyses. It is beneficial to specify these limits in the i Core Operating Limits Report so that the cycle-specific variation.in the limits can be accommodated without the necessity of amendments to the Technical Specifications. The determination and function of the f ( AI) limits is somewhat analogous to determination of the AFD operating limits (which is a COLR parameter), in that variation of the axial power distribution is limited on a cycle-specific basis by the AFD limits to ensure that the accident initial condition peaking limits are preserved. Similarly, variation of the axial power distribution is limited on a , cycle-specific basis by the f(oI) limite to ensure that the OTAT and OPAT trip equations preserve the core protective limits. The cycle-specific power distribution analysis performed for the reload safety evaluation of Sequoyah Nuclear Plant Unit '. 1 cycle 9 produced the following flaI) limit breakpoints and slopes, as compared to the limit valuet for-cycle 8: Comparison of f ( AI) and f,( AI) Limits for SQNP Unit 1 Cycles 8 3 and 9 l Parameter Cycle 9 Values Cycle 8 Values l l f3 (AI) breakpoints -23%, +5% -29%, +5% l f3 (AI) slopes 2.50%, 1.20% 1.50%, 0.86% l l f:( AI) breakpoints -35%, +28% N/A since f ( AI) =0 l f ( AI) slopes 1.50%, 1.70% N/A since f ( AI) =0 l ! The comparison in the table shows that the limits for the cycle 9 design are more restrictive than those for cycle 8. In chio comparison, some of the difference would be expected to be due to the methodology difference and some is attributable to the difference in core design and peaking i FCF Non Proprietary
. factors resulting irem the core design. However, the comparison in the table illustrates that the f(AI) breakpoints and slopes can vary significantly from one core desi3n to the next.
There is precedent for relocation of the f( AI) limits to the Core Operating Limits Report. In an SER dated May 31, 1994, which responded to a request from Duke Power company for
- McGuire Nuclear Station Units 1 and 2 (Docket Nos. 50-369 t l and 50-370), the NRC approved the relocation of the f(AI) limits (breakpoints and slopes) to the Core Operating Limits ,
l Report. The SER stated that "There have been recent :
- instances where one or more of these parameters have been j changed in TS revisions and, therefore, they may be 4
considered cycle-dependent......Thus, the relocation of these parameters to the COLR is consistent with the ' e provisions of Generic Letter 88-16". With this SER, the NRC ! issued Amendment No. 143 and Amendment No. 125 to the McGuire Nuclear Station Facility Operating Licenses NPF-9 l and NPF-17, respectively. These amendments revised i Technical Specification Table 2.2.1, its associated Bases, and Technical Specification 6.'.1.9 ) to reflect relocation of , the f(AI) limits to the-Core Operating Limits Report. 3 Topical report BAW-101630-A was approved in 1989, but was written in 1986--1987 before Generic Letter 88-16 was
- published and use of COLRs began. The topical report also pre-dates NUREG-1431, Revision 1. The purpose of the sample Technical Specifications provided in Appendix A of BAW-
- 10163P-A is primarily to illustrate which parameters must be limited and. monitored in order to comply with FCF's power distribution methodology as presented in the topical report.
The SER described above illustrates that relocation of the
! f(AI) limits to the COLR was considered and approved by the i' NRC subsequent to the approval of the topical report on the basis that they are cycle + specific limits, and hence comply with-the criteria specified in Generic Letter 88-16.
,l - ! With regard to the question on quadrant power tilt ratio L (QPTR), a limit value of 1.02 will be used and QPTR will not F be relocated to the COLR, as agreed at a meeting with the NRC, TVA, and Framatome Cogema Fuels on February 27, 1997. Technical Specification 3/4.2.4 (Quadrant Power Tilt Ratio) will be consistent with the example Technical Specifications in Appendix A of BAW-10163P-A. , ) FCF Non Proprietary v
.,,_-- -. ,.,,n,--__,,--,.. , . , - ..n,. _ , . . , , . ---_._.._,c .
-n_ .,,wn,.__n.,., ~ , . , . - ._n. . _ _ , . , . , - . , , , , ,-
- 3. provide additional basis for not including uncertainties if i the F"a(XY) and F"g(XY) in the footnotes of TS SR 4.2.2.2 and 4.2.3.2.
- Response SR 4.2.2.2 and 4.2.3.2 specify the precalculated limit quantities that will be used for comparison of measured peaking factors.
l The pre-calculated limit values are determined in accordance with the methodology described in topical report BAW-10163p-A. Section 6 of the topical illustrates how the allowable limit values include consideration of the applicable calculational and measurement uncertainties. Therefore, since the allowable limits l- are reduced to accommodate uncertainties, multiplication of the measured peaking factors by the uncertainties is not necessary prior to making the comparison to the limit during. peaking factor 1 surveillance. The peaking factor surveillance process is automated.. A plant and cycle-specific data base of limit values is provided for the
! core monitoring software. The limit, values provided in the data base include the applicable uncertainties so that external application of uncertainties is not nee'ded. .
4 i ? i i u i 4 . i 4 FCF Non-Propri6tary
- 4. Why is F"a (XY) not reduced by 2% over what is specifie d i n the COLR as the approved BAW-10163 prescribes in TS 4.2.2.2.C. 4.e.17
Response
() Surveillance requirement 4. 2. 2. 2. c 4. e .1 of the sample technical specifications provided in Appendix A of BAW-10163P-A accomplished this by requiring that the measured peak be increased by 2% (i.e. a reduction in measured margin) when two measurements extrapolated to 31 EFPD beyond the most recent measurement yield Ff(x, Y, Z) to be greater than the expected value (BQNOM (X, Y, Z) ] . For Sequoyah, the constant ( 2 tr) factor is replaced by a cycle-specific parameter and refocated to the COLR. () Therefore, the surveillance for'the Sequoyah technical specifications was revised from that in BAW-10163P-A to state the
" appropriate #=ctor specified in the COLR" instead of "2%."
4 t e FCF Non-Proprietary
- 5. Explain why Fa and Fu are not verified each time the excore QPTR is verifled with the incere detectors as the approved BAW-10163 methodology prescribes in TS SR 4.2.3.3 and 4.2.2.2.c.4.e.
Response
The methodology provided in BAW-10163P-A was developed to define limits that preserve fuel design criteria related to power peaking and reactivity for Westinghouse PRRs utilizing Mark-BW fuel. The sample technical specifications provided in Appendix A of BAW-10163P-A were created by incorporating these limits into the framework of the Westinghouse Standard Technical Specifications. .. The Westinghouse Standard Technical Specifications included the requirement to verify F, and Fu whenever the quadrant power tilt ratio (QPTR) indicated by the excore detectors is calibrated. This requirement is not part of the limits required to preserve the fuel design criteria as presented in BAW-10163P-A. It exists in the sample technical specifications of BAW-10163P-A because it was originally present in the Westinghouse Standard Technical Specifications. However, the current Sequoyah Nuclear Plant Technical Specifications for power peaking factors (Specifications 3/4.2.2 and 3/4.2.3) do not contain the requirement to verify the peaking
- factors whenever the QPTR indicated by the excore detectors is calibrated. This requirement was not included in the technical specification change package for FCF fuel because its introduction is not specifically related to the fuel design change and would have imposed a separate requirement that does not currently exist for the sequoyah units.
FCF Non-Proprietary
~ '
- 6. Justify the assumption that 15% of the steam generator tubes have been plugged. What effect will that have on the LOCA analysis results if more or less tubes are plugged in the broken or unbroken loops? (see p. 5-6) Additionally, the non-LOCA analysis assumes that 20% of the tubes are plugged.
Justify why this is limiting for all transients analyzed.
Response
Transients were analyzed for a maximum tube plugging of 15% in all generators. The 20% tube plugging for non-LOCA transients on page 6-4 is incorrect. The correct value is'15%. In general, 15% steam generator tube plugging is used as a desig". limit for the reload transients. The use of an upper-bound tuce plugging limit is conservative in the majority of cases because of its adverse effects on core cooling and primary heat removal. Core cooling is adversely affected by the lower RCS flows associated with increasing tube plugging. The margin to DNB is reduced. Primary heat removal interruption is exaggerated by using an upper-bound plugging limit.~ Secondary pressure is also reduced by maximizing tube plugging. Following turbine trip, the secondary pressure increases from its lower value to the steam line safety valve setpoint. The primary heat sink temperature changes relative to the steam pressure increase. The pressure t.ransient and, therefore, the sink temperature band are enlarged by maximized tube plugging and the effect of turbine trip on primary heat removal is conservatively simulated. For large break LOCA, the primary impact of steam generator tube plugging on PCT occurs during the core reflood phase of the transient. The loop pressure drop is a function of steam flow and tube flow crea. The core flooding rate decreases as the tube plugging increa'ses due to increased resistance to steam flow through the steam. generators, the steam binding effect. Therefore, assuming the maximum plugging in all generators produces the minimum core flooding rate. During reflood en additional 5% plugging in all generators is included in the calculation to account for seismically-induced tube collapse. The steam line break analysis is an exception to the use of an upper-bound steam generator tube plugging limit. For the steam line-break transient, tube plugging is conservatively neglected to maximize heat transfer across the steam generator tubes. Both flow and heat transfer surfaces are adjusted commensurate with the assumption of ot steam generator tube plugging. Increasing the RCS cooldown in this manner results in a maximum positive reactivity insertion and return ,to power; thereby, maximizing the severity of the event. . FCF Non-Proprietary
. . , - - - -o- g-
7: The loss-of-coolant accident (LoCA) analysis assumes that the reactor coolant system (RCS) flow is 348,000 gpm; however, the TS Figure 3.2-1 allows flow down to 342,000 gpm ' l if power is derated. Show that the doration is sufficient to assure that no limits are exceeded. Additionally, the TS minimum RCS flow is being reduced with this submittal and ' the analysis on the new Framatome fuel is performed using the lower flowrate. However, TVA is relying on the current Westinghouse analysis to show no limits are exceeded for the Westinghouse fuel inserts. Justify the use of'the current , analysis when the TS minimum flowrate is going to be lower . than was assumed in this analysis. ; a l Additionally, in section 7.3.2, what is the basis for the equr' b>n reducing Fa with reactor power?
Response
The loss-of-coolant accident (LOCA) was analyzed using the minimum thermal design flow rate of 348,000 gpm identified in Section 5.2 of BAW-10220. TS Figure 3.2-1 provider reduced power operation flexibility in the event that the measu.ed flow rate,
- less flow measurement uncertainty, is less than the minimum thermal design flow rate of 348,000 gpm. The power to flow trade off allowance is based on maintaining the margin to the DNB safety limit at the reduced flow condition at a level that is greater than or equal to the margin at the full flow condition.
The reduced flow condition is advantageous with respect to LOCA analysis results. .The effective core peaking limit increases in
; direct proportion to reduced power. A LOCA " hot pin" >
initialization case at 95% flow and equivalent peaking was conducted and concluded that there was no effect of fuel pin temperature on flow reduction in this range of flow. The hot pin initial temperature in lowered flow and power LOCA transient i analyses would, therefore, be equivalent to that currently 1 4 analyzed. During LOCA, it is the dynamics of the average core and not the hot pin that dictaten the progression of the transient. Initial core average fluid temperature would effectively be lowered, reflecting the imbalance between RCS flow and power reduction. Experience indicates that, all other things being equal, initially lower coolant temperatures produce a more subdued RCS core coolant flow transient during blowdown and i slightly hotter pin temperatures at the end of blowdown. However, during reflood, the lower core power level would more than make up for any blowdown penalty. Lower core power results in lower fuel dechy heat and the core would be quenched more rapidly. The net effect of a 2-for-1 tra'de off in power relative to flow would be lower large break LOCA-peak clad temperatures. The flou/ power trade-off would also be beneficial in the progression of a small break LOCA. Small breaks are longer term transients and initial core fluid temperaturer have little effect on their progression. With lower powers and lower decay heat FCF Non-Proprietary 2
,. mn.- - , , - -----m ---,,w.-,, -
rates, lower depths of core uncovery would be predicted for small breaks and the psak clad temperature would be effectively reduced. Therefore, the power to flow trade-off defined by TS Figure 3.2-1, would result in lower peak clad temperature predictions for , LOCA. The present Weatinghouse fuel LOCA analysis was reviewed by l Westinghouse with respect to a reduction in reactor coolant , l system (RCS) thermal design flow from 362,000 gpm to 348,000 gpm. l The Westinghouse review concluded that the reduction in thermal design flow is inconsequential since the core flow is immediately i dominated by the influence of the break during the blowdown phase a ^ of the LB LOCA transient. The influence of the break is such that the core experiences a full flow reversal very early in the transient, with lower core fluid exiting the break via the downcomer path to the cold leg break location. During the LB LOCA, the RCS is almost completely voided during the blowdown phase of the transient. The reduction in thermal design flow res'alts in an RCS Thot increase of approximately 1*F and Tcold redaction of approximately l'F. At the lower Teold condition, the break flow will be slightly greater initially due to-the slightly higher liquid density in the cold leg. This effect will be offset by decreased break flow when the higher Thot fluid reaches the break. During the entire blowdtwn phase of the transient, RCS Tave dominates the blowdown out the break. Minor changes to Thot and Teold at the same Tave have a negligible
'effect on the transient results. Initial RCS flow and temperature distribution have no direct effect on the subsequent refill and reflood phases of the transient. These phases are influenced primarily by the performance of the emergency core cooling system (accumulators and safety inj ection) . Given the small changes in initial RCS flow and temperature distributions, 'Weatinghouse concluded that the thermal design flow reduction will not result in any changes to the reported LB LOCA peak clad temperatures for either the Westinghouse standard or Vantage 5H fuel assemblies.
The power to flow trade off allowance shown in TS Figure 3.2-1 is based on maintaining the margin to the DNB safety limit at the reduced flow condition at a level greater than or equal to the margin at the full flow condition. The sufficiency of this relationship has been demonstrated by analyses performed for the Sequoyah Plant. . The relationship .shown in TS Figure 3.2,-1! requires that power be derated by 2% for each it that flow is below 348,000 gpm. The first test of this relationship was performed at two statepoints; one representing the nominal operating point (nominal conditions with a symmetric axial flux shape), the other representing the transient initialization point (nominal conditions with an outlet skewed axial flux shape). For this'part of the analysis power was reduced, the design peak was scaled up by the (1 + 0. 3 (1 - FCF Non Proprietary
P)] relationship (where P is the fraction of rated thermal , power), and the flow rate that provided DNBR equivalence to the full power, full flow condition was determined. Results showed that the 2-for-1 relationship is extremely conservative relative , to the true power to flow relationship required to maintain DNBR equivalence (by using the 2-for-1 relationship the DNBR margin is increased as power and flow are reduced). l The second test of the relationship verified its conservatism
- across the broad range of axial flux shapes permitted by the l Maximum Allowable Peaking (MAP) limits generated using full power l and full flow conditions. For this part of the analysis, reduced power, reduced flow statepoints were established, and DNBR's were ,
l calculated using the radial and axial peaking combinations ' ! permitted by the MAP limits. In each case, the MAP limits were L scaled up to reflect the reduced power condition (again using the (1 + 0.3(1 - P)] relationship). Results showed that at the reduced power, reduced flow statepoints, the c. culated DNBRs are ! greater than the original target DNBR's, confiraang an increase in DNBR margin. The equation FL m.t4 = FLomn.1 (1 + 0. 3 (1 - P)] presented in l Section 7.3.2 of BAW-10220 allows an increase in the allowable radial peak at core power levels less than 100% power. This ! relationship is also presented in the Sequoyah Tech Spec Bases ! Section 2.1.1 and in Section 2.6 of the Sequoyah 1 Cycle 8 COLR V as FL = F"(u (1 + PFa (1 - P)) where P = THERMAL POWER , RATED THERMAL POWER F"$a = the FL limit at the RATED THERMAL POWER (RTP) specified i in the COLR, and PFa a the power factor multiplier. for FL specified in the COLR. The same relationship was submitted and approved as part of the thermal-hydraulic methoda report for the transition to Mark-BW fuel at the Trojan Plant (See Section 4.2 of BAW-10178P) . This relationship defines limiting heat flux conditions that are higher than those calculated for the range of all control rods fully withdrawn to the maximum allowable control rod insertion , assuming the axial power imbalance is within the limits of the f (Delta I)- functions of the overpower D' elta T and overtemperature Delta T trips. When the axial power distribution is not within the tolerance, the axial power imbalance effect on the Overpower and/or Overtemperature Delta T trips will reduce the setpoints to provide protection consistent with core safety limits. FCF Non Proprietary
. . - _~
s ] j Question 7 additional information on mixed core penalty j The LOCA. mixed core evaluation performed by FTC for Sequoyah is based on the following:
- 1. There exists a valid set of LOCA calculations for each-of the fuel assembly types that may be loaded into the Sequoyah core and each of these seta of calculations presumes that the particular fuel assembly
- type comprises the entire core. The peak cladding ,
j temperatures.from these calculations may be, in fact . probably are, indexed by the originator. Indexing has i j been widely employed by Westinghouse to correct for l 1 potential or real problems identified with calculations. Indexing does not invalidate a
- calculational set it merely adjusts the peak clalding
- tempurature result by the amount of the index.
- 2. The only significant difference between the fuel designs involved in the mixed core configuration that ,
4 would lead to a possible mixed core LOCA effect is the + 2 pressure drop or flow resistance of the assembly, ! i
- Sections 5.10 and 5.11 of BAW-10220 address these conditions. The text of Section 5.10 along with the comparisons offered in Table 5.10-1 clearly establish the
- pressure drop across the assemblies as the only design difference capable of mixed core interactions. However, the i i existence of appropriate, non-mixed core, LOCA calculations i for the Westinghouse Standard and vantage 5H assemblies was not clearly identified.
I Sequoyah first loaded Westinghouse Van: age 5H fuel ! assemblies in the Spring of 1990. At that reload, two sets of LOCA calculations existed; one for the Standard design and one for Vantage 5H. A' comparison between the calculations for the Standard and Vantage SH designs, for the same assembly' power and distribution, showed a deviation of approximately 100 F with the Vantage SH having the lower peak cladding temperature (PCT). Further, a small mixed core effect of less than a 20 F increase in PCT was assessed for the Vantage SH assemblies (Reference 1). A similar but negative effect would be appropriate for the Standard assemblies. TVA, with Westinghouse's concurrence, decided to apply the Standard assembly.LOCA calculations P.d the - resulting peaking limitations for the licensing of both the Vantage SH and the StandardD designs - (Re.ference 2) , This precluded the need for direct referral'to a mixed core effect-because the applied peak cladding temperature (PCT) remained at least 80 F conservative (100 F - 20 F) . This licensing approach remained in effect until 1995. In
-1995, TVA replaced the Standard assembly LOCA calculations
- +
_v w-m- 4w, n-**=- v v r ' -e e
-m- s v ee -- =- w -er - w v -- - * +w M e w-*
. - - - -. - _ . ~ _ _ - - - __ __ __
t F i with the Vantage 5H LOCA calculations with no mixed core ' provision. Because the core was, by this time, essentially j all Vantage SH. assemblies with only a few Standard assemblies present. the maintenance of the mixed core dPCT s and the use of the Standard assembly based LOCA calculations ' l no longer-made sense. LOCA coverage for the few Standard i j assemblies that would remain in the core was based on the 1 low energy potential for those assemblies. All Standard l ) assemblies at the Sequoyah plant had, at that time, i
; experienced at least two cycles of operation and no longer retained the ability to operate at or approach limiting l 1
power conditions. Thus, when loaded, there is no ! possibility that these assemblies will be limiting with j j respect to LOCA. To provide further assurance, the fuel : 1 management procedures contain a proscription based on the ! assembly average power in that no PCT inde.ing is required for Standard assemblies if the bundle average radial peak is > less than 1.2B for the cycle loading pattern. If a Standard [ 3 assembly should exceed that limit, a +100 F indexing would be applied. j j > For the 1997 reloau, Sequoyah is adding Mark-BW fuel l l assemblies supplied by Framatome Cogema Fuels (FCF) and j reducing the thermal design flow to 348,000 gpm. For this reload, three LOCA calculational sets are required, one for i j Westinghouse Standard fuel, one for Westinghouse vantage SH ; j fuel, and one for FCF's Mark-BW fuel. Calculations for the-
! Standard and the Vantage SH exist but were performed at the ,
older thermal design flow of 362,000 gpm. To apply these : lt calculations at the new thermal design flow, TVA requested j that Westinghouse determine an appropriate index (APCT) for tbt shift in thermal design flow. The evaluation conducted i by Westinghouse showed that there would be no change in i calculated PCT for the f.our percent reduction in system > l flow. Thus, the existing calculational sets for the Standard and the Vantage 5H designs apply directly to this reload. The calculational set for the Mark-BW fuel was originally done for this reload and used the reduced thermal design flow of 348,000 gpm. All of these calculations . presume a full core of the reference fuel design. The- '
, applicable sets are: for the Standard fuel, the I calculations for the Vantage SH fuel plus appropriate ;
j indexes-(In accordance with the limitation outlined-in the , i above paragraph, the cladding temperatures will be indexed j by 100-F if'the bundle average radial peak is 1.28 or ;
. greater.);. for the Vantage SH fuel, the Vantsge 5H 1 calculations plus the appropriate indexing; and for theThe
' . Mark-BW, the new calculational set performed by FCF.
- current calculated large break LOCA PCTs of record for ;
i Sequoyah are: 2.911 F for the Standard assembly and 1911 F for the Vantage SH assembly. These values were reported to The j
-the NRC in the' June 26, 1996 Annual 10CFR50.46 Report. l PCT for the Mark-BW assembly is 2115 F.
i. l x s - l
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.__._____.-.._m.-_.__.
l l
< 1 i
' The inixed oore of facts are obtained from an FCF evaluation that cond ders all three co-resident designs. This ). evaluation has shown that the mixed core configuration has
- - only aw mall effect on the results of LOCA calculations such
- that the full covet calculat. ions, with some indexing, are
{ sufd cient for the licensing of transition cores, a ! -hs a r_esult of thic reload, the mixtd configuration of the ] sequoyah co:e over the next several cycles will comprise 5 i i I l i 6 , i [ e' [ i ! 1
~
r 1 j- + i 1 j- , i i i . d a .
.e 44 1
Cycle Mark-BW Vantage SH Standard assemblies assemblies assemblies 1st a 1/3 of core a 2/3 of core small number of assemblies 2nd a 2/3 of core a 1/3 of core small number of as+amblies 3rd & a full core small number of small number of on assemblies assemblies l The evaluation of these configurations on the reference LOCA calculations was described in Section 5.10 of BAW-10220 but it may not have been clear in its applicability to all resident assembly designs. The evaluation of the pressure drop effect proceeds from consideration of the-hot channel simulations during the blowdown and refill phases of the LBLOCA calculations. Both the Westinghouse and Framatome evaluation models separate the calculation of the hydraulics of the hot and average channels during blowdown and compute both. During refill and reflood both evaluation models compute the fluid flow in the hot channel from the average core flooding rate. This essentially means that the mass flux in the hot channel for the reflooding phase is determined by an average core calculation. Only the thermal I effects within the h?t pins are determined by hot channel-specific calculations during reflooding. The expected mixed core LOCA effect can be determined by comparing the pressure drop within the average channel for a full core of the reference design to that of a mixed core. If the average core for the mixed configuratier. would develop a nigher pressure drop than that appropriate for a full core of the reference design, two things will happen: some flow will be diverted into the hot channel during blowdown, lowering the cladding tempora".ure slightly; and the refill rate (dependent on the more resistive average channel) will be slightly slower, raising the cladding temperature slightly. In this analysis, it makes no differcnce how the average chnnnel became more resistive, one alternate fuel design or multiple alternate fuel designs, it ir only important that as an aggregate it is more resistive. If the opposite is true, the average channel is less resistive et,an the full implementation of the reference fuel design, fAow will be divarted from the hot channel during blowdown and the cladding temperature increased slightly, but the reflooding rate will increase slightly which will decrease the temperature somewhat. In either case, a .* wing ir cladding temp.erature during blowdown will be compensated for by a change in the opposite direction during reflooding. The configuration for Sequoyah for each of the fuel assemblies will be as shown in the following table. 1 - s
~
. yc e g
r $ IIf the"not The Average The Affect on Average ,
- Channel Fuel Channel Is Made Channel Resistance relative e is: Up Of: to Reference Calculation , !
i Ist Westinghouse Mix of Mark-BW Average channel resistance Standard and Vantage SH, increases above reference
- Fuel some Standard calculations r
! Westinghouse Mix of Mark-BW Average channel resistance . Vantage SH and Vantage SH, decreases below reference [> l j Fuel some Standard calculations . Mark-BW Fuel Mix of Mark-BW Average channel resistance l i } and Vantage SH, increases above reference calculations 7 some Standard FCF has criculated the ef fect for a representative core in j which the average core resistivity was increased above the > j reference calculations when Mark-BW fuel assemblies replaced Westinghouse OFA fuel assemblies at McGuire and Catawba . (Reference 3). The pressure drop for the OFA assemblies was measured to be 1 psi higher than the Mark-BW assemblies. This is the same as the sequoyah configuration where the ' vantage 5H assemblies also have a pressure drop that is The approximately a 1 psi higher than that of the Mark-BW. difference in cladding and fuel pellet average temperatures, for the Mark-BW assembly in mixed core operation when compared to a full core loading of Mark-BW assemblies, was a decrease of 30 to 50 F at the end of blowdown. This was caused by the diversion of some liquid from the more resistive average channel to the hot assembly. As expected, the temperature rise during reflood increased due to the , slower flooding rate. The increase was 30 F. Because the peak cladding temperature occurs during reflood, the .; differences in temperature rises are additive and the net change was small decrease in PCT, less than 20 F. This decrease in temperature is appropriate for the Mark-BW assembly-during mixed-core operation with the. Vantage 5H design. For the Vantage 5H fuel assembly, the average channel of the- 7 mixed core will have a lower resistance (1/3 of the core will be comprised-of the 1. psi less resistant Mark-BW). Here, the opposite results would occur. Some flow would divert away from the hot assembly durirg blowdown and the blowdown temperature in the hot channel wvuld increase by 30 to 50 F. The temperature rise during reflood would then decrease by 30 degrees or so. The cumulative result would be-a potential increase in PCT of lessLthan 20 F for tbs ' ' l Vantage GH. fuel. t The Standard assembly would evaluate much like the Mark-BW. Because it has a lower resistance than the Mark-BW, it would experience an even larger drop in PCT. Additionally, these } ,
?
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assemblies will not be prevaletnt in the core, only a few may be used in each cycle, and all, having experienced 2 cycles of operation prior to this reload, have very -low enerryy potential. They can not become the LOCA limiting assemblies l in the Sequoyah core, i l For each of the fuel designs to be used in the Sequoyah core l the mixed core effects are presented in the following table. ;
- 6 k
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a r v
't om,,ew n s ~-- m 4- ~ r ,mm,rerr-
f Mined Cors Effecta per Assumed Limiting Assembly { Fuel Designs Affect on Average Potential , Hot Channel Change in Fuel Design in Average Channel Resistance Channel relative to Peak Reference Cladding Calculation Temperature Westinghouse Mix of Average channel PCT Standard Mark-BW and resistance decreases Vantage SH, increases by less Fuel than 20 F l some I l Standard Westinghouse Mix of Average channel PCT Vantage SH Mark-BW and resistance increascs Vantage SH, decreases by less Fuel than 20 F some Standard Mark-BW Fuel Mix o'f Average channel PCT Mark-BW and resistance decreases Vantage 5H, increases by less some than 20 F i Standard In accordance with the depletion of energy production capability for twice burned fuel, fuel assemblies in their third cycle of operation can not comprise the assemblies Therefore, no mixed within which the LOCA PCT will occur. core considerations will be applied to these fuel assemblies. The Mark-BW, which benefits in the mixed core configuration, will be licensed with its full core reference calculations. The Vantage SH assemblies, becaure of the l I potential for a slight increase in PCT, will have a +20 F PCT index applied to the reference full core calculation results for all assemblies that have not gone through 2 cycles of operation. The indexing of the Vantage SH assemblies will be removed for their third cycle of operation. With the APCT applied to the Vantage 5H fuel the fully indexed PCTs for each fuel type are: 1911 F for the Standard assembly, 1931 F for the Vantage 5H assembly, and 2115 F for the Mark-BW. In light of the discussion provided, two specific statements in BAW-10220 need correcting. In Section 5.10. page 5-100, the statement is made, " Additionally, the relative average power limitation assessed by Westinghouse,for side-by-side operation of Vantage SH and Standard fuel assemblies is unaffected by this evaluation and continues to be applicable to those fuel types." This statementhdoes not reflect the f act that the Standard design assembli'es are all in their third cycle of operation and can not comprise the assembly of LOCA PCT. It is not correct and no side-by-side limitation will be applied to these assemblies. Further, the vantage SH assemblies will have a +20 F indexing applied
.. e ,c. .
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. .' v n a
- _ - - . - . _ . - - . . . _ - . - ~ . _ - . _ . . - - - - - - .- -
1 l !_ but that is caused by their side by side operation with the
- Mark-BW assemblies. The sentence should be removed from the j topical. j A similar statement, in Section 5,11, _ page 5-102, is; "Moreover, PCT penalties assessed by Westinghouse for the !
mixed core operation of Vantage SH and Standard fuels, 'I remain applicable to those fuel types." This statement is ; incorrect. The only 6 PCT necessary will be applied to the i' vantage 5H LOCA calculations. The sentence should be disregarded in reading the topical. . i
References:
1
- 1. Addendum 2-A of WCAP-10444-P-A or Addendum 2 of WACP. ;
10445-NP- A, S.L. Davidson, Vantage SH Fuel Assembly, Westinghouse Electric Corporation, April, 1988. Note, the , main. body of this report is appropriate for the Vantage 5 i design and may not be directly' applicable to Sequoyah. Addendum 2, however, is specifically for the Vantage 5H design.
- 2. B.W. Gergos and L.V. Tomasic, THFL-89-762[" Plant Safety Evaluation for Sequoyah Nuclear Plant Units 1 and 2 Vantage SH Fuel Upgrade, December 1989, Westinghouse Electric Corp.
- 3. BAW-10174A revision 1, Mark-BW Reload LOCA Analysis for Catawba and McGuire,. Babcock & Wilcox/Frnmatome Technologies Inc., November 1990.
p p
I l i
- 8. Please describe the changes made to the approved Babcock &
Wilcox Nuclear Technologies (BWNT) recirculating steam generator (RSG) evaluation model (described in p. 5-79) in greater detail and discuss any implications on the prior staff review and approval. The following response completely replaces the response provided ; in the submittal of additional information provided to the NRC on ' 2/7/97,
Response
FCF's statement regarding an evaluation model change on page 5-79 of topical report BAW-10220 is mis 3sading and incorrect. The use of equilibrium conditions in the fictitious leak volume (CV 276) is not an evaluation model change. We apologize for failing to correct that same misstatement in the 2/7/97 response to Question 1. FCF's NRC-approved LOCA evaluation model, BAW-10168P-A, Revision 3, calls for the use of equilibrium core control volumes and nonequilibrium control volumes in the remainder of the reactor coolant system. The leak volume is an artificial, nonphysical node, attached to the reactor coolant system. It is discussed in Section 4.3.2.2 of Reference 1. The purpose of the fictitious volume is to assure the bypass of broken loop ECC water to the containment, conservatively limiting the flow of ECC to the reactor vessel. The volume is neither classified as nor restricted to equilibrium or nonequilibrium. Also, the SER and TER do not impose any limitations on the control volume. , Calculations supporting the licensing of the small break LOCA evaluation model topical report have been performed using both equilibrium and nonequilibrium leak volume configurations. The work presented in Appendix A of BAW-10168, Volume II uses a nonequilibrium leak volume. The simulation of the ROSA LSTF Test SB-CL-18 did not use an artificial leak volume, since-high pressure injection was not actuated in the test and the use of. the artificial volume during the accumulator injection phase of the transient would have only caused an already conservatively predicted peak clad temperature to be even more conservatively predicted. Licensing materials submitted in Reference 2 use an equilibrium leak control volume. The decision to use equilibrium-or nonequilibrium is considered a user prerogatives it provides the analyst with a degree of control over certain nonphysical events. The type, equilibrium or nonequilibrium, of the artificial leak control volume is not of: primary importance, l provided, that the volume is properly _ performing its role of L limiting the flow of~ECC water to the reactor vessel. FCF's preference is to use nonequilibrium in the fictitious volume, making it consistent with other loop control volumes. FCF's initial Sequoyah SBLOCA spectrum calculations showed severe nonphysical leak flow oscillations. In some instances, it proved t FCF Non Proprietary l
difficult to run the spectrum cases to completion. This type of behavior was not seen in previous topical report or application calculations. The Sequoyah results were investigated to determine the cause of the oscillations, but that effort was not totally successful. It was determined that by running the artificial leak volume as equilibrium the nonphysical flow oscillations were reduced, but not completely eliminated. In l either event, the control volume was properly performing its role as ECC mitigator, so it was determined that the leak volume should be modeled as equilibrium for the Sequoyah calculations. l The SBLOCA spectrum presented in BAW-10220 was the result of that decision. Subsequent to our NRC-TVA-FCF conference call (on 2/20/97), FCF again reviewed its Sequoyah SBLOCA analysis. A mismatch in the plant model inputs for the leak junction was discovered. The two-phase Moody choked flow discharge correlation with the no-slip (homogeneous) option - as required by the evaluation model (Reference 1, Table 26-1, page LA-99) - was used at the leak junction (Junction 504) . However, in the regular RELAPS inputs, applicable to all junctions including the leak junction, for~use with the momentum equations, the nonhomogeneous junction option was specified; hence, the input mismatch, Junction 504 was specified as both homogeneous and nonhomogeneous. At sach time step, RELAP5 performs an intermediate velocity calculational using its momentum equations. If countercurrent flow is predicted at the leak junction, it is assumed that there is no choking and the choked flow calculation is bypassed. If coeurrent flow is predicted, the leak flow calculation is performed using the Moody choked flow correlation. In this current investigation, it was also noted that the nonphysical oscillations correlated well with instances of countercurrent flow being predicted between the leak and containment control , volumes. It was discovered that the leak was not choking at
-pressures incompatible with the prediction of nonchoked_ flows.
The countercurrent flow situation can be resolved by using consistent homogeneous inputs for the leak junction. Since the Moody leak flow calculation is based on no-slip (a homogeneous junction), the solution of the momentum equations should also be homogeneous. This precludes the prediction of a countercurrent leak flow. The leak flow will be coeurrent, the Moody correlation will be used to predict two-phase choked leak flows, and the leak junction will be choked when appropriate. While this insures conformer.ce with the approved evaluation model, it still does not explain the cause of the countercurrent leak flows. The actual cause of the countercurrent leak flow and the nonphysical leak flow oscillations was determined to be a large momentum pressure drop between the containment volume and the artificial leak volume. In the Sequoyah SBLoCA plant model, the leak control volume area was set aqual to the cold leg pipe area, ; while the containment control volume area was set equal to the - FCF Non-Proprietary -
leak junction area. The areas are used to compute control volume velocities, which enter into the momentum flux calculations as velocities squared, creating a containment dynamic pressure about 10,000 times greater (for the 2.5" break) than that in the leak volume, necessitating the RELAP5 momentum equations to calculate nonphysical flow behavior at the leak junction, including countercurrent flow. The reality of the situation is that the containment velocity should be close to zero and there should be no momentum flux gradient forcing flow from the containment to the leak control volume. When the containment volume area is set equal to the fictitious volume cross-sectional area, there was no countercurrent flow, thera was no significant momentum pressure drop, and the leak flow was choked, stable, and predicted by the Moody critical flow correlation. Simply put, the leak flow no longer varied between choked and unchoked f3ow and the leak flow oscillations disappeared. A nonequilibrium 2.5" break case was rerun correcting only the momentum pressure drop problem. The results showed that there were no nonphysical leak flow oscillations, further verifying that the basic leak modeling issue was the improper setting of the containment volume area, the. momentum flux problem. The results of this leak modeling investigation are consistent with the behavior of previous application and topical report calculations. A review of these calculations showed that the leak and containment control volumes were modeled with areas equal to the cold leg pipe area. Nonhomogeneous leak junction cases, such as these, with proper leak and containment volume area inputs, choking was predicted until the system depressurized sufficiently to unchoke the leak. Since previous non-Sequoyah cases properly modeled the momentum flux between the leak and containment volumes, they experienced no countercurrent leak flows and no oscillatory leak flows, and they remain as valid cases. With the leak modeling problems identified, understood, and brought into compliance with the LOCA evaluation model (Reference 1), the SBLOCA spectrum was rerun using consistent homogeneous-leak junction inputs, a nonequilibrium fictitious leak control volume, and the containment and leak volume areas set _ equal to the cold leg pipe area. The results of the reanalysis are discussed in-detail in the response to Question 1, as revised._ A modest clad temperature excursion, approximately 1200 F, has been _ predicted for the 2.75" break case and no nonphysical leak flow oscillations are noted. The work fully complies with the NRC-approved SBLOCA evaluation model, as, documented in BAW-10168, Volume II. No evaluation model excepti,ons have been taken and no evaluation model difficulties were encountered.
- 1. BAW-10168P-A, Revision 3, RSG LOCA, BRNT Loss-of-Coolant Accident Evaluation Model for Recirculating Steam Generator Plants, B&W Nuclear Technologies, Lynchburg, Virginia, December 1996. ~
- FCF Non-Propnetay E - . _ _ . _ . _ _. -_ _ _ _ _ _.__.-____. _ -_.
- . . - . . - . - - - - . ~ . .
.-....;-.~_..----..-_. ..__~ - -. - . _ . . . . . - - . . -
i
- 2. . J. H. Taylor-(FTI) to Documeint Cor. trol Deak (NRC) , .
" Supplementary Information to FTI'c Response to NRC's Request for Additional Information . . . , " JHT/96-46, July 15, 1996. l t
i O t FCF Non-Proprietarj
- 9. Has t5ie core down flow bypass l'n the baffle reg' ion-been explicitly modeled in-the LOCA and non-LOCA analysis? Describe '
how it-is modeled,'
-Response LOCA Nodalization is included in the RELAP5 large break model for blowdown analysis,-and in the small break model, that characterize the baffle region (pipe component 350) . The baffle i region-discretization is similar to that of the core bypass region -- (pipe component '34 6) . Representation of the baf fle region is the same as that of predecessor McGuire/ Catawba and Trojan RELAPS models with the exception that flow enters the top of the
,' region and flows down to the vessel lower plenum, reflecting the down flow design at Sequoyah. The flow through the baffle. region is small-and has minimal effects on either LOCA transient. In the REFLOD3B large break model, the core baffle region is combined with the downcomer. This i's conserva*.ive for reflood as > the "effeetive" downcomer fills more slowly and more ECCS mass must be added to the downcomer before it is filled to the bottom of the_ cold leg nozzle. i Non-LOCA The downcomer model for che non-LOCA analysis includes the volume of the downcomer and the core downflow bypass in the. baf fle 4 region. The downcomer is divided into two azimuthal regions (control volumes 304 and-306, and control volumes 370 and 372) as shown in Figure 6.1-1 to allow charact'erization of asymmetric loop conditions. . i c if FCF Non-Proprietary
4
- 10. With regard to the fuel design features changed from the approved topical _ report BAW-10172, describe the changes in greater detail. The submittal is unclear with regard to the bottom nozzle changes. Have these changes been approved by the staff?- Describe what testing and reanalysis has been performed on assemblies with the described changes (structural, flow, CHF). Please verify that the structural analysis performed in Chapter 8 of the topical report includes the changes identified.
Response
As described in Section 3.1 of BAW-10220P, there are four fuel-assembly design dif ferences on tha Mark-BW Seguoyah fuel design that differ'from the design defined in the approved BAW-10172 topical. These are:
- 1) reduction in-number of restraining guide thimbles from twelve to eight per grid,
- 2) attachment of the ferrules using dimples rather than resistance welding,
- 3) utilization of a debris-resistant bottom nozzle rather than the standard bottom nozzle, and incorporation of gadolinia fuel pellets and axial 4) blankets of either natural or low enriched 00 fuel 2
. -pellets.
The axial positions of the intermediate spacer grids on the Mark-BW for Sequoyah are maintained by the use of ferrules that are attached to ( } guide thimbles. The short ferrules, attached to the guide thimbles, provide direct axial interference with the spacer grid interior strips around the guide thimbles to inhibit the grids from shifting above allowable positions. The ferrules are attached at positions that allow the spacer grid to " float" over a short axial distance. Refinements on the ferrule positioning requirements and manufacturing tolerances have permitted the reduction of necessary ferrules from ( ] per grid to ( ) per grid. ,
- The ferrules were previously attached to the guide thimbles using However, FCF has a resistance weld as described in BAW-10172P.
developed an improved attachment connection using a dimpling of the guide thimble into the surrounding ferrule produced from a force applied to the guide thimble inner wall. This connection has the improved performance characteristi'c of a softer interface that distributes the axial loads into the restraining guide thimbles and compensate for telerances on ferrule positioning.
~
The bottom nozzle on the Mark-BW for Sequoyah contains small circular flow holes that are more effective for filtering debris - f rom the coolant than the large flow slots on the earlier Mark-BW design described in BAW-10172P. The debris bottom nozzle design FCF Non-Proprietary
> 9 is the same design as described FC in BAW-10172P. ~as an alternative Mark-BW bottom nozzle
- nozzle design after perfo,F selected.the rming debris-resistant debris trapping bottom effectiveness tests to verify the adequacy of the flow hole size and performine ~
full-scale pressure drop tests to verify the acceptablility of , the pressure drop impact. The small pressure drop increase associated with-the debris-resistant bottom nozzle has been incorporated into the analyses dependent on the hydraulic
- characteristics of the fuel design.
Gadolinia fuel pellets and axial blankets are being utilized in 4 c.te Mark-BW for Sequoyah to provide desired operational flexibility and improved economics. Both of these features-have proven records of acceptable in reactor service in other FCF ! supplied cores. All of the aforementi6ned fuel assembly design changes have been developed in the last decade, incorDorated into the Mark-BW product line, and have accumulated extensive reactor performance experience. As each design change was developed and implemented, FCF followed previously approved analysis methods to verify the change was acceptable. However, FCF did not submit a request for NRC approval for each intermediate design change as product upgrades or minor evolutionary changes to the fuel assembly design can be expected in any of the vendor fuel assembly design. BAW-10220P is the first submittal that collectively identifies all the design changes. The incorporation of the debris-resistant bottom nozzle was the only change that required full-scale pressure drop testing to quantify the impact on the fuel assembly pressure drop. Tests performed in a cold water flow loop have shown -the design change increases the fuel assembly pressure drop by () This increase has been factored in the reload analyses for the Mark-BW. The hydraulic impact of the' changes in the number of restraining guide thimbles and the type of ferrule attachment is small-and has been determined analytically based on relationships benchmarked to pressure drop tests. None of ,the design changes have required the need to perform additional critical heat flux (CHF) tests. A detailed finite element model for the anti-debris bottom nozzle was developed for use in the structural analysis. The finite element model was benchmarked against bottom nozzle strain data obtained from the fuel assembly drop test (FADT). Based on the benchmark of analysis results to FADT data, it was assured that the anti-debris bottom finite element godel provides a close representation of the anti-debris bottom nozzle design. Further, l it was determined that the maximum stresses determined from this - model would be a good representation of maximum stresses expected - to experience by the anti-debris bottom nozzle given the same loading. FCF Non-Proprietar/
~
.+
- The structural analysis performed in Chapter 8 of BAW-10220P, Rev. 0 for the Sequoyah plant specific loads- includes the changes
- identified. The stresses in the anti-debris bottom nozzle of the Mark-BW fuel assemblies for normal operating conditions as well as for 'f aulted conditions are reported in Tables 8.1 and 8.5 of BAW-10220P, Rev. O respectively. The worst case loading for normal operation is at the design flow rate plus ( ) scram
, load. The minimum margin of safety is () for metrbrane plus bending stress, so that the stresses are acceptable. The minimum nargin of safety for the Sequoyah LOCA plus SSE loads is [] for membrane plus bending stress. The allowable stresses are determined as set forth in the ASME Boiler and Pressure Vessel code. c 1 1 1 i FCF Non-Proprietary
- 11. The pressurizer heaters and sprays are not modeled for the ,
non-LOCA analysis. The results of some transients are worse if , these control features function (i.e., the peak steam generator j . pressure can be higher if the sprays act to delay a reactor trip on high RCS pressure). The staff safety evaluation (SE) on the methodology requires consideration of the control features. Describe why these control features are not modeled.
Response
There are a total of six safety analysis events analyzed for Sequoyah reload. Each one will behave differently when the pressure control system is assumed to function. The effects of l the pressure centrol system are considered for each one of the j transients analyzed. They are described below.
~
1 RCCA Withdrawal at Full Power The 75 pcm/sec (maximum withdrawal rate) RCCA withdrawal case i provided did not assume pressure control system operation so that a maximum primary syst'em pressure could be calculated to assure that the RCS pressure acceptance criterion is met. However, a large number of RCCA withdrawal cases with various reactivity , insertion rates were analyzed to verify that the minimum DNBR is
- greater than the limit value in cll cases. Those analyses assumed that the pressure control system, most notably the pressurizer PORVs, controlled the primary system pressure to 2350 psig. Use of the pressure control system in there cases provided the minimum DNBR by limiting the increase in system pressure and-by requiring a reactor trip on over-temperature AT.
In conclusion, independent analyses were performed for this event that either utilized or ignored the effects of pressurizer control to minimize the margin to the DNB and RCS pressure limits, respectively. Loss of Electric Load In an LOEL, operation of the pressurizer pressure control system j mitigates the primary system pressure increase, providing a less l conservative peak primary system pressure relative to the RCS l pressure limit. Therefore, FCF originally analyzed the event l without primary pressure control and presented these results in l Section o6.. 2 7 of the reload topical, BAW .'1022 0 . I l Operation of the pressure control system, specifically, the l pressurizer PORVs and sprays can. delay reactor trip. This can l-increase the peak secondary system pressure as compared with the i case where the pressure control system is not assumed to l function. l FCF Non-Proprietary
. - _ - . .. -- .- -__ .~--.-.--._ _ - - - . -
Y FCF has reanalyzed the LOEL event assuming operable pressurizer l pressure control. The sequence of events for this analysis-is l attached ~as Table 11-1, Plots of important parameters resulting l from the analysis of LOEL with. pressurizer pressure control are l included in Figures 11-1 through ll-2. These calculations, [ combined with-the results of the LOEL calculations previously l reported in-the-reload topical report, demonstrate that the plant l design is such that a total loss of external electrical load l without a direct or immediate reactor trip - with or without l pressurizer pressure control - presents no hazard to the l integrity.of the reactor coolant system or the main steam system. l Pressure-relieving devices incorporated in the two systems are l adequate to limit the maximum pressures within the design limits. l I The integrity of the core is maintained by operation of the l reactor protection system and the minimum DNBR wil'. be maintained l above-the design value - with or without pressurizer pressure l control. Thus, no core safety limit will be violated as a result l of LOEL. I Table 11-1 LOEL (With Pressure Control) Secuence of Events l l Event- Time, sec l w/ PC l l Transient initiation 0.0 l l OTDT signal 6.8 l I High pressurizer pressure signal 7.9 l
}
Rods begin to drop 9.9 l l Peak pressurizer pressure reached 12.0 l l MSSVs lift. 10.8 l l Peak secondarv oressure reached 16 6 l l Loss of Forced Flow The system response shown for this event assumes that the pressure control system does not operate,. This provided the greatest peak pressure for the event. However, the core DNB response is calculated assuming that the core outlet pressure remains at the initial value throughout the event, yielding the most restrictive value for DNBR. Consequently, the results for this event are conservative, regardless of the availability of the pressure control system. FCF Non-Proprietary
e e Locked Reactor Coolant Pumo Rotor The system response shown for this event assumes that the
~
pressure control system does not operate. This provides the greatest peak pressure _forlthis event. However, the core DNB response is calculated assuming that.the core outlet pressure
- remains at the initial value throughout the event, yielding the most restrictive value for DNBR. Consequently. the results for this event are conservative, regardless of the availability of the pressure control system.
Main Steam Line Break-
- This accident results in a depressurization of the reactor coolant system. Because pressurizer heaters would act to increase the system pressure, which is non-conservative with respect to minimum DNBR, they are not modeled.
Steam Line Break Coincident With Rod Withdrawal at Power This event is a Condition IV steam line break with a coincident l withdrawal of .the regulating control rod bank by the reactor i control system. The event is terminated by reactor trip on a low steam line pressure SI signal or on an over-power AT trip. Neither of these functions are affected oy the primary system pressure control system. Furthermore, tae DNB response to this event is calculated assuming the core exit pressure remains at the initial value. Consequently, the calculated results are conservative regardless of the availability of primary system. 4 pressure control. [ a f 4 . FCF Non-Proprietary i 1
Fip*1re 11-1 Loss of Electric Load - w/PC xi o" -Neutron Power 48
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, i 8- . .j . . . - . . . . . . . ..p., .,,,.j._...,,,......,,,
i e I
~l
! 2 i 0
O 4 8 12 16 20 24 Time (sec) Figure 11-2 Loss of Electric Load - w/PC Reactor Vessel Lower Plenum Pressure 3000 i ! 8
! i -
! i ..
3200- --- - -- --. - - - -
?-- --
7-4 ! 4
- - - -- - j-- -- j g2800- I k
i LL$ ! i ! E l ! D ' m 2400- - -..-. -
-- t- -- - -
,j z :
0- +- - - .-- . r- f- - - -- l l .: i i
; i 3
I k h 1600 O 4 8 12' 16 20 24 Time (sec) FCF - rJoa Pap <ie-ke y
Figuro'11-3 Loss of Electric Load - w/PC Steam Generator Downcomer Pressure
.1600 i i i 1400- --~~ - , - - - - - 1 - -----4 --- - .--,:
< l !
g1200- ~~-~~- - - -
+ - - -
~ ~ ~ - - . - -
N o i 40 l 1 f' 13 4 i x 1000-n.
. - -g j
-~=r- -- n - ~~-
i i i i 6
; i i i -
1 - t ! 800- ~~ 1 -- r--- -- 7- ; l 600 O 4 8 12 16 20 24 Time (sec) Figure 11-4 Loss of Electric Load - w/PC Average Core Fluid Temperature 605 j 4 i i i ! 600- - - - t-v--- ~~ 4- -~ ~ - -- i l 595- 4- i-
/-+! -
+-i
- i ! I ut i ) i
< x sgo. . . . ; -
. ..L M I i i i i ;
i ! g 585- --: 4- + - -- G Y . 580- --
; r *- -+
575- -
- -? -e ?- + - - - - - - - .
570
~
0 4 8 12 16 20 24 Time (sec) ((f ko4 fA4f S c. hA Y
)
Figure.11-5 Loss of Electric Load - w/PC i Pressurizer Water Volume 2400 . . i i i = i
! $ I e i !
2000- - j -.. -.4.--- -
...l.
I i l
. 1600- -- - -
i l d 1 2 1200- -- 4.- ; .;- v-- 1. .. . p i i t J l i O 7 I i i ! e i i 800- - - - -
+
.+ i
..-4....-..
i 4QQ. - . . . - . . . . . . ~ . . 7,,...... ,,,p.,..,.,,,,,,,,,,,,, 1 l I O O 4 8 12 16 20 24 Time (sec) 4 t 9 fcf Ab" Atef ri' b" Y
-- -;~- . - - . -- .-. _ . - - - _ _ .
- 12. Discuss in greater detail the implications of no longer modeling a " hot channel" and an " average channel" for non-LOCA transient methodology (described on p. 6-4).
Response
The consolidation of the hot channel with the average core is l not, in FCF's opinior, a change to the NRC-approved safety l ! analysis topical repor;, BAW-10169. The obj ective of BAW-10169 l is to show that RELAPS/ MOD 2-B&W, with appropriate modeling, is a l viable tool for predicting the transient reactor system response l of recirculating steam generator plants as part of a non-LOCA l safety analysis. BAW-10169 provides examples of representative l safety transients using a Westinghouse 4-loop plant nodel. l However, calculational examples of all safety transients and l varied plant configurations are not presented. l l Nodal variations from Figures 6.1 and 6.2 (Reference 2) are l allowed by the SER; they need to be justified and approved by the l NRC'for the application (not generically) under consideration -l (Reference 1, page 14, Conditions and Restrictions 1). l Specifically, the SER accommodates node changes by requesting l that noding details should be justified on a plant-specific l basis, that is, for each specific licensing application. l Therefore, within the context of the SER, noding changes are l allowed. The change in question is technically justified beluw. l l The hot channel originated from FCF's LOCA model. It was a l residual modeling appendage from the RELAPS large break model. l While meaningful in LOCA calculations, the RELAPS hot channel l plays no role in transient safety calculations. Furthermore, the l hot channel is a detractor with regard to computational. I efficiency because it requires added documentation and l -computational resources, hence, FCF's desire to combine the two I channels. l 1 Thermal-hydraulic calculations of the hot channel are performed, I independent of the RELAP5 model, using FCF's NRC-approved LYNXT l code. They are not performed using RELAP5. RELAP5 predicts the l general system transient response, a role requiring the use of l only an average core channel. Only RELAPS system and average l core channel predictions are used in the. calculational chain of a l safety transient. Considering that the hot channel is composed l of one of 193 total _uel assemblies (0.3 percent), it is easy to l conclude that modeling or not modeling a' hot channel is l irrelevant to and will have no influence on the RELAPS-predicted i system and average core transient progression. Tc further I demonstrate that fact,. FCF has rerun the loss of electric load I (LOEL) transient using a RELAPS mo?'l of Sequoyah with hot and l average core channel simulation as shoun-in Figure 6.1 (Reference l FCF Non-Proprietary
_ = - -. -. . - ,_. - . - - . - . _ _ - . . - _ . . . . . 2). FCF'has compared the results of the " hot-channel model" .) reanalysis with those presented in BAW-10220. Comparison plots l
- of important parameter responses to LOEL event are attached as l Figures 12-1 through 12-4. Note that the results of the two l cases are nearly indistinguishable, demonstrating the lack of l significance of hot channel modeling. l l
FCF's modeling change, incorporating the hot channel into the l 3 average core, is within the SER guidelines for BAW-10169. The l l noding change has been justified and shown to have no l calculational impact. FCC believes that consolidating the hot l channel into the average core channel is an appropriate noding l change and that the modeling technique has been justified for use l in the Sequoyah non-LOCA analysis. l
'l
. 1. A. C. Thadani, NRC, to J. H. Taylor, BWFC, Acceptance'for l Referencing of Licensing Topical Report, BAW-10169, "RSG Plant l Safety Analysis," August 20, 1989. l ! I j 2. BAW-10169P-A, "RSG, Plant Safety Analysis," B&W Safety Analysis 'l Methodology for Recirculating Steam Generator Plants, B&W Fuel l Company, Lynchburg, Virginia, October 1989. I l < l 4 1 0 6 4 FCF Non-Proprietan T . -, , . . _ . , . - .
4 s f; k
. Figure 12-1 Loss of Electric Load Normalized Core inlet Flow 1.2 I 4
LEGEND 1 r j: X Hot Channel Model 1.0 - 0 - Original Model 3=
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. T F 1 l
_ o.2 - ---- t - 0 -. -+ - - - -- c e i 0.0
! O 5 10 15 20 25 30 35 40
}~ Time (sec) 4 [ Figure 12-2 Loss ~of Electric Load Core Exit Presssure 3000 r LEGEND j-... , , j ! N Hot Channel Model O original Model
=2750- *-- + - - -
s 1 i
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j .- -0: 5- -10 15 20 25 3J 35 40-Time (sec) 4 4
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Figure 12-3 Loss of Electric Load Normalized Core-Power 1.2 LEGEND M Hot Channel Model 1.0 . - + - - . - - - - - + -e-- O original Model l 0.8 - -~~~"- 4--- - -r-~--m-~~~~ +; v= e----- - - - - - B- I. s
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0 5- 10 15 20 25 30 35 40 ' Time (sec) i Figure 12-4 Loss of Electric Load Core inlet Temperature 620 1 !
~
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_ _ _ _ _ __m __ _ _ _ .e
- 13. The staff SE for BAW-10169 states that the acceptance criteria for atlocked.RCP rotor is the 95/95 DNBR criteria; however, the analysis predicts DNB and applies-the acceptance criteria for infrequent incidents (Condition III). Please correct this depart.ure from the approved methodology.
Response . The 95/95 DNBR limit is used as the fuel failure criterion as per ' BAW-10169. Any pin : hat fails this limit is considrred as failed in calculating the effects of failed pins on the radiological consequences of tne event. The prediction shows leas than 5 percent fuel pin failure, using this criterion. Exiating dose calculations assume 10 percent failed pins. Because less than 5 percent pin failure is predicted in the calculations, the dose consequences of this event are bounded by the existing calculat. ion. The fuel / clad temperature limits are used to demonstrate that the core stays in a coolable configuration. e FCF Non-Proprietary
14: T.he acceptance criteria established (pp. 2-3_ and 7-1) for events of moderate frequency (condition I and II) includes a 99.9% probability that "DNB will_not occur core wide." The SRP acceptance criteria requires "at least 99.9's of fuel rods in the core will not experience DNB" (SRP 4,4 -3) . The two acceptance criteria are not equivalent. Please correct or clarify the difference.
Response
The wording of the DNBR acceptance criterion as established in BAW-10220P (pp. 2-3 and 7-1) was a misstatement, However, FCF analysis _ practice is correct and consistent with SRP 4.4-3, Specifically, as stated in BAW-10170P-A " Statistical Core Design for Mixing Vane Cores" (p. 1-7) , "The ' SDL of 1.345 (subject to core-specific verification) developed in this report (BAW-10170 P - Al provides - 95 percent protection at a 95 percent confidence level-against hot pin DNB. The corresponding corewide protection on a pin-by-pin basis using real peaking distributions is greater than 99.9 percent.' This agrees with SRP 4.4-3. FCF has performed the ' core-specific verification of the SDL for the Sequoyah core and shown that the 1.345 SDL remains valid. 1 FCF Non-Proprietary l
i 15, it' is unclear from the_ submittal which of the Chapter 15 analyses were redone to support the - fuel change .and - which ones ~ were .re-evaluated. Table- 6.1-1 is not censistent with the "
. verbiage in text on a number of examples. Table 6.1-1 .only lists six transients that were reanalyzed; however, analysis results are discussed for other transients (for example, misaligned RCCA discusses results in Section 6.2.3).
To clarify the situation, state for each Chapter 15 transient whether it was reanalyzed using the Framatome methods with acceptable results or why reanalysis _ is not necessary (why , current analysis remains bounding with the new fuel and a lower RCS flowrate). Include both the effects of the new fuel and the lower RCS flow.
Response
4 Table 6.1-1 presents the non-LOCA " transient events" that were
- analyzed and those that-were evaluated in support of fuel reload at SON. The analyzed events include:
- 1. ' Rod Cluster. Control Assembly (RCCA) Withdrawal at Full Power
( 2. Loss of Electric Load
- 3. Four Pump Coastdown
. 4. Main Steam Line Break (MSLB)
- 5. Locked Reactor Coolant Pump Rotor
- 6. Steam Line Break with Coincident Rod Withdrawal at Power o
Several' events, the misaligned RCCA included, are not transient events in the strictest sense; time-dependent hydraulics (RELAPS) calculations-are not performed for these events. They may be, however, either evaluated or analyzed as static physics exercises for each cycle in support of-individual core designs, A new table has been constructed - Table 15 to indicate the non-LOCA events that _are: (1) analyzed with RELAPS to support the loading of Mark-BW fuel at SQN, (2) evaluated only, and (3) analyzed / evaluated as part of the 2 core design. Note that errors contained in Table 6.1-1 (the_ loss of , feedwater was not analyzed, the loss of coolant flow was) have been Lcorrected in Table 15-1. Section 6 of the topical contains an event by event evaluation of FSAR: chapter 15 transients in support of the fuel reload,_only. The 6 transie.nts chosen for re-analysis compr'ise the most limiting event in the SRP event classifications --overcooling, heatup, loss of RCS flow, and reactivity anomaly - and are re-analyzed with FTI(Framatome Technology Inc.) methodology to demonstrate compliance of the Mark-BW hs'
_ _ _ _ - , - m._ _. . . - _._ .. __. . _ _ _ _ _ _ _ ._ _- ._ 1 T
. fuel-with the relevant success criteria for each event. Analyses in chapter 6 are a direct application of, and conform to the approved ,
methods'of, BAW-10169. Safety analyses supporting the operation of SQN with Mark-BW fuel , were performed with lower-limit flow rates of 348,000 gpm. Use of lower flows in the analyzed events is conservative, with the exception of steam line break.and steam line break coincident with rod ~ withdrawal at power (the steam line break transient was performed with zero percent steam generator tube plugging to maximize primary heat removal and the RCS flows for the event are higher as a result) , j of'the various acceptance criteria, the greatest negative impact.of + RCS flow reduction is on DNB margin. It should be noted that the !- limiting events with respect to RCS flow are contained in the events chosen fer the support of the fuel reload as the events that are also i bounding in DNB (Condition II - RCCA withdrawal at power, Condition 4 III - four pump coastdown, and Condition IV - locked rotor) . All of the limiting event successfully demonstrate-adherence to the relevant acceptance criteria. FCF will, howc ar, compile an evaluation of the Chapter 15 events that is independent of this response. The evaluation will directly demonstrate compliance with the NRC topical query regarding lowered RCS technical specification limits. [" The safety evaluation supporting the reduced thermal design flow for l SQN is attached. , i n 4 4 i d 5 4 e d t
7 Table 15 - 1 Summary of Non-LOCA Assesseent for Reload with Mark-BW Fuel Pull Safety Qualitative core "U
* (RELAJ5) Evaluation Design Analysis Analysis
~
Uncontrolled Rod Clurte.r Control' Assembly
- Bank Withdrawal From a Subcritical Condition Uncontrolled Rod Cluster Control Assembly Bank Withdrawal At Power Dropped RCCA/ Bank l x Misaligned RCCA
- I x Boron Dilution _ l x X
Partial Loss of Reactor Coolant Flow [__ Loss of External Electrical Load and/or
- Turbine Trip '
Loss of Normal Feedwater x , Loss of Non-Emergency AC Power to the Station x Auxiliaries Excessive Heat Removal Due to Feedwater x System Malfunctions Excessive Increase in Steam Flow I x Accidental Depressuri:ation of the Reactor x Coolant System Spurious Operation of the Safety Injection x System at Power I . Inadvertent Loading of a Fuel Assembly into x an Improper Position Complete Loss of Forced Reactor Coolant Flow x Single Rod Cluster Control Assembly x
- Withdrawzl at Pull Power Rupture of a Main Steam Line x Major Rupture of a Main Feedwater Pipe X Steam Generator Tube Rupture x Single Reactor Coolant Pump Locked Rotor x Rupture of a Control Rod Drive Mechanism X Housing (Rod Cluster Control Assembly ,
2 Ejection) . I I Steam Line Break at Power With Coincident Rod
- Withdrawal FCF Non Proprietary
. .. . .. . , ~ - .
l 1
- 15. Thia request is in addition to the original question 15-regarding the sequoyah' safety evaluation for reduced thermal 1 design flow.
i-Arguments made in defense of DNB acceptance criteria'for the feedwater malfunction (4 0.2.10), excessive load increase i- (4.0.2.11), and feedwater line brcak--(4.0.4.3)- stated that , margin to DNB was assured because there was no. trip'on overtemperature AT or overpressure AT for these events. Please~ clarify these arguments.
Response
-These events generally result in a change in and the t establishment of an increased power level. The extent to l which power is changed is not sufficient to produce a
~
- reactor trip on either overtemperature oT or overpressure-AT. The transient, if it produces a reactor trip, produces
.. a trip that is not affected by a reduction in thermal design j flow. l MA? Limits have been developed at operacional statepoints i corresponding to overtemperature AT or overpressure AT , boundaries that indicate an adequate margin to DNE. Since the transients do not' trip on the overtemperature AT-or overpressure AT trip, they are within the Fu@ boundary. Margin to DNB is, therefore, assured. I Revisions to the relevant pages of the safety evaluation are attached which clarify the overtemperature AT - overpressure AT arguments. 2 1 I l 0 4 6 .
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