ML19322B860
| ML19322B860 | |
| Person / Time | |
|---|---|
| Site: | Oconee  |
| Issue date: | 07/31/1973 |
| From: | BABCOCK & WILCOX CO. |
| To: | |
| References | |
| BAW-1388, NUDOCS 7912060679 | |
| Download: ML19322B860 (150) | |
Text
__ _ _ _ - _ - _ _ __ _ - _ _ _ _ _
m EAV-1388, Rev. 1'
/ .C' July 1973 c,... . .. .
. . . . ,. , , , c . . ,
Iti. l U' 51'i, s
.,{ t ..w., a a w.4 I bSit Iik,5 i ~~
ROOM 018 OCONEE 1 F17.L DENSIFICATION RE.NRT l
(Nonproprietary Version of SAV-1387)
' - ?e.tsion 1 -
klu,in,
@g] g,[- J 6. 4 . _ 3, 1
i Babcock &'Nilcox l
7912060h79
BAV 1388 Rev. 1
, July 1973 l
l>
OCONEE 1 FUEL DENSIFICATION REPORT
- Revision i -
BABCOCK & WILCOX Power Generation Group Nuclear Power Generation Division P.O. Box 1260 Lynchburg, Virginia 24505
Babcock & Wilcox
, Power Generation Group l Nuclear Power Generation Division Lynchburg, Virginia Report BAW-1388. Rev. 1 July 1973 Oconee 1 Fuel Densification Report Key Words: Fuel, De.tsification Ef fects ABSTRACT In July 1973, Babcock & Wilcox filed topical report RAW-10055 (Rev.
1), " Fuel Densification Report", which descr;bes the methods to be used in analyzing fuel densification effects. The body of this report, with Appendix A, presents an analysis of effects on fuel for Oconee Nuclear Puter Station, Unit 1 and supports the operation of Unit 1 at the rated power level of 2568 MWt. Appendix B ansvers the questions of the AEC Directorate of Licensing Regulatory Staf f pursuant to their review of this report.
1
- 111 - Babcock & Wilcox l
- _ _ _ _ _ _ _ _ _ )
I I
CONTENTS Page
- 1. I N I Rol>tTT t o:4 . . . . . . . . . . . . . . . . . . . . . . . . . L-2
- 2. rdScirSloNS. . . . . . . . . . . . . . . . . . . . . . . . . . 2-1
- 1. W r Sl'I.rS . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-1
- 1. f . 6".wer Spike Madel. . . . . . . . . . . . . . . . . . . . 3-1
- 5. J . Ih.rmal An.il vs is . . . . . . . . . . . . . . . . . .. . 3 - J.
1.1.1. tuel reaperature Analysis. . . . . . . . . . . . 3 '.
l.J.J. DNHR Analysis. . . . . . . . . . . . . . . .. . 34 4.'.l. Sumnary. . . . . . . . . . . . . . . . . . .. . 3-I.
5.l. Nu le er Analvsis . . . . . . . . . . . . . . . . . . . . 3-9 l . l .1. React or Prot ection Systers. . . . . . . . . . . . 3-9 s . 1. 2. An.nlysis of Power Distributions Before Densitication. . . . . . . . . . . . . . . .. . 3-9
- l. s.1 Analysis of Power Distributions With Densitication Effects. . . . . . . . . . . . . . 3-11
- l. l.4 Summary. . . . . . . . . . . . . . . . . . . . . 3-15
- 1. ~. . S.i f e t y Analysis. . . . . . . . . . . . . . . . . . . . . 3-29 3 . ?. . l . Gener.nl Safety Analysis. . . . . . . . . . . . . 3-29
- 1. f. 2 . 1443 Analysis. . . . . . . . . . . . . . . . . . 3-31 t.>. :tec h.en i c a l An.aly is of Oconee 1 Fuel . . . . . . . . . . 3-36 3 . '. . l . Cladding Collapse. . . . . . . . . . . . . . . . 3- 36 l.5.2. Cladding Stress. . . . . . . . . . . . . . . . . 1-37 g
- 1. i . 1. Fuct l'e l l e t trradiation Swelling . . . . . .. . 3-17 g l . i . '. . R. . ision input to Creep-Collapse Calculations. . 3- 37 ArtTNul xES A. S urna ry of Stat ist ical Analysis for Oconee 1 As-Huiit Data. . . . . . . . . . . . . . . . . . . . . A-1 15 . Answers t o 101. Quest ions . . . . . . . . . . . . . . . B-1 I.i n t of Tables l
T.ib I ..
3.2-1. Fuel re:nper.it ure Comparison at Low Power Density . . . . . 3-5 1.2-2. Fue I lemperature Comparison at High Pcwer Densicy. . . . . 3-5
- 4. 2 _l . Etterts el fuel Densification on DNBR and Power Margin . . 3-5 E
m .
-iv- Babcock & Mcox
v Tables (Cont'd)
Page I sh t e
- 1. 1 - ! . %)!fientions to Reactor Protection System Setpoints
.md Des ign Pa r.ame t e rs . . . . . . . ... . . . . . . . . . . 3-In I.1-J. "conee 1 - Nominal Rod Positions Fuel Cycle Data.
14uilibrium Steady State . . . . . . . . . . . . . . . . . 3-17
- 1. 1- 3. Four-Day Design Tr.insient Data Tr tsient Bank 1. . . . . 3-18
- 1. 1- 4 100-Day Design Transient Data. Transient Bank 1. . . . . . 3-19 3.1-5. 100-Ny Design Transient Data, Transient Bank 2. . . . . . 3-20 1.1-6 200-Day Design Trans ient Data. Transient Bank 2. . . . . . 3-21
- 3. 3- el. 2H5-Day Design transient Data Transient Bank 3. . . . . . 3-23 1.5-1. Cladding Circumferential Stress . . . . . . . . . . . . . . 3-40 A-1. Sumnary of Stat istical Analysis for Oconee 1 As-Built Fuel Pellet Data . . , . . . . . . . . . . . . . . . . . . A-2 List of Figures Fig ire 4.!-1. M.sximum Cap Size Vs Axial Position . . . . . . . . . . . . 3-2 1.1 - J. Power Spike Factor Vs Axial Position . . . . . . . . . . . 3-3 1.1-!. Maximum Fuel Temperature 's Linear Heat Rate - Oconee 1. 3-6 3.1-2. Average Fuel Temperature Vs Linear Heat Rate - Oconee 1. . 3-7
- 1. J- 1. I;. p I:.,e f f i c ient Vs Linear Heat Rate - Oconee 1 . . . . . 3-8 1.5-1. I rip 5et point s Vs Axial Imbalance Without Densification I. I f ec t a. . . . . . . . . . . . . . . . . . . . . . .. . . 3-24 3.1-2. c'.il culat ed of fset Limits Vs Power. . . . . . . . . . . . . 3-25 1.1-3. Irip Setpoints Vs Axial Imbalance With Densification I.ffects. . . . . . .. . . . . . . . . . . . . . . . . .. 3-26 3 . 1 *. . Percentage Change in Peak Power Vs Indicated Tilt. . . . . 3-27
- 1. 1 - 5 Maximum kW/ft Vs Rod Inde x for Equilibrium and Trmient Xenon Conditions at 102:: af 2568 MWt (BOL) and LOCA Limit of 19.8 kW/ f t. . . . . . . . . . . . . . . . . . . . 3-28 l . '. - l . Oconee Scram and Coastdcvn Curves With 0.62 Trip Delay lime for Four-Pump Coastdown . . . . . . . . . . . . . . . 3-34
- 3. f.- 2. DNBR and Fi ltn Coef ficient Vs Time for Four-Pump roastdown. . . . . . . . . . . . . . . . . . . . . . . . 3-35
-v- Babcock & Wilcox l
v l
- 1. INTRODUG ION This report docueents the ef fects of hypothesized fuel densifica-I tton for the Oconee 1 core as calculated in accordance with guidelines set forth in the AEC report of November 14. 1972. The application of these guidelines for the results presented in this report are t 3 cussed f eallv in revision 1 of topical report BAW-10055 " Fuel Densification Report."
The analysis of Oconee 1 is limited to an examination of the first fuel cycle. Although Babcock & Wilcox has no operating data on the f uel densi f icat ion phenomenon, data to be released f rom pressurized wa-f .* r reac t s.rs (PWRs) now operating with prepressurized fuel are expected to .illow relaxation of the current guidelines.
l 1-1 Babcock s. Wilcox l
l
r I
(
l
- 2. CONCLUSIONS Enhed on t he analysis performed for Oconee 1, which utilized the methods given in BAW-10055, Rev. I the following conclusions are made even if the fuel pellets should densify to 96.5% of theoretical density:
- 1. The cladding will not collapse because all B&W fuel rods are prspressurt/ed.
- 2. The mechanical performance of S&W fuel rods will not be in-pa t t e J.
- 3. The interim acceptance criteria for the emergency core cooling system (ECCS) will not be violated.
4 The reactor can be safely operated at the rated power level of l 2368 MWt with only minor modifications to reactor protection system (Hps) setpoints. These modifications ensure that the thermal design criteria are not exceeded.
- 5. The modifications to the RPS are a reduction in the ovr.rpower trip setpoint, from 114 to 112% of rated power, and a minor reduction in allowable isbalance limits as shosn in Figure 3.3-3.
2-1 Babcock & WiMox I
sa
- 3. RESULTS 3.1. Power Spike Model The guidelines outlined in " Technical Report on Densification of
- 'ght Qter Reactor Fuels," November 14, 1972, have been used to de-
- t. wine t he adFimum axial gap as a function of Core height. The prob-
.shility value: (F ) given in the same report (Table 4.2.A. column 4) have been used in calculating the power spike factor. This factor, as calculated in section 2 of BAW-10055, Rev. 1 is applicable to individual reactors. The maximum gap size versus axial position is shown in Figure 3.1-1. .and the p .,er spike factor versus axial position is shown in Figure 1.1-2. Th.se figure:: also show the initial and final theoreti-c.1 der. sit.' s (TDI, TDF) used in the calculations. These data form the basis for the following analyses.
Th13 section of the report covers four main topics: thermal
.in.a l y s is , nuclear analysis, safety analysis, and mechanical analysis.
the thermal analysis section considers protection of the fuel melt and DNBK criteria. The nuclear analysis s ection considers thermal design criteria, imbalance trip limits, and ccre operational limits. The saf ety analysis section re-analyzes all postulated accidents analyze:1 in the Oconee FSAR assuming that densification occurs. The mechanical analysis section sununarizes the input suminary and results, cladding creep and collapse, cladding stresses, and fuel pellet irradiation swelling.
I I
31 Babcock & Wilcox
2!l ". '.* f f. Iu D" $ % : >=- 2*.08 0
' 4 1
s e
0 h
' 2 c 1 n I
)
Z
(
0 L
' 0 1 ,
h t
g n
e L
0 e 8 v i
t c
A g
L n
0 o i 6 l
- 4- A 0 n 0 o i
0 t
+ i s
i 0 o l
4 P C l T a 0 8 i
. 2 5 4 s 6 3 A F 9 9 O
0 0
[-
T
. 0 2
=
F 1 L O 0 A T 1 _
_ - - - 0 _
4 3 2 1 0 Ug= d J" a *u
. a5. E T @flIw l ,' -
I ll I
I I
, a g u r .- 1..-;. 1 oves Spike Factor Vs Axia; h sitlen I 3 E
I 2.-
I E
I -
l \
E _
m 5 g :
O E2
. =
XQ l o o n H a, l i
E "o
I E
i I a T U. E. E. E. I sotoej avds Josed i
Babcock a.Wilcox 3-3
I L 2. *n.rnal Analys!3 i
....l. l'u. l Temperature Analysis iJti1 izing the analysis established in BA*J-10055, Rev 1. , the fuel-to-cladding cold diametral gap af ter densification was calculated as
- 12. 4 mil . 1his gap size and the TAFY fuel pin temperature and pressure tode ( !.N.'- 10044 ) were used to perform a fuel temperature analysis.
6 rage and maximum f uel tec:peratures for densified fuel at the beginning ut l i f - ( 801.) . at 100 days, and at the end of the first cycle of opera-tIan are compared in Tables 3.2-1 and 3.2-2 and in Figures 3.2-1 and 3.1-2. The gap heat transfer coef ficients for these same times are com-pared in Figure 3. 2- 3.
3.2.2. DNBR Analysis lhe thermal effects due to densification can be divided na into twa categories: (1) the result of the reduced stack height and (2) tir (umbined result of the reduced stack height with the power spike "
superi posed. Therrul ef fects are then imposed on calculations of the mininu2 departure fruct nucleate boiling ratio (DNBR) used to set thermal dt ign 1imits.
The reduceo active length was calculated to be 141.8 inches, which represents a reduction of 2.17 inches from the nominal active length of 143.97 inches. The as-built information given in Appen-J n A i.a s used in calculating this densified active length.
The axial flux shape that gave the maxt:su:s change in DNBR _
iron the original design value was an outlet peak with a core offset of u All.SI. The spike runnitude and the maximum gap size used in the analy-sia are 1.125 and 2.62 inches. The results of the two effects are sun-marizel in Table 3.2-3 in terms of percentage change in minimum hot channel DNBM and peaking margin.
- 3. 2. 3. su c=na r y This analysis assumes that densification and associated pheno =wna will affect the hot channel, which has the most limiting therzul-hydraulic characteristics in the core. In addition, the power spike is assumed to be located at thu hot channel posir. ion that mini-miacs M Bn. The rea.ultant 4.46% DNBR loss or 2.01% power peaking 3-4 Babcock *_ Wilcox i
- o I margin reduction will be compensated so that the plant can function at I rated power without violating the initial design criteria for D*f5R and/
or fuel melting. The allowable power shapes and the new offset limits are discussed in section 3.3.
I Table 1.2 - 1 Fuel Temperature Comparison at Low Power Density I Density, Cold gap. Time in Surface Avg fuel Max fuel
- theor mils life kW /ft fuel temp. Ftemp. F .e m p. F
'd . . ; 12.8 ilO L 5.6 904 1228 1384
'" . . ; 12.8 End 100 days 5.6 920 1247 1605
%.; 12.8 End cycle 1 5.6 940 1271 1633 I Table 3.2 -2. Fuel Temperature Comparison at High Power Density Density. Cold gap. Time in Surface Avg fuel ."ax fuel
- theor mils life kV//ft fuel temo. F tee , F ter:o. ?
90.5 12.8 Bf 17.6 133P 2741 4210 96.5 12.8 End 100 days 17.6 1362 2774 *242 9 /. . ; 12.8 End cycle 1 17.6 1389 2808 4274 I Table 3.2 -3. Effects of Fuel Densification on DSTR and Power MarNin @ 114% 2568 W(t)
Densified active Densified ective length length and power spike Axial DNDR %A %A DNBk %A %A power shape (W - 3) DNB Margin (W-3) DNB Ma rcir.
I Outlet peak with +11.8%
1.521 -1.62 -0.67 1.477 -4.46 -2.01 core offset I
I 3-5 Babcock s.Wilcox
I l Figure 3.2-1. Maxinne Fuel Tc=perature Vs Linear Heat g, Rate - Oconee 1 5200 Welting Temperature U02 9 15.000 MID.uTU ,
4800 I
4400 '
End 100 Days J
4000 En3 of Cycle I 3500
\ 2 y % BOL g 3200 n
es 5
E
/ I 1 e I j 2400 End of 100 Days = 5.260 Mr0/MTU j
End of Cycle I = 15.000 EIN11TU 2000 g'
5' 1000 2 f I I
l la00 I
ao 20 50 10.0 14.0 18.0 22.0 26.0 I Linear Heat Rate, kW/ft 3-6 Babcock a.Wilcox
l Figure 3.2-2. Average Fuel Temperature Vs Linear Heat Rate - Oconee 1 3600 3700 //[
End of cycle i .
2800 Eno cf 100 Days 00L
, 2400 /
E O
o 2000 T
h End of Cycle I = 15.000 usD/MTU i E End of 100 Days = 5260 Es0/ETU
~ ISCO ,
f i r 1200 800 i l 0 4 8 12 16 20 24 Linear Heat Rate. kW/ft 3-7 Babcock s.Wilcox
I I
Figure 1.2-3 Cap Coefficient vs. Linear Heat Rate - Oconce 1 1400 I
BOL /
End of 100 Days 1300 )/
End of Cycle i /
1200 I
i1.00 /
~
r
$ 1000 /ll l
b 900 i
/ E 800 /
= / / E
/ End of Cycle I = 15.000 unD/MTU End of 100 Days = 5260 Mr0'MTU 700 / l
/ I 600 500 l
2 6 10 14 18 22 26 Linear Heat Rate, kW,'f t I l l
3-8 Babcock & Mcox l
l - ._ - _ . _ _ . - _ - _ , _ . _
l l
- 3. ?, . Nuclear Analysis l 3.3.1. heactor Protection Sysces The safe operation of a reactor core requires an extensive ;
ar ..l ys is eif power dist ribut icns resulting f rom the various modes of plant o pe ra t i on. The primary considerations and results of thin analysis are ;
as follows:
- 1. Ass u ran ce that thermal criteria are not exceeded; i.e., !
specified minimum DNBRs and centerline fuel temperature a )
may not be violated.
- 2. Definition of imbalance ILnits to prevent adverse power peaks that would exceed the foregoing criteria. l
- 3. Definition of core operational limits anil recommended l
operating procedures to prevent unnecessary reactor trips.
1 The complete maneuvering study entails a combined nuclear-thermal analysis of the gover distributions. This section describes the methods and criteria usert in develcping the RPS setpoints and in modify-ing the set points required to account for postulated densification ef fects.
3.3.2. Analysis of Power Distributions Before Dens i ficat ion The three-dimensional PDQ07 code with thermal Ie back ef-tects is used to analyze power distributions. This analysis detern .
power dist ribut ions for all modes of reactor operation except accidents and other rapid transients. The design power transient (100-30% power and return to 100% at peak xenon) is analyzed throughout core life. Th e-fuel cycle and transient analyses determine po 'r distributions for nor-mal equilibrium and transisnt conditions , respectively. The extremes of core ope rat ion, such as control rod bank insertion beyond normal limits sad maloperation of ax111 power shaping rods, are also examined. The i l
catreme control rod bank conditions define the limits for the inbalance protection systen.
l 1.3.2.1. Correlation of Power Peaks to I Thernal Design Criteria l l
The power peaks from PDQ cases are corrected for l
j calculational uncertainty and are analyzed to determine the margin to the l the rmal criteria: centerline fuel melt and departure from nucleate boil-ing (DNB). The margin to centerline fuel melting is defined as 39 Babcock s. Wilcox 1
I j
l l
l l
I I !
Fuel melt margin = [\ Max calculated peak -l\1001.
/ l The ruxirrum allowable peak is defined as the pointwise power that yields centerline f uel rel t ing:
Max allowable pean, = 22.2 kW/ft 5.656 kW/ft = 1.014 = FDP 22.2 kW/ft = fuel melt limit, 5.656 kW/ft = average heat rate at 2568 MWe, j 1.014 hot channel factor, TOP = f raction of power.
The maximum calculated peak is the largest total peak f rom the PDQ power increased by a f actor of 1.075 to account for calculational uncertainity.
The determination of DNB margin requires a more complex analysis. DNBR is a function of peak location, magnitude of the powe r peak component parts (radial and axial), and other core parameters.
To arrive at t rue DNH conditions, each power distribution is analyzed explicitly. From the PDQ power distribution, the maximum calculated total peak 1s ebtained and adjusted for uncertainty. The DNB margin is then defined as Allowable total peak Max calculated total pea.x , 1}/ 100I. )
I The basis for the allwable total peak is the reference design DNBR at design conditions, or a 1.30 DNBR associated with the protection system envelope, or a quality limit based on model applicability, whichever is g
most limiting. 5 l l
1.3.2.2. Of fset-Margin Relationship Core e f fset , a measure of the axial power im-balance, is delined as the f ract ion of total core power *n the top hal f ,
of the core minus the f raction of total core power in the bottom half of the core: l E i 3-10 Babcock a. Wilcox l
l l
Wer P) - PWer Got td offset =
Power (top) + power (botton),
1 The relationship between hot channel power peaks (i.e. , thermal margins) and core offset defines the protection system setpoints. Power imbalance is the primary signal to the protection system for flux shape protection.
The maneuvering analysis defines the relationship between core imbalance and thermal margin.
Limiting offsets are determined to prevent the violation of thermal criteria for all operating conoitions and power levels.
To yield the imbalance trip envelope, the limiting of fset values are cor-rected for potential instrumentation errors, imbalance detection bias, and calibration. The labalance trip envelope defines the range of allow-able operational imbalance and ensures that a 1.3 DNBR and/or the central fuel melting limit will not be exceeded. Figure 3.3-1 presents the trip setpoints based on these criteria. It is important to note that DN8R criteria define the positive tabalance slope, whereas centerline fuel e,elting is more limiting for the negative imbalance limits. The over-power t rip setpoint shown in Figure 3.3-1 is centrolling for overpower trar.sients, whereas the solid borizontal line is the trip for loss of flow transients.
3.3.1. Analysis of Power Distributions With Dens i fi ca t ion Effects Provision for possible fuel densification requires modifica-tion of the imbalance t rip system for two reasons: (1) the fuel melt (kV/ft) criterion change, and (2) an additional never spike is included in the reactor pwer dist ributions. Since the power spike f actor is a function of axial position, the appropriate power spike factor is used to increase each PDQ peak to account for potential densification.
The modified offset limits with fuel densification effects included ate presented in Figure 3.3-2 and compared with the present limits. The primary dif ferences between the two sets of calculated limits are as follws:
3-11 Babcock a.Wilcox 1
l 1
1
1 I
- 1. The DNBR loss of -4.46% res ul t s in a peaking margin t oo o t ~2.01:
- 2. The cent ral fuel melting limit changes from 22.2 kW/f t I
before densi f icat ion to 21.8 kW/ft.
- 1. A 2.21-inch decret.se in fuel column length increases I the nonsnal heat rate at 2568 MWt f rom 5.656 kW/f t before densification a=
to 5. 7 4 '. k W/ f t .
?. . the loc.at power spike f actor is applied to the calcu-lat est pmer dist ribe fons. um
- 5. The overpower limit in the imbalance protection system.
is refelined as 111' of 256M MWr . The ef fect of the reduced overpower limit is one-to-one for local heat rate and approximately two-to-one for DNBR.
The t rip setpoints are obtained f rom the calculated of fset limits by adjust ing f or potential electronic errors and of fset measure- M rae n t bias by the out-of-core detectors. The error-adjusted limits for densified fuel are shown in Figure 3.3-3. The imbalance trip points and over power t rip provide operating flexibility with assurance that thermal criteria are not exceeded. Furthermore, potential relaxation of the l imi t s may be realired from the physics tests, in which the offset bias inst rumentat ion behavior and the neasured flow in the RC system will be determined.
1.l.3.1. ECCS Considerations ECCS calculations that include fuel densification ef fects have shown that a peak fuel rod cladding temperature of 2291F is achieved with a linear heat rate of 19.8 kW/f t at 102% of 2568 Mit. The margins associated with 14.8 kW/f t have been re-examined for normal e
___ I The 21.8 kW/st central fuel melt limit was calculated for densified fuel g for a maximum, first-cycle fuel pin burnup of 14.694 Pfdd/mtU. The ende g described in HAW-10044, "TAFY - Fuel Pin Temperature and Cas Pressure Analys is," was useil f or this calculation. A first-cycle burnap of 14.694 MVd/mt U includes a 10% uncertainity margin for first-cycle burnup.
3-12 Babcock a.Wilcox 9
2 e qui i fler les, e ps rat lon .and maneuvering transients during the fuel cycle.
Ihe IJHJ. margin la delined as g , _ !.imit in g heat rate (19.8 kW/ft)
,5.744 kW/it = calculated peak = 1.02 - 1 100%
5.744 kW/ft - average heat rate at 2568 MWt for densified fuel, Calc mak = peak calculated with PDQ increased by a 1.075 uncertainty factor, 1.014 hot channel factor, and the power spihe factor (Figure 3.1-2) as a function of axial position, 1.02 - 102% of 2568 Mic.
3.1.1.2. Normal Fuel Cycle Operation Table 3.3-2 summarizes the fuel cycle calcula-tions. the maximum n otal hot spot peak is presented with the correspond-Ing peak linear power density (kW/ft). De IDCA margins represent the n.irgin at the hot spot for a balanced core and normal centrol rod posi-tions. Af ter establishing equilibrium xenon, the IDCA margin exceeds WI throughout the remainder of the rodded cycle; i.e., the peak linear power is 14. 5 kW/ f t compared to the 19.8 kW/f t li;ait . The margin in-creases t h rou ghout the fuel cycle except for the period from 285 to 310 days. At 110 days, the margin is calculated to be 31.27.. The decrease in m.irgin at the end of the first cycle is due to the transient rod bank being fully withdrawn at 310 days to enable the reactor to meet end-of-life reactivity requirements.
3 .1.1. 1. Transient Data 1
i 1
A 100-30-100 power transient is the design transient for the core. The xenon override control rod bank (transient i
bank) is designed to recover f rom 30 to 1002 power at maximum xenon, which occurs approximately 8 hours9.259259e-5 days <br />0.00222 hours <br />1.322751e-5 weeks <br />3.044e-6 months <br /> af ter the power reduction. Tables 3.3-3 through 3.3-8 summarize the calculated data for six transients at various times in core life. Power peaks result f rom the rod positions required to meet the design transient. Af ter recovery to full power, control rods are inserted beyond their nominal insertion range to com-l pensat e for xenen undershoot. The largest operating power peaks occur at this t la;. In Tabic 3. 3-3, the [ ]-hour peak is the largest operating l
3 13 Babcock a.Wilcox
t l
l I i 1
pwer peak. A het spot heat rate of [ } kW/ft occurs at this time, I which gives the t wes t operating LOCA margin of ( }. Th 's power peak is .e re sult of bank 7 being fully inserted and bank 6 being inserted to
.sppr..x a m.s t elv t he midplane. Note that this peak is unrealistically con-serv.it i ve s ince 100 ef fect ive full poser hours will be achieved during 5 start up test ing before the reac'
- is near full power. Af ter 100 full power days, t he max t: sum heat r. are below 16.4 kW/ft and the LOCA nirgins are in excess of 21%.
- 1. 't . 3. 4 ouandrant Tilt in the unlikel) event of a quadrant power tilt during normal equilibrium operation and during design power t ransients, 3 the pri sary concern is r. ensure that the limiting heat rate of 19.8 kW/ft is not exceeded. The present technir- specifications limit quad-rant power t ilt to 102 at rated power. The relationship of increased reaking and radial tilts has been investiRated for various conditions
.end core designs. Fi gu re 3.1-4 presents results for increased peaking
.ns .i function of quadrant tilt as indicated by the out-of-core detectors f or dropped cont rol rods and single and multiple control rods out of sequence. The magnitude of a radial tilt as indicated by out-of-core det e. t ers depends on the relationship of the tilt axis and the out-of- 3 c o rt- Jetectors.
A maximum indication is obtained when the tilt axis is aligned with the out-of-core detectors. A minimum indication is obtained when the tilt axis is at an extreme angle f rom the out-of-core detector.
An uncertainty of 231 is included in the tilt indication in Figure 3.3-4
.. a.osunt for the potent ial dif ference between the nomir:al quadrant tilt
.ind the minimum indicatinn. Figure 3.3-4 is a plot of peaking increase
.ns a function of tilt indication for a minimum indication by the out-of-core detectors. The data were determined with three-dimensional, ther-mal-hydraulic feedback analyses (PDQ-7), except for the dropped rod cases, 5 which were determined f rom two-dimensional PDQ-7 analysis. The data are A(kW/ft %)
limited by a line with a slope of 1.84 A(Indic,' %).
- 1. 3.1. % Rod insertion Limits A minimum LOCA margin of 18.4% is needed at 102Z of 2%N NWt t o accomodat e a 10% quadrant tilt and not exceed LOCA crf-teria with densificatica effects included. A 5% quadrant tilt will 3-14 Babcock a. Wilcox 1
require a minimum LOCA surgin of 9.2% at 102% of 2568 IWt to prevent ex-ceeding 1.OCA criteria with densification effects included. Table 3.3-2 shows that early in core l'.fe under present operating conditions, the minimum LOCA margin is [ ] a: the time of xenon undershoot. Since pcwer peaks in a rodded core are primarily a function of rod insertion, the total power peak can be reduced by centrol rod insertion limits. Figure 3.3-5 is a plot of maximum heat rate as a function of rod index for equilibrium xenon and t ransient menon conditions. Rod index is defined as the sum of the percentage insertions of banks 6 and 7. The figure indicates a range of control rod indices which yield lower peaking con-dittons. linen the core is at power, the selection of a control rod op-erat ing band between rod indices of 52 and 117% will result in lower peaks. Ut 111 ring the foregoing control operating band, the maximu=a heat rate at 102% of 2%8 MWt, with all densification ef fects and a 7.5% un-certainty factor, is 17.82 kW/ft. This rate yields a LOCA margin of II.l!.
The inclusion of a rod index range for power operation will yield a minimum LOCA margin of 11.1%, which is sufficient to .seconunodate a SZ quadrant tilt. Furthermore, maintaining the rods within the 52 to 1172 index range will 1 cad to larpr margins during the early part of core life. The rod index limitation would be required only for operat ton above 90Z power. Below the 90% power level, the LOCA mar-gin . are large enough to handle a 5% tilt without rod index restriction.
Fu r t h e r, rest rictions on a rod index could be removed af ter the first 100 .laya of ful1 power operation. Table 3.3-4 Indicates that even for the normal ro<l Index of 141.7%, the LOCA margin is in excess of the re-qui red 9..f Z.
3.3.4. Summary Fuel densification and associated design limit changes have produced nodifications to the RPS and cperational changes. The power phaking margin loss of 2.01% f rom the DNB analysis, the lower fuel melt-ing limits, and the additional power spike factor have been compens sted by a 2% reduction in design overpower and by more stringent of fset limits.
The revised RI'S allvws operation at 100% power with assurance that the r-mal criteria, with all densification effects included, are not exceeded.
1he modi f f rat ions are summarized and compared with the previous' system in Table 3. 3- 1.
3-15 hock & Mcm
I I
I
'I a ble 3.1-1. Modifications to Reactor Protection System Setpoints and Design Parameters A. Imbalance Sy stem Previous Modified system l'a ramet e r system (densificati on)
Fuel melt limit. kW /ft 22.2 21.8 IJND peaking margin penalty. % --
-2.01 Nominal heat rate, k W /ft 5.66 5.74 Overpower. **,- of 2 5i.M MWt 114 112 Offset limits at rated power l'ositive offset +39 +26 Negative offset -65 -62 T rip setpoints at rated power l 'o s it i ve imba lance +'t +8 N. 2at is e imb.ilanee -30 -32
< pike factor None 1.00 to 1.125 Nuclear power peak uncertainty 1.075 1.075 B. ECCS Coasiderations The normal operating pc,wer peaks incorporating a power spike fac- g tor for densification result in heat rates at 102% of 2568 MWt, which 3 are less than the limiting value of 19.8 kW/ft. The peak heat rate without rod index limits is [ ] kW/ft.
C. Radial Tilt For pows levels above 90% of 2568 MWt, an operating rod index rance from 52 to !!7% will yield a minimum LOCA margin of 11.1%.
This value contains sufficient margin to accommodate a 5% quadrant powe r tilt . The peak heat rate with rod index limits is 17.82 kW/ft.
i t
3-16. Batsock a.Wilcox ,
e
[
[
[
[ i Table 3. 3-2. Oconee 1- Nominal Rod Positions Fuel
( Cycle Data. Equilibrium Steady State-Time. days Offset, f.
0 +5.80 4 +3.18 25 +1.61 50 -3.96 100 +3.43 150 -8.48
( 200 +1.11 250 -8.43 l 285 +2 . 6 3 310 +8. 31 l l
1 I
i 3-u sabcock awik:ox i
i
, . _ _ , . . , , _ . . _ , , , . - . . . _ . _ _ . _ . . , _ , - , . _ -,_ _ _ . - _ _ . .. ____._._.m..m. .._,,__. - , , , - .,,.7
I I
I l l
l Table 3. 3-3. Four-Day Design Transient Data. Transient Bank 1 Time, h Offset. %
96 +3.32 104 + 6.88 IM.5 -15.2 105 -13.7 g
106 +0.1 =
107*I I -12.1 108(a) -10.5 109(b) -9.64 110 -5.73 1II -3.98 112 -2.24 I* I Bank 7 = 100% inserted and bank 6 =
4 5.8% ins e rted.
I Bank 7 = 100% inserted and bank 6 =
- 41. 7 % ins e rted.
I I
i I I
3-18 Babcock & Wilcox I
O
l I
I I
I I
lI l
Table 3.3-4. 100-Day Design Transient Data. Transient Bann I Time, h Offset, %
2400 -4.60 2408 -2.28 2408.5 -16.3 2409 -S.96 2410 +1.20 2411(a) _11,3 2412 -14.4 I 2413 -13.3 I*I Bank 7 = 100% and bank 6 = 41.7%.
i I
L 1
l l
3-19 Babcock & Wilcox l
i I
I' I
I l f
I Table 3. 3-5. 100,-Day Design Transient Data. Transient Bank 2 l
I Time, h Off s et. "e 2400 -4.34 2408 -1.40 2408.5 +2.96 2409 -2.38 2410 +4.02 2411 -0.60 2412 -10.1 2413(a) -12.9 2414 -11.0 2415 -11.9 l
2416 -12.5 I 2417 -12.8
! I"IBank 7 = 100'e and bank 6 = 54"..
l I
I I
I 3-20 Babcock & Wilcox
Table 3.3-6. 200-Day Design Transient Data. Transient Bank 2 Time, h Offnet, c~s 4800 -5.24 4808 -9.50 4808.5 -1.42 4809 +1.48 4810 +4.02 4811 -5.82 4812 -12.5 4813 -16.1 4814 -17.8 4815 -17.0 4C20(a) -21.8 4822 -21.0 4828 -19.0 4836 -8.6 4844 -4.8 I*I Bank 7 = 100% and bank 6 = 45.8%.
3-21 Babcock s.Wilcox
I I
I Table
l 4800
-3.21 4808 -1.91 4808.5 -11.8 4809 -6.u4 4810 +2.34 l 4811 +0.11 )
4812 -6.95 !
4813 -9.21 !
4814 -11.4 4815 -11.0 4816 -11.9 4818 -14.4 4820 -17.3 4822 (a) -16.5 4824 -16.1 4828 -12.3 4832 -7.03 4836 -1.35 4844 +1.65 I*IBank 7 = 1007, and bank 6 = 42b '
I I
3-22 Babcock a.Wilcox
l f
Table 3. 3-8. 285-Day Design Transient Data.
Transient Bank 3 Time, h Offset. %
6840 -8.89 6848 -9.01 6848.5 -6.93 i 6849 +1.67 6850 +2.29 6851 -2.91 6852 -11.5 6853 -13.5 6854 -13.2 6855 -12.5 6856 -13.1 6858 -15.4
(.860 -19.1 6862 -20.3 6864 (a) _39,z 6868 -15.5 6872 -10.4 6876 -6.35 6884 -3.15 t
6892 -1.71 6900 -4.28 6908 -6.91 1
I*I Bank 7 = 100f. and bank 6 = 427..
3-23 Babcock s.Wilcox
I
-I i
Figure 3.3-1. Trip Setpoints Vs Axial Imbalance Without ,
Densification Ef feces
@ s l l10 _
i 10 0 verpower Trip l 90 _ j Setpoint 80
$ 70 -
=
N 60 O
', 50 -
E t 5 40 I
30 -
-I !
20 l IO -
I 0 I ' ' '
l
-50 -H) -JO -2D - 10 0 10 20 30 40 50 Core labalance, %
i 1
3-2t. Babcock s.Wilcox
t i
i i
i Figure 3.3-2. Calculated Of fset Limits Vs Power 120 -
PRIOR TO DENSIFICATICal 110 -
WITH OENSIFICATION EFFECTS 103 -
M 3
E E 91 -
.e 80 -
TO
! - ' I r e t , i j 60 -40 -20 0 20 40 Core Offset, s i
1 3-25 Babcock a.Wilcox
I Figure 3. 3-3. Trip Setpoints Vs Axial Ichalance With Densifica 'on Effects 120 -
I i I l 110 I
- i I
- )
i OVEltPOWER l 100 - l TRIP SETPolNT l
l
=
90 -
= 3 3 5 m - 80 - E ,
- u '
d' l
I 70 -
I 60 -
I l
I
-40 20 0
20 40 I
I Core imbalance, 5
\
3-26 Babcock &Wilcox
Figu re 3. 3-4. Percentage Change in Peak Pcwer vs Indicated Tilt 14 .
A Peak Poser eb) = 1.84 I Indicated felt (S) 10 .
b 15 .
[ g e E Reds Out of Seevence
, ,7 ,
X 5 Reds Out of Saguance c
{ 6 6 Stepped Red a
- 6 .
3 O O Single Red Out of Sequence O
4 .
Note West Conservative OC3. Pest Angle Used 0I . > = . . . .
0 2 e s a le 12 is fadicated Tsit. E l
l 3-27 Babcock & Wilcox m
. , . . -, . . . . . -,m ,.,. , . - --
~
I 6 i gia r e-(.<.nd i t l eang at 102% of 2568 MWt (80L) and LocA I.imit <> f 19.8 =W/fr o
I
~
l l
E o
3
=
2 E
E <
! l e
i a
2 g j g
I l -
1 l
l I f f I e v. un an w i
- s.s. .- - l t y my 'steg teaH xead j l
3-28 Babcocka. E cox
-- , -, .,e , , - - - - - - - - _ , ._- --e ,
1 i
3.4 Safety Analysis 3.4.1. General Safety Analysis 3.4.1.1. Introduction The significant effects of fuel densification are a,n increase in maximum fuel temperature and a slight increase in average heat flux due to shrinkage of the pellet stack length. In addition, spikes in the neutron power can occur due to gapa in the fuel.
these combined effects will lead to a slightly decreased initial DNBR ;
for the accident calculations presented in the Oconee SAR. For over- !
power transients, such as a rod withdrawal, the effects are offset by i i
a reduction in the overpower trip setpoint. The parameters used in 1
the analysis are the same as those used in the SAE analysis. The changes )
in fuel geometry and higher fuel temperature will lead to slightly more '
negative values of the moderator and Doppler coefficients; however, to 1 maintain conservatism the original values were used. All calculations were made for BOI. conditions.
j 3.4.1.2. Reactivity Insertion Transients !
The rod withdrawal from zero power was not re- '
calculated since for all combinations of parameters, including the simul-taneous withdrawal of all rods in the core, the peak thermal power at-tained during the transient is always less than 100%. The rod with-drawal from rated power was calculated for two rod withdraws 1 rates:
1.4 =
10- Ak/k/s. which corresponds to the maximum single group with-drawal rate and 7.25 = 10' Ak/k/s, which corresponds to an all-rods withdrawal. The single group withdrawal gave a peak thermal power of 106%, and the all-rods withdrawal gave a peak thermal power of 103%. Both t of these values are well below the 112% design thermal power level; there-i fore, the 1.3 limit on DNBR is maintained for this transient.
The startup of an inactive loop was not consid- l ered in the analysis since the maximann thermal power achieved during the transient is much less than 100% and occurs after full flow is reached.
The rod drop accident results in an initial decrease in power which is followed by a return to 100% power. Since it has been shown previously that neither the withdrawal nor the drop of a single control element i will perturb :he flux shape sufficiently to exceed design conditions l 3-29 Babcock siWilcox l
- W=-
j E
at 1122, such occurrences still do not present any thermal problems.
The moderator dilution accident results in reactivity insertion rates that are very slow, and the accident is terminated by the high-pressure trip well before power reaches the 112% design thermal pcwer level.
Therefore, the 1.3 limit on. DNBR is maintained.
Ejection of the most reactive control rod at BOL with densified fuel was analyzed. The results show that this ac- E cident is no more severe than the equivalent accident presented in the Oconee SAR. The ejection of a 0.5% ak/k rod (maximum worth that occurs at rated power) results in a peak neutron power or' 320% and a peak thermal power of 119.5%. The comparable values presented in the SAR are 277 and 126%, respectively. A cc,aparison of the values shows that a rod ejection with densified fuel is less severe because of the in-creased time constant of the fuel heat transfer, which results in a slightly lower peak thermal power. Appendix B provides additional in-formation on this subject.
Secondary system accidents resulting in a power increase occur at or near end of life (EOL) when a highly negative moderator coef ficient exista. Since more DNB margin exists at EOL, these secondary accidents, such as a steam line break, are not expected to cause thernal limits that are more severe than those presented in the SAR. The SAR analysis of secondary system accidents, such as tube ruptures and loss of electric power, is unchanged since the thermal power remains the same or decreases during the transients and, there-fore, does not increase the potential for reaching design limits.
3.4.1.3. Loss of Coolant Flow The loss-of-coolant flow accident has been analyzed under initial conditions that represent the most conservative that can occur in the core with densified fuel. The case considered is a balanced power peak case with the power spike placed at the point of minimum DNBR. The other parameters normally considered in the l coastdown calculations remain unchanged from the FSAR values. Fig-ure 3.4-1 shows power, flow and the calculared average heat flux 3 for a four pump coastdown initiated from 102%. Figure 3.4-2 shows I
3-30 Babcock & Wilcox l
l I
[ l l
( the calculated DNBR and film coef ficient as a function of time. The gap conductance used for this calculation was 850 Btu /h-ft2-F. The f uel and cladding temperature is not shown since there was no variation in these parameters, bec:use the DNER for this accident did not go below the cri-terion value of 1.3. It is therefore concluded that no fuel damage will occur.
1.4.2. IDCA Analysis 3.4.2.1. Introduction Topical report BNi-IOO34 established the 8.55-ft2 split break at the pump discharge as the worst break location and size in
{
the FC system. This analysis assumed that the axial power peaked 3 feet f rom the bottom of the core and was shown to result in a higher cladding tenperature than that obtained when *.he power peaka at higher elevations in the core.
As shown in Table 3.3-2 the highest linear heat rate for equilibrium operation is on the order of 16 kW/f t. Higher peak heat rates can occur during maneuvering transients. The highest of these occurs during the 4-day transient at 5.5 feet from the bottom of the core. Inc peak linear heat rate was placed at the 5.5 foot level for this analysis. An exial peaking factor of 1.786 was used and the radial varied untti the limiting cladding temperature (2 2300F) was 1 reached. This procedure was accomplished to establish the reference case relative to which LOCA margin could be determined.
3.4.2.2. Initia1' Conditions The normal design basis transient is defined as a 100-30-100; transient consisting of operation ar. 100% power, reduction in power to 30% power, operation at 30% power for about 8 hours9.259259e-5 days <br />0.00222 hours <br />1.322751e-5 weeks <br />3.044e-6 months <br />, and a return to 100% power. The return to 100Z power is made at the time of l
raximum xenon, at a rate of power change of 10Z per minute.
The return to 100% power for a condition of maximum menon causes a rapid depletion of xenon, and control rods must be inserted with a resultant increase in the axial peaking factor. The highest values of power peaking occur at " xenon undershoot." which re-quires the greatest rod insertion. Further, the largest peaking factors occur near the beginning of ccre life.
Babcock a,Wilcox 3-31
~1 4
An 4x141 peakir.g factor of 1.186 was used in the LOCA analysis. This :- the largest value for the design basis transient and includes the usual factors for calculational uncertainty. The maxi-mum axial power occurs at 5.5 feet above the bottom of the core, or 0.5 -
I feet below the core midplane. Instead of imposing a power spike at the 1
5.5-foot elevation due to fuel densification, the local power spike was i
1 replaced by an equivalent radial multiplier over the entire fuel pin. E j This approach leads to a higher calculated peak cladding temperature of i approximately 10F. ,
When these adverse peaks occurred, it was further l assumed that the fuel changed to 96.5% of theoretical density with a decrease of 1.8% in pellet stack 1.eight. Rather than change the stack height in the THETAl-B model, the radial peaking factor was increased by 1.018 to provide an increase in linear heat rate. The diametral gap be-tween the fuel and the cladding was increamed with a corresponding in-crease in fuel temperatures.
3.4.2.3. Analysis Results from CRAFT were input into the THETA 1-B code to determine the peak cladding temperature. The THETA medel util-i ized 13 axial segments divided into eight 12-inch-long, two 16-inch-long. l l one 8-inch-long, and two 4-inch-long regions. The B&W carryout rate frac- l tion was used in the REFLOOD code to determine the FLECHT heat transfer coefficients.
3.4.2.4. Results The maximum linear beat rates accounting for fuel densification up to 96.5% TD for BOL and EOL were calculated according t l procedures similar to those found in part 4, Appendix A, of the AEC In- E E l I
1 terim Acceptance Criteria for Emergency Core Cooling Systems for Light 1
{
Water Power Reactors, dated June 29, 1971, as amended December 18, 1971.
The results are as follows:
1 1
3-31 Babcock 3.Wilcox l
i
.i f
I .
l Maximum linear Maxf-r= cladding heat rate kW/ft temp. F BOL 19.80 2291 EOL (40,000 MVd/c) 13.50 2280 I
The kW/f t values associated with the interia ae-ceptance criterion of 2300F peak clajding temperature were used to show the conservr.tism of the operating power distribution (Tables 3.3-1 through 3.3-8). The results demonstrate that ine operating power distributfors, with all densification effects included, yield if near heat rates that a-*
below the allowable peak values listed above.
I 1
i I
l l
l
\
1 l
1 l
\
3-33 Babcock & Wilcox l
I.0 __
, x :
- HEAI FLUI =
)N 9 -
_ . _ __ .. I.
E Flos ,
2 l .
' \ P a !
- . 7 5
- -
! L - ~
- .8 i \ ,
- R M ! [ !
! POWER E *7 -m 2 -
40 Y #8 W 2 m
- o.
p
.6 _-.
{p E
s 8E C 'I
.5 AT Tite ZERO Fl 2
g HEAT FLUK = 174.130 Stu/nt-Ft " !!
PNER = 1.02 X 2588 Ett 9*
.2 m
R N = 2988 lb/hr E5 g .4 - ( 39 g .=
8 $
x-P g .3 '
6 o 0 .2 .4 .6 .8 1.0 1.2 f.4 1.6 1.8 2.0 Time ($e.)
M M M M M M M M M M M M M M M M M M M
Figure 3.4-2. DNBR and Film Coefficient Vs Time for Four-Pimp Coastdown 1.78 g-5 NAXIMUM CLAD TEMPERATURE = 650*F l.74 -
= CONSTANT THROUGNOUT TR ANSIENT- 80.000
, /* '
/
l.70 / 70.000h
?
- / =
a i /
- /
[ 1.66 -
f 60.000 G y
= f va a f C E f
/
2 L E
- O R
/ : 2 1.62 - 50,000 u =
3
/ 3~
/ *
/
1.5s -
/
/
40 n00 z3
(
/
/
/
/
1.54 --~Y r i I 30.000 0 0.4 0.8 1.2 1.6 2.0 Time (Sec) l l
l 3-35 l
Babcock s.Wilcox 1
1
.. . . . s. ,
. r ,-. . .. - . . , . . . .. ,. ;. 3
(
3.5. Mechanical Analysis of Oconee 1 Fuel I
') . 5.1. Cladding Collapse Input sunetary
- 1. Pellet 03, in. 0.36998
- 2. Cladding ID, in. [ ]
- 3. Pellet density (batch 3) [ ]
4 **re-densif. int . press. [ ]
(cold), psia E g
- 5. Post-densif. int. press. [ }
(cold), psia l
- 6. Cladding ovality, in. [ ] =
- 7. Cladding wall thkas, in. 0.0262 l
- 8. Rod time-temp-press. history:
Cladding Internal press.,
Time, h temp, F psia 0-600 639 600-1,200 6'l9 1,200-2,400 639 3
2,400-4,800 631 5
4,800-6,000 642 6,000-6,840 g '
634 I 6,840-7,440 638 E 7,440-8,640 642 8,640-9,840 639 9,840-11,040 635 11,040-12,240 626 12,240-14,280 14,280-14,880 624 623 g
14,880-16,080 621 3
16.080-17,280 621 17,280-18,480 621 18,480-19,680 622 19,680-21,720 621 21,720-22,320 622
- 9. Post-deasif. pellet OD in. [ ]
- 10. Fast flux level 6.94 = 1013 Results Predicted time-to-collapse: [ ] h when fission ga,s release is ignored.
3-36 Babcock a.Wilcox l
3.5.2. Cladding Stress
_ Input Su ma rv
- 1. Min intepress. (cold) after densif., psia
[ ]
- 2. Cladding temp. As described in BAW-10055 Rev 1 Results
- 1. Table 3.5-1 lists cladding maximum stress calculated at various times in life. In no case does stress exceed yield.
- 2. Cumulative fatigue damage after three cycles is less than 0.9.
3.5-3. Fuel Pellet Irradiation Swelling Input Surzaarv
- 1. Density (mean + 2s local) [ ]
- 2. Initial pellet dia, in.
O.36998
- 3. Cladding ID, in.
[ ]
- 4. Max local three-cycle burnup, 42,000 mwd /mtU
- 5. Pellet dia af ter densif., in. [ ]
- 6. Peak local heat rate, kW/f t 22 Results Circumft;rential plastic strain less than 1% after three cycles.
3.5.4. Revision Input to Creep-Collapse Calculations The application of the creep model to fuel rods in a spec-ific reactor was performed using input from the worst-case rod. The method for selecting this input is described as follows:
1.
An initial gap between the pellet and the claddir.g is calcu- l laced on the basis of as-built dimensions. The mean as-built pellet OD is combined with the mean as-built cladding ID minus two standard de-viations to determine a minimun BOL gap. These dimensions result in fuel pellets with low thermal expansion and fission-gas release, and therefore a rod with a high pressure differential.
3-37 Babcock s.WilCox
- 2. Initial pellet density is the mean as-built density for fuel assemblies operated for three cycles. When the pellet densifies to 96.5%, a change in stack and pellet dimensions produces a change in rod internal pressure. The use of the minimum diametral gap in the fore-going step produces a minimum gap volume that contains backfill gas.
The increase in void volume due to pellet shrinkage will in this case have the greatest effect on BOL pressure.
- 3. The internal rod pressure before densification is assumed to be the lowest allowable pressure for backfill gas. (All B&W fuel will be prepressurized.)
- 4. Rod length, end cap dimensions, and internal spacer arrange-ment and dimensions are assumed to be at values that will minimize the rod internal vold volume.
- 5. Comoining steps 1 through 4, the BOL cold rod internal pres-sure is calculated by the equation BOL press. = backfill press. m Rod void volume after densification The increase in void volume is based on pellet dimensional changes cal-culated in accordance with the AEC report of November 14, 1972.
- 6. Cladding ovality at BOL is the mean as-built ovality plus two standard deviations.
- 7. The wall thickness at BOL is the mean as-built thickness minus E
two standard deviations. 5
- 8. For each plant, the radial x local peaking histories of three-g cycle assemblies are examined to determine a conservative worst-case rod 5 !
peaking history. For each time increment (varying from 25 to 100 days),
the maximum peak for any three-cycle assembly is assumed to be the peak i
for that increment. Thus, the final assembly peaking history represents <
a combination of peaks from many assemblies. Rod peaking history is ob-tained by applying a rod-peak / assembly peak ratio [ ] to the assembly history. A [ ] uncertainty factor is also applied. The resultant peaking l versus time history is used as input to the cladding temperature model. l After relating the history to burnup, it is also used as input to the 3-38 Babcock a.Wilcox I
l
TAtY code (along with the rod geometry and pressure described above) to compute the rod internal pressure versus time curve.
- 9. The cladding OD temperature along a gap in the pellet stack drops to the coolant temperature within about 1/8 inch. The temperature increase across the cladding wall has been calculated to be 3F due to radiating f roen pellets at the boundary. Therefore, the assumed maxt==
cladding temperature for any gap is the coolant temperature correspond-Ing to the peaks described in step 8 plus 3F. .
I l
l l
i l
l
)
l 3-39 Babcock & Wilcox
, a i t
Table 3. 5-1. Cladding Circumferential Stress Bending + Total Yield I lt imate Ext pres s. . Int press. , Cladding membrane Case psia psia 'emp,F stress, strength, strength.
stress, psi psi psi psi 1 2200 460 532 2500
-22500 -22500 48000 57000 460 5 32 -27600 -27600 48000 57000 2 2200 580 650 -20800 -20800 45000 2500 580 5000C 650 -25700 -25700 45000 50000 3 2200 600 650 -20500 2500 -20500 45000 50000 600 650 -25400 -25400 45000 50000 Y 4 2200 580 723
$ -21000 -24900 42000 44000 2500 580 723 -26100 -30000 42000 44000 5 2200 600 733 -20700 -25100 41500 43500 2500 600 733 -25700 -30200 41500 43500 6 2200 460 532 -22500 -22500 48000 57000 2500 460 532 -27600 -27600 48000 57000 9 2200 580 704 -20900 -23800 43000 46000 2500 580 704 -26000 -28900 43000 46000 10 2200 600 711 -20600 -23400 43000 46000 2500 600 711 -25600 -28900 43000 46000 11 2200 460 535 -22500 -22800 48000 57000 g 2500 460 535 -27600 -27800 48000 57000 K 12 1725 400 425 -16300 -16300 50000 62500 8
x I
w M M M M M M M M M M M M M M M
I l
l a
l APPENDIY A Summary of Statistical Analysis l for Oconee 1 As-Built Data A-1 Babcock 8.Wilcox 1
1
I In keeping with the guidelines given in " Technical Report on Dennification of Light Water Reactor Fuels," November 14. 1972, as-built data on the UOy feel pellets have been utilized in the densifi-cation analysis. The results for fuel pellet diameter and density are given in Table A-1.
The standard deviations for the local pellet diameters and densi-tles (o ) were pooled to give a representative value for the core.
l Table A-1. Summary of Statistical Analysis for Oconee 1 As-Built Fuel Pellet Data !
i I !
I I
I I
I i
I I
l i
I l
l I
A-2 Babcock & Wilcox l '
l l
l I .. ._ _ . . _ _ _ . _ _ _ . -___ _ _ -
l l
l I
l l
I l
APPENDIX B Answers to Questions Contained in Letter to Duke Power from R.C. DeYoung dated March 14, 1973 n-1 Babcock & Wilcox
I The AEC issued a list of questions to Duke Power Co: pany on March 14, 197 5 as a culmination of the Staff review of ef fects of postulated fuel densificat ion on the design and operation of the Oconee 1 nuclear station. These questions are addressed in this appendix to facilitate g
ref erral to the answers. E Question I Provide the values for the following physical properties and dicen-sionn of the Oconee Unit I fuel pins:
- 1. Fuel pellet length, diameter, dished end volume and chanfer volume.
- 2. Fuel pellet density.
- 3. Clad inside and outside diameter, initial ovality, and wall thickness.
For each of these parameters, provide the nominal value, the speci-ficd value and tolerances, and the average value (as built) with standard deviation as well as maximum and minimum measured values.
Describe the acthod of measuring that was used, the frequency of neasuring samples, and the production step at which the ceasurements are perforned in the production process of the fuel assembly. If any of these parameters is measured at different times in the production pro-cess, e.g., at the fuel pellet manufacturer cod at B&W, with identical i
or dif ferent measuring techniques compare the results and discuss any '
differences.
I Response l 1
l l 1
1 I
1 I
B-2 Babcock 3. Wilcox
1 l
1.2. Fuel Pellet As-Built Data ' ~ ~
In general, the fuel pellet vendor is the source of information.
All vendor measurements are taken at the end of the pelletizing process.
l_. 2.1. Acceptance Criteria The f uel pellet vendor is required to sample every produc-tion lot for pellet length, average diameter and density based on a variables sar:pling plan. The acceptable quality level (AQL) lot toler-ante percent defective (LTPD) consumer's risk (CR) and producers risk (PR) .are speci fied. These, in conjunction with the tolerance limits, detereine the sample size and the acceptability of the pellet lot.
The fincl pellet vendor is also required to sample every pr odut- t ion lot for dish dimensions based on an attributes sampling based on mpecified AQL. LTPD CR. and PR. Again these parameters and the tolerances determine the acceptance criteria for the pellet lot.
1.2.2. Comercial Nuclear Fuel Plant Overcheck B&W's Commercial Nuclear Fuel Plant performed over-checks on pellet length, average diameters, density and dish dimensions utili-zing attributes saepling. These results are compared with vendor quality control data before acceptance for fuel assembly production.
_1. 2. 3. Method of Measurewnt The vendors quality control program requires:
B-3 Babcock a. Wilcox
l I
- 1. Pellet length will be measured with a disc micrometer I
with a rtiuinum graduation of 0.001 in.
- 2. Pellet average diameter is measured at each end and at the midpoint with a micrometer of minimum graduation of 0.001 in. The three measurements taken along the pellet length are spaced 120* apart radially. Average diameter <
is calculated by adding the two end measurements to twice the midpoint measurene :t and dividing by four.
- 3. Pellet dish depth and diameter are measured by attributes- I type sarpling employing a bench-type indicating dish .
depth gage with a minimum graduation of 0.0001 in. and 5 an optical comparator readable to 0.001 in. for diameter.
4 Pellet density is calculated f rom average diameter and length and weight. Weight is measured to a graduation of 0.01 gran. A correction factor is applied for the dished ends.
1.3. Details of Inspection of Occoce 1 Claddine The cladding vendor was required to perform 100% inspection for ID, OD, and minimum wall thickness dimensions with a scanning instrument l rovering at least 720*/ft. Ilie specification requires all traces to l l
be within the given tolerance limits before the claddf.ng is acceptable for fuel tissembly production.
Available vendor treces on Occoce 1 cladding include 300 traces 1
from 6 lots. In rddition B&W's Connereail Nuclear Fuel Plant performs 1 an ID and OD overcheck at a frequency of 10 tubes /1000. ID was measured =
ct both ends, and OD was measured at the ends and middle of the cladding tubes.
Ducstton 2 In order to assess the B&W evaluation model, TAFY, for stored I energy, f uel pellet to clad gap conductance, and clad temperature of a l
fuct pin for Oconee t'nf t 1 a detailed description of the following g
e items is required. Where applicable, equations and empirical formula-tions should be provided.
34 Babcock a,Wilcox l
lI l
I 2 .1. Question 2(a) .
We amount and ccesposition of the sorbed gasen assumed to be pre-sent in the fuel including the analytical riethods used to describe the release rate of the sorbed gases.
Response
ne content of the sorbed gas in the fuel is assumed to be air and is considered to be completely released into the gap at beginning of life. ne sorbed gas is accounted for in the gap conductance deter-ntn.ition. TI,ie initial concentration of sorbed gas is 0.02 cc/gm 2.2. Question 2(b)
The auount and content of the gas in the gap between fuel pellet and clad.
Response
The backfill gas is 100% He. Oconee 1 fuel rods were prepressurized to ( )
I g 2 . 1. Questton 2(c)
Descript ton of gas mixture conductivity model for the gas in the I f uel pellet-clad gap, and the thermal expansion model for the fuel pellet and the clad. Discuss how the fuel cracking is treated.
Response
l The gas conductivity mixture model used in TAFY is called GSMIX l
and it is a sub routine of the main code. 'Ihe CSMIX sub routine vas obtained from Argonne National Laboratory and detailed explanation can he found in the followin,q reference: Holmes, J.T. and Baerns, M.G.,
" Evaluation of Physical Ptoperties of Cases and Multicomponent Gas Mix-tures". Report number ANL-6951, Argonne National 1.aboratory, November 1964 The TAFY-3 loput tc, CSMIX is the gn-soles of sorbed gas, the gn-notes of residual air, the gm-moles of He backfill, and the ga-moles of f18810n gas released. The mass of sorbed gas, residual air, and He backfill gas will remain constant with lifetime, however, the relative concentration of the mixture will change with lifetime since the initial B-5 Babcock A.Wilcox
l I
mixture is being diluted with fission gas throughout burnup. The con-statuents tha make up the fission gas are 15.uZ krypton and 84.2% xenon.
Thermal Expansicn of Fuel and Clad
!.un l Ti.e thermal expansion of the fuel in the axial and radial directions are considered separately in TAFT-3.
The axial or longitudinal fuel swelling is a function of the mean sec-tional temperature and the mean coef ficient of expansion. The mean sec-tion temperature is calculated by the following equation:
T = nax - T, 2
1 l where:
T =
maximum centerline fuel temperature for the axial segment.
T, =
fuel pellet surface temperature. !
(
The mean coef ficient of axial expansion (0,) is temperature dependent and I is calculated as follows:
if T, 1 1950*F a, = 5.699 x 10 -6 (2) l and if f, > 1950*F l l a, =
4.005 x 10~0 + 8.575 x 10-10 y, (3)
The longitudinal differential expansion for each segment can be calgulated
(
cnd summed to attain the total stack differentiel expansion using the following equation: g 3
l s-6 Babcock s. Wilcox I )
I .
i I
6 i
n AL g -
I o (T8 -
T) L i=1 *1 1 0 1 (4)
T, =
initial ambient temperature L =
initial segment length 9 e bient conditions g
AL g -
total axial differential expansion of fuel due to thermal effects.
n -
total number of axial segments The pellet radial thermal expansien is also a function of a mean coef-ficient of Ifnear expansion as given in equations 2 and 3. The equiva-lent temperature (Tr) is based on the following equation.
i=1 .
I 1~1 ( i-1 i}
2 r* __
l l
(R max (5)
Rmin) where:
Tg -
fuel temperature at i th radial segment
=
Tg ,g fuel temperature at 1-1 radial segment Rg _g= radius of 1-1 segment Rg =
radius of i th segment R =
radit5s of pellet 0.D.
R =
- " radius of central void. If central void doesn't exist R = 0.0 n =
total number of radial segments l
B-7 Babcock & Wilcox l
)
~
I The total radial expansion of the fuel due to thermal growth can be written as:
M g = o r
(R - R,g ) dr ~
o} (}
where: o = o, 8 tea af the values for longitudinal and radial thermal growth will vary taroughout the lifetime of the fuel since the fuel temperature will vary with lifetime, clad The thermal expansion of the clad in the radial and longitudinal direction are considered separately in TAFY-3.
The longitudinal clad growth is a function of the mean clad temperature and the mean coefficient of linear expansion which is:
l a = 3.1346 x 10-6 + 3.878 x 10-10 (7)
I where:
u = coefficient of mean expansion l l
T = numerical average of the tempera- l ture of the cladding's inner and outer surfaces.
t i
The total axial elongation due to thermal expansion is given by:
I
=.
-Ane s - s r.> <e>
I I
B-8 Babcock a.Wilcox G
w -
l where:
T = initial ambient temperature l
L = initial segment length 0 ambient conditions tal. = total axial elongation due to thermal expansion The clad radial expansion due to thernal effects is assimed to be ade-quately described by a " thin cylinder" approximation. For thermal ex-pansion the inner cladding radius is computed by:
Ric =
Ric, 1+o (5,- T ,)
d=re:
.c Ric, =
inner clad radius @ ambient conditions a
=
mean coefficient of linear expansion (see equation 8)
T, =
ambient temperature T =
numerical average of the temperature of the cladding's inner and outer surfaces.
The value of cladding axial and radial thermal ex-pansion will vary directly as the mean cladding temperature varies with lifetime.
Fuel pellet cracking is not treated explicitly by the BW fuel pin tem-perature and pressure code TAFY. Fuel pellet cracking, however, is implicit in the empirical data used to formulate the various models con-tained within TA7Y.
2.4 Questton 2(d)
I A listing of input values used for the TAFY code, including fuel and clad surface roughness and the fuel pin plenem volume.
B-9 Babcock & WilcoK
1 1
Response
I The AEC codel used to predict the fuel pellet geo:netry and fuel pin in-ternal pres.sure utilizes the following equations to determine the densi-fled diameter.
Dg = D - (o{ - o{ + 2o) p i (g) 3 where:
- D =
g final densified diameter of the fuel (0.36523)
Dg =
nominal value of fuel pellet diameter (0.36998) c 7
=
final density at which densification is complete (96.5% T.D.)
bg *
[ ] as built data 20 =
[ ] T.D.
The densified stack length is given by:
If
- Lg - (E r - o)L g g (2) 2 where:
E r
=
final densified stack length, (141.8 in.)
Lg =
initial nominal stack length from as-built data, (143.97 in) (143.97)
The total initial void volume is calculated by the following equation.
Th= first term represents the volume in the gap, the second term re-presents the total plenum volume, and the last term is the dish void volume. "'
,, =
t,.
(< -+ . 1 i. t,. g i i (n I
4 4 ,
g B-10 Babcock & Mca
l
)
where:
V = the initial void volume present in g ,
the fuel rod [ ]
D = the statistical diameter of the clad I.D. [ ]
The final void volure is given by: ,
i V
g
=
Lg = ,(D -D)+[
g ]
4 2 2 l
+ AL w (D ) + LvD p g[ ]
(4) 4 4 .
where:
l Yg =
final total void volume after densifi-cation [ ]
AL = L g -L g l
i I
i Knowing the initial pre pressurization level and the initial and final void volumes, the final internal pin pressure can be calculated from:
i l
P = Pg V, g ($)
Vg i
I where:
(
Pg = final internal pin pressure after densifi-cation [ ]
Pg = initial total pre-pressurization level i
. I l The remaining terns unaffected by densification which are input to TAFY-3 are given below:
1 l
B-11 l Babcock & Wilcox i
l I
l
.I
I Film coefficient =
5000 Btu /hr-f t *F
=
Coolant Tesperature- 582*F Sorbed Gas = 0.02 cc/gm
-6 Roughness = { } (fuel / clad) RHR 10 in.
2.5. Question 2(e)
A listing of the following parameters calculated with TAFY; hot gap size, fuel pellet diameter, conductivity of gas mixture, tem- '
perature jump distance, gap conductance and the contribution of each of the additive terms in the gap conductance. The information should be ,
provided as a function of linear heat generation rate (kW/f t) and as a funct ton of fuel burnup.
Response
The following tables give the requested parameters as calculated with TAFY for densified and undensified fuel.
I Table 2-1. Without Densification [ ]
Nominal Cold Diametrical Can) 1.inear Conductivity Temperature
!! cat Of Jump B Rate Burnup Hot Cap Hot Fuel Pellet Dia. Gas Distance 5 (Ew/ft) (MVD/KfU) (Mils) (in) (Bru/hr-ft *F) _ (in) 5.6 BOL 0.127733 0.3317x10[5 5.6 End cycle 1 0.095218 0.2011x10 20.1 BOL O.139597 0.2029x10[5 20.1 End cycle 1 0.104022 0.3795x10 Table 2-2. With Densification [ }
Diametrical Cao) 5.6 Bol 0.127778 0.5107xib[5 5.6 End cycle 1 0.106418 0.3829x10 20.1 BOL 0.148057 20.I End cyc1e 1 0.125849 0.6382x10[f 0.4793x10 I
B-12 babcock & Wi!COX
/5L J l
l Provided in the two tables below are the individual contrubutions of each additive tern in the conductance equation for no densification ef fects and with densification effect. These rest:lts pe:tain to the Oconee 1 l
P ant and thev were obtained from TAFY-3.
g ,
Table 2-3 Without Densification IIcat h H Rate total E#' "
=
('w/ft) radiation Eurnuo (Bru/hr-ft *F) (Btu /hr-ft
- F) (Stu/hr-f t *F) (B tu /hr-f t2.p) 5.6 ROL 1364.0 5.6 1302.0 60.0 2.0 Emi cycle 1 1128.0 1061.0 65.0 2.0 20.I 1911. 3505.0 20.1 3347.0 155.0 3. 0' End cycle 1 3235.0 3037.0 195.0 3.0
_ Table 2-4 With Densification 5.6 BOL 798.0
%.h 758.0 38.0 2.0 End cycle 1 700.0 660.0 37.0 3.0
'O.1 not. 1189.0 1147.0 37.0 5.0
.'O . ! End cycle I 1114.0 1069.0 40.0 5.0 1.6 Unestion 2(f) l'rovide comparf non of TAFY vith fuel perforniance data. Of in-terest is comparison with gaps of approximately the size assumed after denstfIcation.
Response
The TAFY computer code was developed as an accurate model for the calculation of fuel temperatures within a cylittirical fuel rod. In this endeavor a realistic approach has been taken with regard to the indivd-ual components of this model. Where little data was available for a particular area of concern a conservative approach was taken. Thus, the complete model should be capable of predicting fuel temperatures in a realistic but generally conservative manner. The ultimate evalua-tion of the complete model should be its ability to reproduce fuel per-formance data available in the published literature.
rAFY has been compared to published data with respect to two per-formance criteria: 1) measured fuel centerline temperatures 2) measured B-13 Babcock 8.Wilcox
3 s I I
or " inferred" gap conductances.
Tabular and graphical comparisons of TAFY with data from four sources are presentec.:
- 1. " Microscopic. Autoradiographic and Fuel / Sheath Heat Transfer Studies on l'O Fuel Elements". A.S. Bain, Chalk River, June
- 1966.AECL-25k8
- 2. "In-Pile Determination of UO2 Thermal Conductivity Density Ef fects and Gap Conductance", G. Kjaerheim and E. Rolstad.
Decceber 1967. HPR-80.
- 3. "In-Pile Measurement of UO2 Thermal Conductivity", M.G. Balfour, J.A. Christensen and H.H. Ferrari, March 1966, WCAP-2923.
- 4. "Densification Considerations in BWR Fuel Design and Perforaance".
D.C. Ditmore and R.B. Elkins, December 1972, NED*A-10735.
For each investigation, the pertinent physical parameters have been listed followed by the tabulation of experimental results and the cor-responding TAFY prediction. Also included is a separate figure for each reference graphically presenting the tabulated results.
The results of measured centerline temperatures as reported by i Cohen, I.ustman, and Eichenberg (WAPD-228) were not selected for this comparison due to the extremely low heat rates involved (less than 6 kw/ft). Any comparison with published data in this power range would he of limited value in light of present peak heat rates.
I e
The fuel-to-clad diametral gaps considered in these investi2ations
. ranged fron 0 to 33 mils, which brackets the gaps encountered in the -
SW fuel products, with and without the assumption of densification.
Due to the inherent uncertainties in " measuring" gap conductance the, measured fuel centerline temperatures of HPR-80 and WCAP-7'23 were g sclerted as the most appropriate indication of TAFY's ability to re- E produce published data. The composite of the TAFY comparison with these two sources is shown in Figure 2-1. As seen in this Figure TAFY has been shown to be capable of predicting seasured centerline temperatures in a realistic but generally conservative manner.
I B-14 Babcock a,Wilcox
l l
l Figure 2-1. TAFY Courparison With Published Data O HPR-80 0 WCAP-2923
, o'- 5000 u;
E O
5 8
$ 14000 -
U l
W o h O O
l U 3000 - o 5
o a .
O U
o O 5 2000 - O O
t e a '
O 99 9
1000 -
0 f f f f 1000 2000 3000 4000 5000 MEASURED CEllTERLINE TEliP 'F B-13 Babcock s.Wilcox
I
- 1. Comparison of TAFY vith published data source 1. " Microscopic.
Autoradiographic and Fuel / Sheath Heat Transfer Stuides on UO 2 Fuel Elements" TAFY predictions of gap conductance are compared to gap conduc- l I
tances inferred from measured melt radii. The experiment represents a vide range of fuel-to-clad diametral gaps. fuel clad interface atmospheres and clad materials. The two tests were the Hydraulic Rabbit Test (cool- !
1 ant temperature = 50F and pressure - 100 psia) and the X-23201. S-23303 )
X-23401 Locp Tests (coolant temperature = 482F and pressure - 1420 psia) .
The specific conditions for the various elements are tabulated on the following page for the two tests.
I I l 1
i I
I
( I r
i t l l
I B-16 babcock 8 WilCOX 1
l l
l l
l 1
l l
l l
Table 2-5. Test Data rt.D'F5T CIAD NO. TILL U-235 IVEL PATERIAL CAS kT. Z DENSITY CLAD I.D. Cl>D O.D. LINCTH TEST Z T.D. IN. IN. IN.
s-4 r-14 APPARATYS S.s. AR. 1.0 97.8 .669 .721 F t-31 2.76 PHASE 13 S.S. AR. 1.5 HYDML'LIC RAB8IT 97.8 .669 .721 1 F.-4 5 : 1:-47 S.S. ^
2.76
- AR. 2.2 97.35 .669 .721 F-16, r.~48 2.76 a S.S. HE. 2.2 97.35 .669 .721 I-49 E-59 2.76 =
21-4 AA. 2.75 97.2 .669
- -0 .721 2.76 5 10 ZR-2 a
AR. 2.0 96.0 .669 .719 511, 512 3.01 x-23400 2R-2 VAC. 2.0 96.0 .669 ETDRAtlLIC RABBIT a-15. ft-16 .719 3.01 =
ZR-2 AR. 2.75 96.0 .669
%C. YD .722 2.76 5.5 VAC. ppa 37, go 3.99 95.8 .666 VI. VF .686 2.76 S.S. AR. 3.99 95.8 .666 W. YH .686 2.76 -
S.S. VAC. 4.82 %.7 .666 VI. VJ .686 2.76 a S.S. AR. 4.82 96.7 .666 FAO .686 2.76 a 2R-2 AR. 1.5 95.8 .720 X-23301 FAT .794 6.0 y-23303 AR-2 AR. 1.5 95.9 .720 X-234n1 t'Ar ,794 6.0 a 29-2 AR. 1.5 95.5 .720 FAa* .794 6.0 a 2a-2 AR. 1.5 95.5 .720 FAs .794 6.0 a 7a-2 AR. 1.5 95.7 .720 FAI .794 6.0 a ZR-2 AR. g.$ 95.7 .720 FA5 .794 6.0 a ZR-2 AR. 1.5 96.0 .720 FAY .794 6.n a ZR-2 AR. I.5 96.1 .720 FA3 .794 6.0 a 7R-2 AR. 1.5 95.7 .720 FAZ .794 6.0 a 7R-2 AR. 1.5 %.1 .720 FAE .794 6.0 a ZR-2 AR. 1.5 95.9 .7397 FAF .800 6.0 a ZR-2 VAC. 1.5 96 9 .7397 .800 6.0 a FAY ZR-2 1.5 AR. %.2 .7401 FAX
.800 6.0 -
1R-2 1.5 VAC. 96.2 *7397 .800 6.0 a B-17 OCk & MCOX
_ _ _ __ _ I
Table 2-6. TAFY Comparison With Pha w 13 Hydraulic Rabbit Test Results ELEMDfT NO. DIA. GAP IIEAT RATE U Ar w.u o A.wt.
TAET EXPERIM1.WT (BTU!HR-IT 2 .y) l m
l (IN) (KW/FT) (BTU /HR-FT 2 F)
E-4 .033 15.4 187 125 E-12 .033 15.1 170 117 E-17 .020 21.6 355 248 E-18 .026 21.3 263 236 g E-20 033 19.8 199 185 E-21 .020 21.0 335 233 E E-22 .033 20.3 202 199 E-23 .020 21.3 345 254 E-24 .033 20.6 205 199 E-26 .026 20.1 245 217 E-28 .033 19.6 197 188 E-29 .026 20.1 245 217 E-31 .026 19.7 238 203 .
E-34 .026 25.0 327 336 E-36 .033 26.6 265 308 E-37 020 26.1 499 444 E-38 .033 25.5 253 294 E-39 020 26.3 507 520 E-41 .020 27.7 562 493 E-42 .026 27.3 377 402 E-43 .013 28.5 1055 749 g E-44 .033 26.7 266 314 g E-45 .026 25.5 337 395 E-46 .026 25.4 392 486 E-47 .026 25.4 266 361 E-48 .026 25.5 394 486 E-40 .006 35.9 1667 2325 E-52 .006 35.6 1655 2149 E-53 .020 33.7 862 1075 E-54 .020 33.7 862 1013 E-55 020 34.0 883 978 g E-56 020 33 876 1145 3 E-57 .006 34.8 1655 1762 E-58 .020 33.1 822. 942 E-59 006 35.3 1655 1762 I
B-18 Babcock a.Wilcox
l Table 2-7
= TAFY Comparison with X-23' A 1. X-23303, X-23401 Loop Test Results I LI.EMENT NO.
DIA. CAP IN.
HEAT RATE T> R (BOL) TAFY (8000 MWD /MT) 2 .F EXPERIMENT KW/FT BTU /HR-FT BTU /ER-FT2 -F FAP 015 20.5 580 477 481 FAW .029 20.5 254 207 276 FAR .015 20.5 5 80 477 490 i FAX FAS
.029 015 18.6 17.6 228 456 183 373 257 370 FAY 029 15.6 198 149 231 l FAN .015 22.0 658 541 528
] FAZ .029 22.0 281 231 317 FAO .015 17.4 449 366 352 FAT .029 18.6 228 183 234 Table 2-8. TAFT Ccaparison with X-23400 Hydraulic Rabbit Test Results ET.N' DIA. GAP IIEAI RATE TAFT EXPERI!ENT NO. IN. KU/FT 8TU/HR-FT2 -F LTU/HR-FT2 .7 H-9 .004 35.9 1584 3259 M-10 .004 33.0 1571 3523 M-11 .004 33.0 1651 7222 H-12 .004 33.0 16!.2 4404 M-15 .004 39.5 1591 2373 H-16 .004 39.5 1578 2114 M-17 .004 39.5 1669 3347 H-18 .004 39.5 1656 21490 Table 2-9. TAFY Cceparison with Phase 10 Hydraulic Rabbit Test Results ELEMENT DI . dAP' liEAT RATE TAFT EXPERIMENT
_ _t:0. IN. KW/FT 2 BTU /HR-FT -F BTV/HR-FT2 _y
- SK 003 33 2220 2290 l SS 003 33 2215 2290 VC 0039 33.8 1870 2079 VD .0019 33.8 1870 2079 VE .0039 33.6 1812 1585 VF .0039 33.5 1810 1585 VG .0039 37.2 1859 2889 VH 0039 37.0 1856 2008 VI .0039 38.0 1794 VJ 1585 k
.0039 37.8 1791 1585
{
B-19 Babcock & Wilcox s
l I
I Figure 2-2. TAFY Comparison With Published Data I
10000 as E
DATA FRC;4: BAIN, A.S., " MICROSCOPIC, AUTORADl0 GRAPHIC AND FUEL /SHEAlH l
HEAT TRt.NSFER STUDIES ON U0 2 FUEL ELEMENTS", JUNE 1966 AECL -2588
~
U g Oo LE @ &O t ;-
se o
O l I m o 1000 =
g
.a ? 6*
N> g UN g S" O o
')
0 db o o
O O I
100 1000 10000 I
GAP CONDUCTANCE, Btu /hr-ft 2 _.y EXPERIMENTAL
-l I
n-20 Babcock a.Wilcox
~_
- 2. Comparison of TAFY with published data source 2. "In-Pile Deter-mination of UOy Thermal Conductivity. Density Effects and Cap Conductance I ,
Table 2-10. Test Data
- Fuel Fod Thermocouple Diametral Cap Density (IN) (! T D i
!!8A 11TF-1 .00189 98 -
l IIbB 11TF-2 1
.00205 96 l
!!PC 11TF-3 !
.00650 96 IICA 21TF-1 .00185 98 IILC 21TF-2 .00661 96 'I frCD 21TF-3 .00228 96 t
i
)
The followinJt conditions were constant for all the fuel rod samples:
Pelle t stack ht, in.
67.5 Clad material Zr-2 Clad surface roughness. uin. 157.5 Pellet roughness. uin.
49.2
- Clad surface temperature. F 473 B-21 Babcock & Wilcox i
l
I Table 2-11. TAFY Coerparison With Halden Test Results I
l Fuel Rod HBA - Thermocouple 11TF-1
! HEAT RATE _ CFNTET11NE TO(PEPATURE CAP CONDUCTANCE TAIT HA1. DEN TAFY HALDEN Kv/TT (F) (F) B'20/HR-FT2 -F BTU / fir-FT2-F l
2.90 870 734 710 1370 3.81 980 824 750 1430 4.72 1110 932 820 1!00 8.23 1640 1418 1060 1780 8.99 1750 1499 1130 1860
'> . 30 1810 1580 1140 1870 l 9.45 1880 1580 1150 1880 v.45 1880 1544 1150 1880 10.06 1950 1652 1210 1920 11.60 2260 1994 1350 2060 13.47 2590 2228 1500 2150 15.54 3030 2516 1500 2200
( Fuel Rod HBB - Therzocouple 11TF-2 I
l 0. 76 560 518 -- --
2.90 860 752 670 1330 3.66 960 842 720 1380 4.63 1100 968 770 1450 5.79 1290 1168 850 1540 8.08 1700 1436 3020 1730 8.84 1820 1544 1080 1790 9.20 1890 1670 1110 1820 9.33 1900 1634 1120 1830 9.30 1900 1688 1120 1830 9.75 2000 1724 1160 1870 11.58 2350 20]2 1320 2020 ,
13.26 2710 2246 1460 2160 t 15.39 3180 2624 1500 2200 ruel Rod HBC - Thermocouple 11TF-_3 0.76 620 518 --- -
2.59 1020 788 250 590 3.66 1250 932 280 610 4.63 1450 1098 310 630 5.79 1730 131 1 330 650 7.99 2210 1670 400 710 3.75 2380 1796 430 740 E m
8.99 2430 1904 440 750 9.20 2490 1850 440 750 9.70 2490 1904 440 750 9.4 ?570 1940 450 770 31.34 2;F0 2264 510 840 1/. 94 3:34 0 2150 720 1080 l
B-22 Babcock a.Wilcox
Table 2-11. TAFY Coeparison With Halden Test Results (Cont 'd)
_ Fuel Rod HCA - Thermocouple 21TF-1 HEAT RATE CD;TERLINE TEMPERATL*RE CAP CONDUCTANCE TAFY HALDEN TAFY HALDDI (KW/FT) (F) (F) BTU /HR-FT2-F BTU /HR-FT2 -F 4.72 11D0 950 84 0 1530 5.49 1190 1040 890 1600 5.94 1270 1112 920 1630 7.22 1460 1274 1000 1720 8.84 1700 1526 1130 1850 9.20 1750 1562 1160 1880 10.06 1900 1670 1220 1950 10.97 -
2100 1814 1300 2020 11.13 2120 1868 1310 2030 12.30 2380 2111 1420 2140 13.56 2630 2264 1480 2190 14.69 2950 2516 1480 2180 Fuel Rod HCC - Thermocouple 21TF-2 4.48 1450 1049 280 640 5.33 1620 1184 310 670 5.70 1710 1292 320 680 7.16 2050 1562 360 720 8.53 2320 1796 410 770 8.84 2400 1841 420 770 9.37 2530 1940 430 800 9.60 2570 1940 440 810 10.52 2740 2102 470 830 10.8? 2810 2156 480 850 12.10 3060 2372 520 890 12.50 3120 2466 540 910 13.11 3240 2534 570 930 13.56 3310 2543 600 960 14.17 3400 2606 630 990 l Fuel Rod HCD - Thermocouple 21TF-3 4.57 1070 887 750 1370 5.30 1200 1C46 800 1430 1
5.79 1280 1148 830 1460 l
7.01 1520 1328 910 1550 1
7.32 1570 1400 930 1580 C.69 1850 1562 1030 1680 S.99 1900 1634 ?.360 1710 9.66 2010 1706 1110 1770 9.75 2050 1724 1120 1780 10.73 2240 1841 1180. 1850 10.97 2270 1904 1200 1880 12.05 2530 2138 1300 1990 12./. 5 2609 2246 1340 2020 13.32 2740 2318 1390 2070 13.73 2C40 2444 1430 2090 14.33 2950 2534 1470 2090 Babcock & %Vilcox B-23
1 I
Figure 2-3. TAFY Cotaparison With Published Data 1 I U
DATA FRON: KJAERHEIN AND ROLSTAD "IN PILE DETERMINATION 3 0F 002 THERMAL CONDUCTIVITY. OENSITY EFFECTS o E AND GAP CONDUCTANCE" HPR -80 .
. l
~
3000 l .
3 5
o o
o oo
. I 4
9 y# l a
w a
r*o *. - l oo u
2000 .
O@ l l
_J
$ & o%*
l* O O O O G m
- l
! a a." i
'[ 1000 -
g O '
3 0
~
! I 1
s I I
O 0
1000 2000 3000 l
MEASURED FUEL CENTERLINE TEMP *F g
B-24 Babcock & WilCOX l
i
l l
l l
- 3. Comparison of TAFY with pelished data source 3, '*In-Pile Measure-ment of CO2 Thermal Conductivity" Table 2-12. Test Data gpsule 1 Average heat rate 22 W/IT Fuel Density 95% T.D.
Enrich:ent .64 w/o U-235 Fuel OD 1.25 in.
Fuct ID 0.068 in.
Clad ID 1.2745 in.
Fuel-to-Clad Dia. Cap .0245 in.
Pellet and Clad Roughness 70 pin.
Clad Material 348 SS Fuel-Clad Atraos. I atra lie.
, Capsule 11 Average heat rate 23.6 W/FT Fuc1 Density 95% T.D.
Enrichment .82 w/o U-235 Fuel CD 1.25 in.
Fuci ID 0.068 in.
Clad ID 1.2745 in.
Fuel-to-Clad Dia. Gap .0245 in.
rcilet and Clad Roughness 70 pin.
Clad Material 348 SS Fuel-Clad Atnos. I atm lle.
l l
B-25 '
- E Tats te 2-13. TAFY Comparison With Test Results I
I
, Capsule I: I Tic:e d' T (F) T (F)
I (llour s) KW/IT T (*C) (Experiment) TAFY 6.5 16.2 1750 3180 3230 .
8.0 21.2 2030 3690 3962 g 9.5 20.8 2050 '
3720 3905 g 12.5 21.2 2070 3760 3962 16.5 21.7 2089 3790 4029 18.0 21.95 2095 3800 4064 20.5 22.34 2313 3840 4115 l
5 24.5 22.92 2131 3870 4193 I
Caysule II : '
Time d' T (*C)
I !
T (F) T ( F) l (Hours) KW/IT (.brperimut) TAFY l 1.5 2.16 894 1640 523 2.5 8.23 1399 2550 1503 3.5 13,35 1720 3130 2584 5.0 16.61 1904 3460 3272 6.0 In.96 2073 3760 3614 l 7.0 19.20 2098 3790 3649 l b.5 21.79 2209 40:0 4013 9.5 21.95 2115 4020 4034 16.5 23.62 2239 4060 4253 11.5 23.73 2245 4070 4265 13.5 23.99 2250 4080 4299 15.5 24.23 2260 4100 4330 19.5 24.69 E g
? 's . 5 25.21 27.5 25.73 29.5 25.97- -
I B-26 Babcock & Wilcox l
l Figure 2-4. TAFT Comparison With Published Data j
DATA FRON: BALFOUR.M.G., CHRISTENSEN, J.A. ,
m AND FERRARI H.N., "IN-PILE NEASUREMENT OF U0 2 THERMAL CONDUCTIVITT"
[o
)
- , NARCH 1986 WCAP -2923 o I E 4000 h R 00 e
- r 5
o !
u 3 1 5
c :
o .
a \
3000 -
O U
E o
W 2
2000 = =
2000 3000 4000 MEASU3ED FUEL CENTERLINE TEMP 'F I
n-27 Babcock s. Wilcox
- w w--
l l
I I
4 Comparison of TAFY vith published data ource 4. "Densification Considerations in EWR Fuel Design and Performance".
Table 2-14. C.E. Test Data Flement DP4 AEC Pellet Diameter .489 in .488 in.
Clad O.D. .565 in. .544 in.
Clad I.D. .505 in. .5038 in. ,
Pellet Density 95% T.D. 95% T.D.
Enrichment 2.5 we % U-235 2.5 we Z U-235 f Atmosphere 1 ata He 1 ata He Fuel Surface Roughness 39 u in. 39 p in.
Clad Surface Roughness 20 p in. 20 u in.
Coolant Pressure 1050 psi 1050 psi Coolant Temperature 590 *F 590 'F Fuel 1.cngth 30 in. 30 in I
I I
I I
I I
d-28 Babcock 8.Wilcox
o Table 2-15. TAFY Prediction of C.E. Test Data Results lod AEC - (Beginning-of-Life): _
Cap Co'nductance Heat Rate (BTU /ER-Fr2_y)
(KW/FT) W/Zr-2 Clad W/SS Clad 10 435 478 11 446 485 12 460 494 13 473 503 14 488 514 15 506 526 16 525 540 17 547 557 18 572 575 19 599 597 20 630 620 21 663 647 22 716 688 23 773 736 24 837 789 Sorbed Gas Content = 0.05 cc/gm i
l B-29 Babcock & Wilcox
l 1
Table 2-15. TAFY Prediction of C.E. Test Data Results (Cont'd)
I Rod DP-4:
E Cap Conductance I
(BTU /HR-7T -F)
Heat Rate BOL 8000 MWD /Mr KW/rr Serbed Gas content Morbed t'a=
0.0 cc/gm 0.05 cc/gm 0.0 cc/ga_
10 684 436 223 11 694 448 236 12 707 461 249 m 13 721 474 265 l 14 736 490 282 15 754 507 300 16 774 526 321 17 796 549 343 18 821 573 368 19 850 601 393 20 881 631 436 21 916 668 477 22 955 723 521 23 1017 782 569 24 1091 847 622 I
I I
I I
B-30 Babcock & Wilcox
l s
l l
l Table 2-5. TAIT Comparison With Pablished Data l
GAPCON PREDICTIONS FOR GAP CONOLCTANCE l
10000 8000 0 Prediction (GAPCOM) 6000 O Data 4000 O Prediction (TAFY) c C O :
O m0 C O l
= C
- 3 C ?..
1000 C ..
g 5 0.0 Sorbed Gas SOL ,
g w v {
E /
j 600 _
- b l
2 M 0.05 Sorbed Gas BOL ~
{
l 8000 MWD /MT
/ 0.0 Sorbcd Gas I I I I I I I I I I I I i 10 Il 12 13 14 15 16 17 18 2D 21 22 19 23 24 Heat Rating (Kf/FT) l l
l i
1
"~3I Babcock a.Wilcox l
i
I-Table 2-6. TAFY Comparison With Published Data E
O naca cro. ucar-2923 capsule I 24.5 mil Diametral Cap fi 46M --
O CAPCON Prediction O TArv Prediction 6/ I 4400 - -
l 4200 -
w
,b I 4000 - -
h-Y !
U l 2 L1 l w O g D
ti 3800 -
CO 5 a l D ,
d U D e
I i g 1600 - -
~
5 l
~
1400 -
c 3 3200 g _
l l
I 3000 - -
l l l l I l 16 17 18 19 20 21 22 23 ;
Heat Rate KW/FT I
s-32 Babcock a.Wilcox l
4 4
Question 3 Provide additional information on the fuel cladding creep testa that were performed in the BAWTR and that form the basis for the B&W clad creep mo-del, CREC0L (ed. B&W collapse model), which is used to calculate the ex-pected collapse time for the Oconee Unit 1 fuel cladding. The requested information should include:
1
[ 3.2. _uestion Q 3(b)
Physical dimensions of samples including measured outside diameter, ovality, and thickness.
l Response CD = 0.430 2 <0.001 Initial ovality = <0.0005 Thickness = 0.0265 2 <0.0005 l
l l l
l l
l i
! B-33 1
f Figure 3-1.
Typical Tubing Profilemeter Trace E
a
- I g i . . I g
I I
l g i. . . _ . . I w -
+-
w .I x _..;...
i H l 0 : E w E a
H nn o
o a
z N
m o
l l
. g . - 1 I .
.l_.. . . . _ , . .. .j... _ . : _ _ .I _ ._.;' ._ _ l _
~
l . I ~ i . . .
t g . .
I e j
. i !
I I 8
- i- l I l 3' e
3
? .
t .
i ; ! ! I i i
! t :
- I_ _ _ - ._.. _ -_ .
! 8 ! !
! I I -
] l l _ ' . .ll
. . } . . ._ .l. . ! .
l TUBING AXIAL LENGTH I
s-34 Babcock a.Wilcox I
l
.i Figure 3-2. Typical Tubing Profilometer Trace e . . . i -
, I !
8 e
i I '
i , : :
l
.i. _ f . i. ..... .I .u ....s. l i
l .
, f 1
- . ._- 1 t
m:
w t-w 1
ct H
O I
w i o
l N
- vi )
W i
n o
o z
H m
3
>- -) t -
.._..w. :
, i :
l : ! !
. iI : !
I.
( . . e . .
. . . . . . . _ .. w . . _. i . . . . I
,e i .
i l.. -
i i i '
i
.t i
- i I : l i -
' i :
i .
I i '
i
, l
. . l . _ .. , ,
q . __ p .
i i i : .
! . i .i !
l
' ' ! . . ._ . i ....-
-l -
l 1 ; I : j ' ' :
l l l.
i l i .. l l l
t 1
l l TUBING AXIAL LENGTH i
I s-35 Babcock & Wilcox 1
t l
, . . . . , _ . _ - . - - - , , - . , ,.,,.--,- _,,-.,__.--..-.,._--.,_,..,,....n-,
1 4
I Figure 3-3. Typical Tubing Profilcceter Trace g
l.
i .
- . . __ t. . -
e
.. l. . . --. l 1
I e .
I .
cz-w i
y .
l I !. . --
H -
O !
w .. , - - -
a M
u>
p ..
3 O
O z ..
H m
D I
l
- . - I l
l l
, ia ,
i I
e
- l .
. e h !
- i l l a .
f
. 6 a
. ;. . _ . . :.. .; . ...? .. ; . . . . _ _____.:= 3. ..
l l l . t i. .
- i
- I !
l ! . : . ; i TUBING AXIAL LENGTH h s-36 BaW s Mm ,
I g
Question 4 In order to assess the B&W collapse model, CREC(L (ed. B&W collapse model),
that is used to calculate expected collapse tir,e for the fuel pins in Oconee Unit 1 the following additional information is required:
4.1. Question 4(a)
The equation used to calculate collapse ovality and a justification for not using the creeprate of irradiated ~uel in that calculation.
Response
(1)
,O(a 12 gr__ , g 3 ,(i) , 12 AH I
, ,ft-1) _ hJ (1-Y) ,g o h>
, ANg (i) _ g ,
(Ac + vac )
h g,v 2 l
y(i) o
, ,(1-1) , g(i) o o Refer to CAMD-9623 for a derivation and an e'xplanation of t e ms .
When the effective von Mises stress reaches the material yield strength, collapse is assumed to have occurred. The vfi) at this point is defined as the collapse ovality.
CRECOL is utilized only to calculate the ovality at which the cladding becomes unstable, given a certain temperature and differentail pressure. Since irradiation effects on physical properties (i.e., yield strength) are ignored, collapse ovality becomes a function of geometry only.
4.2. Question 4(b)
A detailed description and justification for the extrapolation of BAVIR test data to the collapse time by use of the Larson Miller i
Parameter (IlfP) including specific literature eferences to this method of extrapolation.
Response
In WAPD-nf-585 Figures 121, 122 and 123 depict a series of curves of hoop stress versus IRP for a variety of textures and geometries.
For the range of pressure from 2000-3000 psi, these curves are approp- l riately linear and they can be extrapolated to 1000 psi. The B&W l data was extrapolated over essentially the same range.
4.3. Question 4(c)
A justification for not including in the cladding stress analysis B-37 N C" I
I' I
such axial forces as cuased by pellet hangup, rod interference on grid platen, and rod bending at the spacer grids. 3
Response
Axial force due to pellet hangup will be accomanodated by the ficx-tble sparers located at both ends of the fuel rod.
All B&W fuel assemblies have sufficient gaps between the fuel rods and the upper end fitting to accoannodate all differential growth between rods and guide tubes. This is true for both operating and refueling conditions.
B&W spacer grios have been designed to permit fuel rod axial slip at forces less tian those necessary to cause significant bow in the fuel rods.
4.4 Question 4!d) - i A coer.arison of CRECet (ed. B&W collapse model) calculated critical ovality and collapse time with experimental clad performance data.
Response
Experimental specimens could only be observed during measurement per-tods, so it is irnpossible to determine the exact collapse ovality.
However, the equations used in CRECdL to calculate when stress reaches j the instability criterion include conservative assumptions. These lead to conservative collapse ovality predictions, as demonstrated by experimental data from unirradiated specimens at both 2250 and 2175 psi differeni.ial pressare. One specimen at 2250 psi had [ ] ails oval-ity without collapsing. Specimens at 2475 psi exhibited [ ],
and [ ] mils ovality without collapse vs. the CREC@L preaiction of about [ ] mils.
Experinental data were used to prepare Figure 3.3-3 in BAW-10054, so experi:nental collapse-times for specimens are implied directly from the curve (tice-to-collapse is defined as the time it takes to reach the conservative CRECWL collapse ovality).
4.5. Question 4(e)
A discussion of how flow induced fuel pin vibrations could affect the fuel pin collapse time.
~
I B-38 Babcock 8.WilCOX
w I
l
Response
As described in the Rancho Seco FEAR, mid-span fuel rod deflect-ions due to flov-induced vibration have been experimentally de-termined to be less than [ ] mil. This experimentation is appli-cable to all B&W fuel assemblies. The stresses imposed on the cladding due to this deflection are insignificant.
4.6 Ouestion 4(f)
A discussion of the clad temperature used in i he CREC9L (ed. B&W collapse model) calculation.
Response
ClaJding OD temperature along gap in the pellet stack drops to the contant tempersture within about 1/8". Temperature increase across the cladding wall has been calculated to be 3 degrees due to heat radiating from pellets at the boundary. Therefore, the assumed max-imum cladding temperatura for any gap is the hot-channel coolant ' temp-erature corresponding to the peaks described in item 5d, step 6, plus l 3 degrees.
4.7 Question 4(g)
A comparison of the B&W CRECdL code and the CPECWL code described in USAEC Report CAMD 9623, CCA, 1969. The comparison should identify any changes made to result in the present B&W version.
Response
1.
A srbroutine ZlRC" was added for Zircaloy-4 material properties based on BAW-3765-6, except for Poisson's ratio from WCAP-3629-41 and creep rate from AECL-2528.
2.
A subroutine "P6LY" was added to evaluate the material properties in 1.
- 3. A creep rate evaluttion change was made in " CREEP" to handle the different form of the AECL-2528 Zircaloy creep equation f rom the Haste 11oy and stainless equations already in CRECdL.
4 In subroutine 'stSTRESS", the initial time increment was increased to im;, rove run time. This had no effect on the 5 change-in-ovality limit per time increment.
B-39 hock & Mox
I quen, tion 5 In cader to perform an independent staff evaluation of the clad
! antegrity for Oconee Unit 1, the following information is requested:
5.1. Question 5(a)
Detailed discussion of the 0.9 value for the usage factor and of l the damage categories included in the analysis.
Response
j The 0.9 usage f acter is allowable cumulative fatigue factor.
not that predicted to occur in Oconee 1. The actual factor is
(
calculated (based on Miner's rule) for bending, thermal and hoop b stresses which occur during fluctuations in reactor power.
5.2. Question 5(b)
Collapse cine calculated with CREC6L (ed. B&W collapse model) with a comparison to one cycle and three cycle operating times.
Response
Predicted time-to-collapse is [ ] hours when fission-gas buildup is ignored. One cycle operating time is 7440 hours0.0861 days <br />2.067 hours <br />0.0123 weeks <br />0.00283 months <br /> and three cycle time is 22.320 hours0.0037 days <br />0.0889 hours <br />5.291005e-4 weeks <br />1.2176e-4 months <br />.
5.3. Question 5(c)
A list of operating conditions and physical properties including:
naxinsam operating time maximum external fuel pin pressure clad temperature clad outside diameter clad thickness initial ovality yield stress vs. temperature and fast flux elastic modulus vs. temperature Wl Poisson ratio vs. temperature internal fuel pin pressure vs. time fast flux
Response
- 1. Maximum operating time - 22,320 hours0.0037 days <br />0.0889 hours <br />5.291005e-4 weeks <br />1.2176e-4 months <br />
- 2. mxta mal fuel pin pressure - 2200 psia
- 3. Gacding temperature and rod internal pressure vs. time -
I I
B-40 Babcock s.Wilcox n-- --- . m w <,-~~ sm- -, - -,--a w~~w-s- g e e
4 Cladding OD - [ j in. mean, o = [ ] in.
- 5. cladding thickness - [ ] in. mean, o = [ ] in.
- 6. Cladding initial ovality -[ ] in. mean, o [ ]
inches.
7 Fast flux - 6.94 x 1013 - nyt (> 1 mev)
- 8. Yield stress elastic modulus, and Poisson's ratio as a function o_f temperature.
~
Temperature - F Jy - psi x 10 E - psi x 10' v 500
.265 525 .264 550 .262 575 .260 600 .258 625 .256 650
.255 675
.253 700 .251 725 .250 750 .248
- 9. Yield stress as a function of fast fluence -
(See following page).
Babcock & Wilcox B-41 .
I E _
I I
I I
I I
5.4 ouestion 5(d)
Discussion of the assumptions for the internal fuel pin pressure .
I vs. time, including the cold and hot BOL pressure with and withm t '
fuel densification.
Response
- 1. An initial gap between the pellet and cladding is calculated on the basis of as-built dimensions. Mean as-built pellet CD is combined with mean as-built cladding I.D. minus two standat3 deviations to determine a minintas BOL gap. These dimensions result in fuel pellets with low thermal expansion and fission-gas release, and therefore a rod with a high pressure differential.
2.
g Initial pellet density equals the mean as-built density for 5 3-cycle fuel assemblies. When the pellet densifies to 96.5%, a change in stack and pellet dimensions produces a change in rod internal pressure. The use of the minimum diametral gap in the above step produces a minimum gap vol-une which can contain backfill gas. The increase is void volume due to pellet shrinkage will therefore in this case have the greatest effect on BOL pressure.
- 3. Internal rod pressure preceeding densification is assumed B-42 Babcock 4.Wilcox
fl
\
to be the lowest allowable under the tolerance for backfill gas (all B&W fuel will be prepressurized).
4 Rod length, end cap dimensions, and internal spacer arrange-ment and dimensions are all asaused to be such that rod internal void volume is a ninf==.
- 5. Combining steps 1 to 4. BOL cold rod internal pressure is eniculated by the ecuation:
BOL PRESSURE = BACKFILL PRESSURE x Rod Void Volume before Densification Rod Void Volume Af ter Densification The increase in void volume is based on pellet dimensional changes calculated in accordance with the AEC technical report of November 14, 1972.
- 6. For each plant, the radial x local peaking histories of 3-cycle asses.blies are examined to determine a conservative worst-case rod peaking history. For each time increment (varying f rom 25 to 100 days) the maximum peak for any 3-cycle assembly is assumed to be the peak for that increment.
Thus t he f inal " assembly" peaking history represents a com-bination of peaks from many assemblies. Rod peaking history is obtained by applying a rod-peak / assembly-peak ratio [ ]
to the assembly history. A [ ] uncertainty factor is also applied. After relating the history to burnup, the result-ing peaking-burnup history is used as input to the TAFY code (along with the rod geometry and pressure described above) to comrute the rod internal pressure versus time history.
14o t BOL pressure calculated by the above steps represents a minimum value for the densified case. For Oconee 1, this pressure is [ ] psia. For the non-densified case, hot BOL internal pres ere is [ ] psia, based on a cold BOL backfill pressure of [ ] psia and a high power rod.
Babcock A Wilcox B-43 i
I question 6 in order toassess the B&W evaluation of transients and accidents, provide a complete and consistent set of design values and operating parameters for conditions with and without fuel densification. Where g
applicable, appropriate information in the Final Safety Evaluation 3 keport. FSAR. for Oconec Unit I should be referenced. The information requested should include:
- a. Core vide radial power map (see Figure 6-1).
- b. Radial local peak for hot assembly (See 6.6).
- c. Axial flux shape (see Figures 6-3 and 6-4) .
~
t
- d. local flux distribution in hot assembly tsee Figure 6-2) ;
- e. Mass inlet velocity to hot assembly (with and without.
one vent valve assumed open). (See 6.4 ).
- f. Loss coefficients for spacer grids and upper and lower cnd fittings (see 5.4).
Re.pon=e lhermal-hydraulic information for the Oconee 1 plant is provided as per requested for fuel densification an3 7ha t densification. This part icular sect ion addresses steady-state - .4 tons. Transiene and ,
accident condit ions are addressed in answers to other questions.
6.1. Core Operatine Conditions
- a. The reactor versel inlet temperatures are given below:
Nominal Maxiseum 100% power 554F* 556F ,
114% power 550.6F 552.6F
- b. The norsinal outlet pressure is 2200* psia and the minimum outlet pressure is 2135 psia.
I
- Values riven in the Oconce 1 FSAR.
Babcock a.Wilcox
6.2. Core Design 6.2.1. Fuel Assevibly Information
- a. There are 177* fuel assemblies in the core.
.h . There are 208* fuel rods / assembly with an outside diameter of 0.430* in. and an inside diameter of 0.377*
In.
- c. There are 16* control rod guide tubes / assembly with dimen-stons of 0.530* in. OD x 0.016* in, wall thickue== and one instrument tube / assembly with dimensions of 0.493*
in. OD x 0.441' in. ID.
- d. The fuel rod pitch is 0.568*in.
6.2.2. Fuel
- a. The s.adensified active fuel length is 144* in.
- h. The active length of the fuel with densification is 141.8 in.
- c. The cladding is Zircaloy-4 (cold worked) with a thickness of 0.0265 in.
- d. The undensified pellet is 0.37e* In. diameter and 0.700 1.1. Iong and the densified pellet is I J in. in diameter.
- c. L'n i t I core 1 is 93.52* of theoretical density (specified).
h . ). Power Distribution *
- a. The design core radial power map is shovn in Figure 6-1.
- h. The maximum fuel assently local rod power peaking distri-bution is shown in Figure 6-2.
- c. The p rcentage of the power generated in the fuel is 97.3%*
- d. The percentage of power generated in non-fuel regions is 2.7%*.
evalues given in the Oconee 1 FFAR.
B-45 Babcock & Wilcox
l l
l I
E 6.4 Fluid Flow
- a. Con 1.,nt F1tvs and Mass Velocities i
One !
Vent vent )
valves valve
]
closed open Total reactor vessel coolant flow 131.32* 132.60 (Ibm /h) (106 ) (106 )
Ef fective core-coolant flow (1bs/h) 124.23* 118.52 (106 ) (106 )
Average mass velocity at the core 2.53 2.41 inlet (Ibm /h-ft2) (106 ) (106 )
Inlet mass velocity to the hot 2.235 2.13 assembly (Iba/h-ft') (106 ) (106 )
- b. He core flow area (effective for heat transfer) is 49.19* ft2
- c. The loss coefficient for each spacer ' rid is [ }.
for the lower end fitting it is [ }, r 4 for the i.pper end fitting it is [ J. Rese form loss coeffi-cients are based on a flow area of [ ] in.2/ assembly.
6.5. Hot Chan::e1 Factors
- a. ne hot channel factor on average pin power (F ) is 1.011. It is applied on the enthalpy rise for the entire channel. The hot channel factor on local sun-face heat flux (F") is 1.014. h is value is applied locally on the calculated local surface heat flux.
- b. Flow area is reduced in the hot channel by a flow area reduction factor (Fg) of 0.98. W s vale is applied over the entire length of the channel.
c' . Flow is reduced in the hot bundle by a flow maldistri-bution factor which is 95% of the nominal isothermal -
bundle flow.
8 Values given in the Ocence 1 FSAR.
B-46 Babcock a.Wilcox i j
- d. The energy mixing coefficient a is 0.02.
6.6. Core Peaking Condis.fons The reference design 1.5 cosine symmetrical aximi power shape was used as s base case to determine if other axial power shapes in any way magnified the variation in IMBR. A 1.78 radial-3ocal asselaar peaking factor (Fah) associated with a 1.5 cosine = f=1 flux shape establish the maxistan dealgn condition resulting in the 1.55 INEE at 114Z of 2568 W (t).
The results indicate that outlet !==ha with the apike showed an overail larger degradation in DNBR than the densified 1.5 cosine axial power shape and its associated power spike. B&W utilized a conserve-tive 1.83 (P/P) outlet axial power chape in conjunction with a 1.49 (P/F) radial-local peak to maintain the reference deci;;= DEEE of 1.55 at 114Z of 2568 W(t).
This set of peaking conditions ==rimf zes the DNM penalty associ-ated with fuel densification and preenrpts the necessity of a reevalua-tion of all DNBR data for the power / imbalance / flow trip system. The penalty det'.rmined in this manner was used to modify tme power /in'>alsace/ l flow system as indicated in section 3.3.4 The 1.83 (P/I) outlet axial !
1 power shape shown in Figure 6-4 is precluded during normal operation as described in the Technical Specifications and as such in r.at a design criterion.
l 1
The 1.5 axial power shape in .onjunction with a 1.783 radial shape I
peaking combination is used for t.ansient and accident analyses. This particular shape results in a more conservative INSE ehme any ot!.er j shape that exists during norms 1 operation. This shape is shrwn in Figure i
63.
For LOCA analysis, the design bases axid power shaps wu a 1.786 peaking 0.5 feet below the core midplane. Th1= shape and pesk in cocjunc-tion with the calculated radial factor is the most conservative for the LOCA peak clad temperature analysis and could occur so-emrily during the period of xenon undershoot following a design basis (100-30-100Z) {
transient. Thf a peaking factor and associated radial are withir. the 1 l
l 1
a.47 Babcock a Wilcox l
j '
. . i ~.
e
- j M*
I DNBR limiting criteria statement in the previous paragraph. The reason I -
- 1. that in LOCA analysis, the important parameter is peak cladding te:n-pera t ure ; Wet as for 1*BR protection, the important parameter is not only heat flux and flux shape, but also the integration of heat input up t he channel ani the resultant enthalpy rise. g 5
A graph of power spiking versua axial length is giv?n in Figure 6-5. g m
The ncn-densified DNBR at design overpower is 1.55 and with den-sif icat ion and the spike utilizing the 1.83 axial power shape, the DNBR is 1.48. The reduction in overpower limit indicated in section 3.3.4 raised the 1.48 DNBR back to the design 1.55. g E
ti . 7 . Heat Flux Conditions The following data is based on the above peaking conditions so that a neaningf ul comparison between non-densified and denmified fuel can be su & .
3S-Densified Condittons
- a. The heat transfer surface area / fuel pin is 1.351 ft#.
- b. The average heat flux (q") is 171,470 5tu/h-ft2,
- c. The maximum heat flux at the minimum DNBR is 457,774 Bru/h-ft'.
19,, (MDNBR) - q,** x 1.55 x 1.49 x 1.14 x 1.014) i Axial (P/P)at PDNBR - 1.55 l i
(P/P) radial-local = 1.49 Max overpower = 114% of 2568 MWe*
!!at channel factor on local surface heat flux = 1.014*
- d. The average power density in the core is 83.38 kW/
liter, and the average linear heat rate is 5.66 kW/ft.
- e. The maximum clad exterior surface temperature at 100%
pouer is 650.0F for a pressure of 2135 psia.
avelues given in the 0:onee 1 FSAR.
j B-48 Babcock & Wilcox I l
l Dennified Condittens
- a. The heat transfer surf ace area / fuel pin is 1.3302 f t2
- h. The averace heat flux is 174.131 S t u /h-f. ' .
- c. T1.c n.eximum heat flux is $21.860 Btu /h-ft?.
( e; ** (*ux) q , x 1.74 x 1.49 x 1.14 x 1.011.]
Axial (P/P) at MDNBR with power spike = 1.74 (P/P) radial local = 1.49 Max everpower = 114% of 2568 MWt*
Ilot channel f actor on local surf ace heat flux = 1.014*
- d. Average volur etric power density in the core is 83.38 kW/ liter and the average linear heat rate ic 5.74 kW/f t.
This assurses all fuel pins have the densified active lengti whicle is conservative.
. The maximum clad exterior surface temperature 100%
power is 650F for a pressure of 2135 psia, i l
i l
- Values given in the Oconee 1 FSAR.
B-49 Babcock & Wilc0x
I I
I i
Figure 6-1. Typical Radial Power Distribution 4
l I
1.52 I.68 1.64 1.47 1.14 .35 i
I.64 1 69 1.62 1.44 1.16 .73 .33 f
I lr g 1.68 1.56 1.36 .99 .71 .31 l
n _
l.33 1.11 .43 ;
t l
.87 .59 .27 ;
~
l ,
FUEL ASSEll5LT
.35_ _
i AVERAGED P*/P
\
1 l
CONTROL R00 POSIT 10N I
I I
DEN & NOK l
t B-50 I i i
1 !
l i
F1sture 6-2. Mami. usa Fuel Red Pcwer Peaks i
i lI.l.
t l
- f. 4 [
- .. [ -
!I i .! !
- l g.
I
..s.
. D.
.n
- 1.
- I.. .. i.
- s. h.
23 '
1 A. mW.
' I r
- 92 .v
.", i .
. .M ... ., . 33 W. . .H.
3 .. .. . .
.I L 5 V...
g .1 .M
, ,..i.
j- ,
... I,,',...
. .- I . , ,I . .
g, g ... .
. .: I :. .. . e. 43_
m,
.... _m
'; ,/A . a. ;I '7 , e . u, .= /.r1./1 v Y ... . .
. .y .. . . Y. ,, Y. . ..i ,
., ., . . . . I ,,, ,,,
! I
-i .... ..., . .... . ,,. ,,
) 11 11 g , 3* . I. . . e; . ep . ..; ... :m y -
0
/f .i=ee..... ....... ...t.'
1 9 .. . . : i s l- i....... ... e.....
- c ... .. .
- :i. .n..e ..c.... . .s. :... ..
- . g ,,, , ,,, ,,,, ... . ..ei . ,e i. . ..s
. cou ... .. .. .. ..si .....i
- O ... c . .. . ..
- : .
- .
- E
- j L ,el' I' ;4 t
l
- i i -. .
lI 3
g
? i w a a
- 4
. I
(
, P !
? I I I 1 I
',T TEl' i l
l B-51 Babcock a.Wilcox i
l l
1
E Fleure 6-3. The Ef fects of Fuel Densification on the 1.5 Costoe Reference Design Asial Flux Shape E
l
- i
, O_
i
/
o
- O ! / ,
N e : l ,
l a 1 E,
=
2 g 2 5 f \
g . o / I
/
__i - __ =
~
l c
/ ' ~
~
l l I ;
/ ' U l l e ..
I . -. .n_ a 5
'y - :
- =
i i* : 3 E
. . .d. a
= l 1
M i
~
S
~ ~
I
~
4
. I J l I o R R . R e
- - o o o If dc d) J3seg ltIty altjaay 0; yead 3-52 BabC0Ch 4. WilC0K
i i
Fliture 64 The Effects of Densification on the 1.833 (P/P) Axial Flux Shape a
p L. - .- . .
_-. ~ . ,
w.
Y r
a w
o Ay --
n -
% . n. w s - o
%s -
LE "
e=, i N
w "E
-=
N N .
g _ -.
=
a \
r
=
\
s O
=
- a , = =
.a,, > w u
= c
-- ._.._ _L. _._ _ =_ = ,~
'1s e
=
. . - i c
- 1 s =
I a a,,. .
=
. P - c Jm 5 E
___ _ . _ _ _ . =_ = -
- =
)
w a
e o
=
. .j __ . _. , , ._
o i
1 I
_..{
k L _
=
l 0
o o e ~ o w a l
~ - . o a V( ud) adeus sesod leity a-53 Babcock a.Wilcox l
l i
i 1
1 l
I
- ! .* u r c 6-5.
Teve r Spike 5 actor Vs. Axial Fmition. occeee 1 I !
. ? I l
\
! o
= I I
E I ,
l
\ _ ,g
_ I ;
s S I T
U E
o 5 i
Ik
= I I
I a
I sotarj siids jased I
B-54 Babcock & WWCOE
iM tlon_f Pt vide .an .irtalysis of t*1 (1) loss of flow t rar . lent and (2) locked rotor .e< . a de nt for Oconee I without and with the man spelen of denstfled fuel. 1te inf ormation provided should acclude the followins:
- a. Surle.sr power decay.
- t. Core coolant flow decay.
- e. Ocre inlet pressure.
.t . DNhR vs. time.
- e. P.* a k c lad temperat ure vs. time.
!. Pe.e k f uel centerline temperature vs. time.
- r. Average heat fluz vs. time.
- f. . c.a p c ond uc tanc e vs . time.
- i. Clad to coolant heat t ransf er coef f icient .
R _e_e p o n s e
- 1. The loss of flow transient is included in the revised ver*ien of en -1387
- 2. I..wked Rot or Acc ident An ana lys t s ha s been per f ormed I.,r t he tecked rotor
.a cc ide n t with the assumptions outlined in Table 10-1 of the rod ejection
- c c r i . .n . II.c mwer dist ribur. ton was asstaned to be a 1.5 cosine with
.e power spite located at the point of minimum DNBR. Figure 7-1 shows the power. hot channel mass velocity, ans' calculated heat flux.
The pressure was assumed to be constant at 2135 psia.
The init tal power level for this accident was 1022 of 2568 MW. Trip occurs at about 3.9 seconds af ter which the power decays to a v lue of about 20* at about 10 seconds. Figure 7-2 shows the maximum f uel ten-perature and the film heat transfer coefficient at the point of maximum fuel temperature versus time.
The fuel temperature is affected very little since the power rises only slightly. Figure 7-3 shows the maximum cladding teiperature. the DNBR and the film coefficient at the point of maximum c ladding temperature. It is seen that the DNBR reaches the cr t-Babcock a.Wilcox n-55 1 P M
.w-
I tersen value of 1. ) at about 1.0 seconds af ter which the cladding tem-I pe r.a t ur.- increases to a va?ue of I X)OF which occurs !. 5 seconds after the i n i t 1.i t 8.*n of the Jr(adent.
I R
I I
I I
I I
- I I
I I
l I l I l I B-56 Babcock a. Wilcox l
_ _ . . _ _ _ . . _____.._________.i
Fi gure 7-1. Oconee 1 Locked Rotor II IO
~-_ _ _ _ _ _ _ _ _ _-' - - - ,s ,,,,
% [ FLUE
\
0.9 N
FLOW h E 0.8 '
i N - \
_E_
O' 7 y a
, IFITIAL Po6TR = 102 X 2366 We k
INITIAL MASS VELOCITY = 2.45 % 10 . .n-f:
2
!N!!!AL FLL*X = 174,130 5:.s/h-f t 2 06 s I i
OWER 0.5 l
0.4 l
0.3 0.0 0.4 0.8 1.2 1.6 2.0 24 Time (Sec)
B-57 Babcock a.Wilcox
i Farure 7-2. Ma x imus Fue l Terpe ra t ur e and F i lm Coe f f ic ient Versus Time for a Locked Fotor Accident 4100 ,
I 4000 ;
3900 \ 12000
,Q Fuel Temperature 3800 / 20000
/
m0
/ \ \ 18 0 l i \
- 3600
\ 16000 l
l
/( rel.\ \ i 8~
3 Coeffscient\ -
3 3500 \ 14000 $
/
E \ E I
/ :
- s \ 3 3 3400 12000 m
\
3300 sg \ 10000 l
\ I N I 3200 8000 Ngap = 850 8tu/hr f1 2 ..r 34 1 3100 6000 l
\
3000 0 0.8 1.6 2.4 3.2 4.0 4.8 5.6 E.4 Tese (sec) a-se Babcock a Wilcox
lI renur, 7-3. %1 . clad Temperature, DNR htlo, and Film Coefftelent for a Locked Rotor Accident.
H - a50 steun-f.2-r 80000
/1 -
760C0 lag _
72000 1
) {FILEC0EfflCIENT -
g 68000
'9 - 3M3 -
l _
% - 64C00 l -
60000 .
I8 - 1;ts l l
- 56000 CLA0 TERPERATURE
[
l -
52000 17 - 11M -
l -
48000 y i I
44000 "s : 6 _ito - I\
. [ _
40000 ;
j -
,E j l -
36000 f
/
" 15 300 -/ -
32000 is l ?
j 6 14 j t j i
"'* " o m -
rin00$
l 5 33e -
2400C 20C00 33 -
700 -
16000 l
l l -
12000 12 -
63s -
8000 l -
l -
4000
, L ___ _ __ _ ___._. _ __ - __ ,____. _ ,'
0 0I 1.6 2.4 3.2 4.0 48 56 Tsee (Sec) 8-5' Babece.k a. Wilcox Im I
r guesticn 6 Provide the technical bases and supporting analyses for your con-clusion that the 8.55 ft2 split break would remain the worst break for )
a l o s s -o f -c oo l an t accides.: tithin the break spectrum considering the effects of fuel densification. For the worst break provide curves 1
showing:
I
- a. Ifot rod axial flux distribution for the steady state i condition.
f
- b. Maximum clad tenperature ary! local rod heat tranafer coef ficient as a function of time. g )
5 1
- c. Ifot channel flew rate as a function to time.
Response
- Densified fuel af fects the CRAFT runs only to the extent that higher core average temperatt.res tend to cause slightly lower core i f lows during a loss-of-coolant accident . CRAFT is used to represent the average core condit ions and the average core temperature experiences a very small change in teeperature in going from 92.5% to 96.5% density during densif icat ion. The results presented in BAW-10034 were based on an average core teeperature equivalent to or greater than those existing wi'h densified 'uel. 1herefore, all of the CRAFT results presented in RAW-100 34 are appropriate for denstfied fuel. 5 As the core hydrodynamics and fluid condition presented in BAW-10034 are appropriate for each break size, the break which resulted in E W
the highest cladding temperature without the power spike is the break site which yields the highest temperature when the power spike is in-cluded. Since this spike is not dependet.t on break size, the worst break size presented in IMW-10034 is the worst break size when a power spike is assumed to exist.
The hot rod axial power distribution for a 4-day maneuvering transient, maximum clad temperature during the loss-of-coolant accident.
local hot spot heat transfer coef ficient, and hot channel mass flux are shown in Ficures S-1 through 8-4 I
B-60 Babcock 4.WilC0K
1 i
l As stated in this report, the radial peaking factsr was varied I until the raaximum linear power density consistent with the Interim Accep- ,
tance Criteria was obtained. The cladding temperature shown in Figure t l 8-2 is for beginning of life conditions with a maximum linear heat rate ,
l of 19.8 kV/ft. l l
l l !
i i
i t
i l
1 1
l l
l l
B-61 1
1 1
j l
- l l E 1
Figure 8-1. Axial Power Shape i
I l i S '
i I l 1 .
I 1 .
I I e
- 5 I E 5 -
E I 5 E- - E
- 20 . 5
. 05 L E g
I "
E E3 - A E 5 55
- E I , < a. => < g i E I -
1 l a
." l w 3 a
~
1 l m
l m
I .
I I
E I
a e a E
~
- 9 o = 9 a o
m - - a yead le!xy B-62 I
Babcock a.Wilcox e
b
N ricure 8-2. Hot Spot Cladding Tesoer.ature for 8.55 f t Cold Leg Break o
a S
. o O
- o 433 2
. =
49 m a a_ .
3 -
Sf
=
l
, o i
m l
t l
i
. n S
. . n . . . .
o o o o e o o o e o o o o o
,o o o ,o o o ,o o e o =
, , m m ~ ~ - -
3.*aintrJadmal Eu! ppg l0 B-63 Babcock s.Wilcox
I I
E Hot Spot Heat Transfer Coefficient for 8.55 ft Cold Leg Break Figure B-3.
I e g I
E
= 3 1
- R \
- o S $
. o l
W i
- M I
~
R I
- S r_ . . . .
E m o 7 a a g
- ~ _
2 S S 'o 5 2 2 2 2
- 3. glI-J4/ntg *)ua!3!;jao3 JalsurJ1 traH I
8* Babcock a.Wilcox
5'i cure S-la. Hot Spot Mass Flux for 8.55 ft Cold Leg Break R
. =
Y I
1 i
, i
)
l l
l
= a en
~
E C
. =
1 i - w t
. ~ ,
e a a a a a -
=
* =
A R R R 2 Z
33' N 'Al got x untj sses ,
l t
a-65 Babcock a.Wilcox
I flues e ion 9 Provide details of the assumptions and justification for establish-Ing the design transient 100-30-100Z power as the limiting transient.
!>i sc uss the axial menon oscillations that are included in the design transient analy=Is.
Respon ne A number of power transients were examined. These included varia-ttons in the range, e.g. 100-50-100; and variations in the rate of power )
recovery. The range of power reduction and t;ie rate of power recovery 1 1
.if fec t the amount of :nserted rod worth prior to power reduction and the
]
rod insertion for menon undershoot reactivity. The 100-30-100Z power a j transient, with recovery at time of peak x;non at the rate of 10% power /
minute, combine the ef fccts of maximum menon undershoot and maximum rod insertton. Hence, the resulting power peaks from this transient (design i power transient) are the most limiting.
The transient xenon effects on power distributions are included in f the nuclear analysis. The design power transient studies include: (1)
Peak xennn buildup during reduced pcaer. (2) rapid renon depletion and xenon underahoot at f ull power, and (3) xenon redistribution (axial oscil-lations). The reactivity ef fects and axia' power shif ts caused by tran-I sient zenon condition.m are controlled by control rod motion and APSR I
( positioning.
l Desian power transients with the resulting menon redistributions were examined throughout core lifetime. The hot channel power peaks =
are shown in Tables 3.1-3 through 3.3-8 of this report.
I ;
1 I
I I
l l
s-66 Babcock & Wilcox l 1
I Questicn 10 Di*cuns. In detail, and justify how the ef fects of fuel denaf fica-
' l on a re- included in the rod ejection accident analysis and comupare this analysis to the one in the FSAR without fuel densification. In particular, for full power cases, provide the initial peaking factors and their rela-tionship to the power spike em!el and to design limits in the Technical Spec i f ica t i on s for Oconee Uni t 1. Describe how the initial pellet density variation has t,een accounted for if other than by a 2 o variation on heat f lux and map increase. Provide the peak fuel temperature (average and centerlinc) and clad temperature as a function of time during the accident. Indicate the nusiber of fuel rods that experience DMB and will fat! during the course of the accident.
Response
An .a na l ys i s of the election of the naminum Technical Specificat ion value of rod v..rth f rom the core with the ef fects of fuel densification i included has been performed. Due to the ses11 changes in fuel geometry f and heat transfer characteristics. it was expected that no changes in the basic kinetic respcese of the core wuld occur and the calculations have verifled this expectation. We basic assuusptions for the calcula-t ions of the plant parameters are the same as presented in the Oconee 1 FSAR. Figure 10-1 shows the neutron power and pressure for the ejection of a 0.5%.%/k control rod at beginning of core life. The neutron power-reaches about 1201 prior to inward rod motion which occurs at about 0.6 seconds af ter which the p<uer decays to a value of about 30%. The pres-wure increane* to about 2320 psia due to the increased energy transfer to the coolant then decreages later on in '.he transient. Figure 10-2 shows the fuel tewperature and the heat t ransfer coef ficient at the point of maxiseum f uel temperature during the transient. It is seen that the maximum tesaperature occurs about 1.0 second af ter the peak neutron power and reaches a maximum value of about 4480F which is well below the assumed hczinnina of life meltina point of 5080F. The gap coefficient used was 160 Stu/h-f t -F this value was chosen to match the TAFT steady state f ue l t empe ra ture .and i t is conservative with respect to ceveerline fuel temperature. Table 10-I shows the important assumptions for the ther-n-67 Babcock s.Wilcox i
l t
ma l ana l y = I s . Figure 10-3 shows the assumed axial power distribution for j the thermal analysts.
Fizurc 10-4 shows the cladding tesperature, clad-to-moderator heat transfer coefficient and DNBR as a function of tise. The DNBR reached
- 1. 3 at about 0.6 seconds after which the maximum cladding temperature reached was 1305F. a value well below the assumed limit of 2300F.
It can be seen from the plot of film coefficient versus time that the film boiling heat transfer coefficient reaches a low value of 435 Btu /h-f t -F at about 0. 7 seconds and remains low for several seconds; however, the clad temperature decreases after about 2.5 seconds due to the decreased neutron power.
A parameter study was performed to determine the percentage of fuel pins t *na t would experience a DNBR less than or equal to 1.3. It w.i s determined that f or the rod worth analyzed (0.5%ak/k) about 13% of I
W tlee pin
- would exhibit a DNBR of 1.3 or lower.
I I
Table 10-1. Assumpt ions For Thermal Analysis Active fuel length, in. 141.8 Pettet diameter, in. 0. M5 Cladding thickness, in. 0.0265 Cap coefitrient 850 Btu /h-ft27 Filse coefficient Variable
- He; Channel Factors Overall pauer factor (F ) 1.0107 lecal heat flux factor (F") 1.0137 Flow area reduction factor 0.98 Assumed DN8 1.30 -
DNB correlation used U-3 Flow (vent va lve open) 95.4% of design flow 5 Errors Tinlet. F +2 Pressure, psi -65 Fluz trip setpoint.% +6.5
- Af ter a DNBR of 1.3 the Bishop. Sandburg. Tong correlations were used for both transition and film boiling.
B-68 Babcock a.Wilcox 1
Figure 10-1. Pressure and Iseutron Power vs. Time for Rod Ejtetf oe Accident (0.5Zak/k) for oconee 1 tisd *ainssesd 3 3 2 3 3 3 3 3
% % % % % % % N N,
- a i a u n
o \
d j ee Y
S l e
=
"- o g 1 O,.
E
_. =
a=
=
W _
a 3 -
- ?
a r
- ?
o 2
I I f I l w I l ,
e o e
" " 9 9 9 9 m e e. n -
unstarsj issed uojtnagg
)
Babcock s.Wilcox s-69 1
l l
I resure 10-2. hximum Fuel Temp Vs. Time and Water Clad Heat Trans fe r Coefficient vs. Time (at hx Fawl Temp) for Rod Ejection Accident h x Design Conditions 4500 I
4400 -
I FUEL TEMP 4200 .- N(2) = 850 -
20*000 COEFF HEAT TRAMSEER
, (H2 = 858)
) /' .
"_ 4000 4 / 16.000 ~
N -
/ \ 7
! E 3 / \
i / \ 5 l
~ \ - 12.000 ;
3300
/ \ : E
- 5 O
h / *-
n f A E 0
I E :
- N - 8.000 _ !
- 3600 -f ~
\
/
H2 = GAP COND. \ $
\
3400 -
4.000
- FUEL TEMP N H(2) = 1600 \ {
\ a 3200 ' ' ' i \ o l
0 1 2 3 4 5 Time (Sec)
I B-70 Babcock 4.Wilcox I
. - - - _ _ _ - ----_g--..--.--_------.-,,w-y..3-,, ,7 ---,------,---,.-,---,-,e
W 5
Figure 10-3. Slumped and Spiked Axial Flux - Oconec 1 I 1
=
l 1
E
/ ~
- /
-/- =
=f S
3 22 E
\
N, =
l
=
9 k *
(d./d)
, B-71 Babcock & WilcoK o
Tigure 10-4. Clad Temp armi Clad-Water llaat Transfer Coef ficient (at l W
l Nx Cladding Temp) for Rod Ejection Accident. Oconee 1 f
f -....... g i
I
- i. -
lIr-NEATTRANSFER
[
'g CDEF ICIENT 1300 - I -
7s. sos g
1.9 -
g 1200 g
'l se.ose : 3 3l [ W 1 I I.8 I [ E
'IO0 -
l l ll g
- 54,884 y
= 5 t Il CLAD TEMPERATURE
- II te I
'7 - - Ii ~_~
- -- g 3 i g 43,000 '-
1000 -
l
- l - 1- .
1.6 - .
, ._ g em u 5 30.000
} 900 -[ [ f.- ,
=
E # , j
- ,, _ , 3 g EINlEUR ONOR 800 -
20.000 1.4 -
100 -
l t .000 i I t.3 I
1 E 600 t i _ 1_ _ _.L '
e 5 3 4 5 g 7 Time (Sec)
Babcock a.Wilcox s-72
_ _ - _ . . _ - _ _ _ _ _ .__-._J