ML19322C136

Fuel Densification Rept (Non-Proprietary Version of BAW-1399).

ML19322C136
Person / Time
Site:	Oconee
Issue date:	03/31/1974
From:	BABCOCK & WILCOX CO.
To:
References
BAW-1400, NUDOCS 8001090554
Download: ML19322C136 (40)

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Oconee 3 FUEL DENSIFICATIO'i REPORT (Nonproprietary Version of EAW-1399) i l

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! Power Generation Group Nuclear Power Generation Division P.O. Box 1260 i* Lynchburg, Virginia e*505 Babcock & Wilcox

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Babcock & Wilcox Poser Generation Group l Nuclear I'over Generation Livisica l Lynchburg, Virginia l

l Report BAW-1400 March 1974 Oconec 3 Fuel Densification Report (Nonproprietary Version of EAW-1399)

Key Vords: Fuel, Densification Effects ABSTRACT In June of 1973, BAW-LOO 55, Rev.1, was filed with the AEC in ac-cordance with the guidelines set forth in the ALC .enort, " Technical Report on Densification of Light Water Reactor Fucis," datad November 14, 1972. This revision incorporated the answers to additional ques-tions from the AEC Staff concerning generic items on fuel densification.

In October of 1973, B&W filed an additional report, BAW-LOO 79 ,

" Operational Parameters for B&W aodded Plants," which sets forth the core operating parameters for B&W rodded plants. This report established the loss 'of-coolant accident (LOCA) basis for determining the m:irimum allow-able heat rate and outlined the analysis used to determine plant operat-ing restrictions owing to the pestulated ef fects of f uel densification.

Questions relating to individual plants (as-built data, etc.) are answered in individual reports which are filed for each plant.

This report, along with the appendix, presents an analysis of the effects of fuel densification on the fuel for Oconee 3 and supports the safe operation of that unit at the rated power level of 2568 K4t.

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I CONTENTS l Page

1. INTRODUCTION . . . . . . . . . . . . . . . . . .. . . . . . 1-1

2. CONCLUSIONS. . . . . . . . . . . . . . . . . . . . . . . . . 2-1 3 R ESUi.TS . .. ........... . . . . . . . . . . . . . 3-1 3.1. Power Spike Model. . . . . . . . . . . . ...... . 3-1 3.2. Thermal Analysis . . . . . . . . . . . . . . . . . . . 3-4 3.2.1. Fuel Temperature Analysis. . . . .. . . . . . 3-4 3.2.2. DNBR Analysis. . . . . . . . . . .. . . . . . 3-4 E 7

3.2.3. Summ ry . . . . . . . . . . . . . .. . . . . . . 3-5 5 3.3. Nuclear Analysis . . . . . . . . . . . . . . . . . . . . 3-11

! 3.3.1. Reactor Protection System. . . . .. . . . . . 3-11 g 3.3.2. Analysis of Power Distributions g Before Densification . . . . . . .. ... . . . 3-11 3.3.3. Analysis of Power Distributions With Densification Ef fects . . . .. . .. . . 3-13 3.3.4. Summary. . . . . .. . . . . . . . .. . . . . 3-15 3.4. Safety Analysis. . . . . . . . . . . . . . . . . . .. . 3-21 3.4.1. General Safety Analysis. . . . . . . . . . . . . 3-21 3.4.2. LOCA Analysis. . . . . . . . . . .. . . . . . 3-24 3.5. Mechanical Analysis of Oconee 3 Fuel . . . . . . . . . 3-36 3.5.1. cladding collapse. . . . . . . . . . . . . . . . 3-36 3 3.5.2. cladding Stress. . . . . . . . . . . . . .. . . 3-36 5 3.5.3. Fuel Pellet Irradiation Swelling . . . .. . . . 3-36 APPENDIX - Design Parameters for Oconee Unit 3 . . . . . . . . A-1 l

I List of Tables Table 3.2-1. Fuel Temperatures at Low Power Density . . . . . . . . . 3-6 3.2-2. Fuel Temperatures at High Power Density. . . . . . . . . . 3-6 3.2-3. Ef fects of Fuel Dencification on DNBR and g Pcare r Ma rg in a t 114 % o f 2 568 HWe . . . . . . . . . . . . 3-7 5 3.3-1. Modifications to Reactor Protection System Setpoints and Design Parameters. . . . . . . . . . . . . 3-16 3.4-1. Thermal Data Input for Safety Analysis . . . . . . . . . 3-26 1

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- Table:s (Cont'd)

Page Table

3. 5- 1. C1.dding Circumferential Stress. . . . . . . . .. . . . 3-37

) List of Figures Figure 3.1-1. Maximum Cap Size Vs Axial Position . . . . . . . .. . . . 3- 2 3.1-2. Power Spike Factor Vs Axial Position . . . . . . . .. . . 3 -3 3.2-1. Maximum Fuel Temperature Vs Linear Heat Rate . .. . . . . 3-d 3.2-2. Average Fuel Temperature Vs Linear Heat Rate .. .. . . . 3-9 3.2-3. Cao Coefficient Vs Linear Heat Rate. . . . . . ... . . . 3-10

3. 3-1. Trip Setpoints Vs Axial labalance Without Densification Effects. . . . . . . . . . . . . .. . . . .'3-17 3.3-2. Calculated Offset Limits Vs Power. . . . . . . . . . . . 3-18 3.3-3. Trip Setpoints vs Axial Imbalance With Densification Effects. . ... . . . . . . .. .. . . . 3-19 3.3-4. Envelope of Maximum Operating Linear Heat Rates as Function of Axial Position. . . . . . . .. . . 3-20 3.4-1. Pressure, Power, and Flux Vs Time for Densified Fuel, Re1 Ejection Accident. . . . . . .. . . 3-27 3.4-2. Slumped and Spik.J Axial Flux Shape. . . . . . . . . . . 3-28 3.4-3. DNBR, Fuel and Cladding Temperatures, and Film Coef ficient Vs Time for Rod Ejection Accident. .. .. . 3-29 3.4-4. Power, Flow, and Flux Vs Tinte for Densified Fuel, Four-Pump Coastdown. . . . . . . . ....... . 3-30 3.4-5. DNBR and Film Coefficient Vs Time for Densified Fuel, Four-Pump Coastdown. . . . . . . . 3-31 1.4-6. Power, Flow, and Flux Vs Time for Densified Fuel, Locked Rotor Accident. . . . . . . . . . . . . . . 3-32 3.4-7. Cladding and Fuel Temperatures and DNER Vs Time for Densified Fuel, Locked Rotor Accident. . . ... . . . 3-33 3.4-8. Sensitivity of Allowable Peak Linear Heat Rate to Moderator Coef ficient ... . . . . . . . .. . . . . 3-34 3.4-9. Axially Dependent Linear Heat Rate . . . . . . . . . . . 3-35 A-1. Design Radial Power Distribution . . . . .. . . . . . . . A-7 A-2. Maximum Fuel Rod Power Peaks . . . . . . .. . . . . . . . A-8 A-3. Effects of Fuel Densification on 1.5 Cosine Reference Design Axial Flux Shape. . . . . . . . . . . . . A-9 A-4.

Fffects of Densification Flux Shape on 1.833 (P/P)A A*i"1

. . . . . . . ...... . . . . . . . . . . . A-lO I

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1. I NTRODUCTION This report documents the effects of postulated fuel densification for the Oconee 3 core as calculated in accordance with guidelines set forth in the AEC report o' November 14, 1972. The application of these guidelinas to the results presented in this report is discussed fully in B&W's proprietary topical report BAW-10055. Rev. 1. " Fuel Densification Report." Further considarations as presented in BAW-10079, " Operational Parameters for B&W Rodded Plants," were also taken into account.

The analysis of Oconee 3 is limited to an exanination of the first fuel cycle. Babcock & Wilcox now has operating plant data on the Oconee 1 fuel, and there are no signs of fuel densificat'.on after 75 EFPD. It is expected that data from other pressurized water reactors (PWRs) now l operating with prepressurized f uel will allow relaxation of the current guidelines. Before the completion of the first cycle, a supplementary report will be filed for Oconee 3 to cover three full cycles of opera _

tion at 2568 Mut. )

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2. CONCLUSIONS i

Based on the analysis performed for Oconee 3, which utilized the

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methods given in EAW-10055. Rev. 1, and BAW-10079, the following conclu-sions are made even if the fuel pellets are assumed to densify to 96.5%

[ of their theoretical density:

1. The cladding will not collapse because all B4W fuel rods are Pressurized.

2. The mechanic.nl performance of B&W fuel rods will not be ime-paired.

3. The interim acceptance criteria for the emergency core cooling l system (Et.CS) will not be violated.

4.  % e re. actor can be safely operated at the rated power level of L 2568 MWt with the reactor protection system (RPS) setpoints outlined herein. n ese modifications ensure that the thermal design criteria are not exceeded.

5 W e modifications to the RPS are a reduction in the overpower s trip setpoint, f rom 114 to 112% of rated power, and a minor reduction in allowable imbalance limits as shown in Figure 3.3-3.

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3. RESULTS L

, This section of the report covers four rain topics: thermal anal-L ysis, nuclear analysis, safety analysis, and mechanical analysis. ne thermal analysis section considers protection of the fuel melt und D.BRT criteria. The nuclear analysis section considers thermal des gn crite-ria, imbalartce trip limits, and core operational limits. The safety analysis section reanalyzes all postulated accidents analyzed in the Oconee 3 FSAR assuming that densification occurs. De mechanical anal-ysis section contains the input summary and results for cladding creep and collapse, cladding stresses, and fuel pellet irradiation swelling.

Since complete as-built data were not available for this analysis, the most conservative values from the specification are used in each analy-sis.

3.1. Power Spike Model The AEC guidelines outlined in Technical Report on Densification L

of Light Water Reactor Fuels " November 14, 1972, have been used to de-

~ termine the maximum axial gap as a function of core height. The proba-u bility values (F ) ggiven in the same report (Table 4.2.A. colume 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 individ-ual reactors. The naximum gap size versus axial position is shown in f

Figure 3.1-1, and the power spike f actor versus arati position is shown in Figure J.1-2. These figures also show the initial and final theoret-ical densities (TDI, TDF) used in the calculations. These data form the basis for the analyses in this report.

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I 3_. 2 . Thermal Analnis, I

h2.1. Fuel Temperature Analysis Utilizing the analysis established in BAW-10055, Rev. I plus taodifications as requested by DOL, a fuel-to-cladding cold diane-tral gap of 12.45 mils af ter densification was nalyzed. The results of this analysis are presented in Tables 3.2-1 and 3.2-2 and in Figures 3.2-1 through 3.2-3.

The modifications are as follows:

1. TAFY" thermal code

a. No fuel restructuring.

b. A 25% reduction in gap conductance.

2. Inputs to TAFY

a. Most conservative specification data used for fuel density and diameter and for cladding ID (Table A-1).

3.2.2. DNBR Analysis The thermal effects due *o densification can be divided into two categories: (1) the result .. the reduced stack height and (2) the combined result of the reduced stack height with the power spike superimposed. Therm al ef fects are then imposed on calculations of the minimum departure from nucleate boiling ratio (DNBR) used to set ther-mal design limits.

The reduced active length was calculated to be 139.94 inches, which represents a reduction of 4.06 inches from the nominal ac-tive length of 144.0 inches. The most conservative specification infor-cation given in the appendix was used in calculating this densified ac-tive length.

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I s I The axial flux shape that gave the caimum change in DNBR from the original design value was an outlet peak with a core affcat of

+11.8%. The spike magnitude and the maximum gap size used ir. the analv-sia are 1.100 and 1.65 inches, respectively. The results of the two en-fects are sumrurized in Table 3.2-3 in terr.s of percent .ge chaage in minimum hot channel DNBR and peaking margin.

3.2.3. Suna.iry_

This analysis assumes that densafication and arsociated phenomena vill affect the hot channel, which has the most limiting h Both the fuel temy.ra-thermal-hydraulic characteristics in the core.

tute analysis and the DNBR analysis were conducted indepeMently with the respective most conservative specification values. In addition, the power spike is assumed to be located at the hot channel position that minimized DNBR. The resultant loss in DNBR of 4.4% .esults in a DNBR of 1.48 at 114% of 2568 MWt. This is equivalent to a 2.1% loss fn I allowable power peaking. "Ihe inclusion of control rod insertion limits as well as the reduction of the overpower from 1142 to 112% of 2568 HWt -

compensates for this loss. The plant can then function at t*w f ul ,

core rated power level v'*hout violating the design criteria "or PNBR and/or centerline fuel metting. "I'ae allowable pos.tr shapes and the new offset limits are disc 2ssed in secti3n 3.3.

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Tabic 3.2-1. Fuel Temperatures at few Power Density Cap coeff, Surface fuel Average fuel Maximum feel Density, Cold gap, teep, F temp, F

% TD mils kW/ft Etu/h-ft 2 'F temp, F 977 1337 1733 96.5 12.45 6.0 680 Table 3.2-2. Fuel Temperatures at liigh Power Density, Surface fuel Average fuel Maximtna fuel Density, Cold gap, Cap coeff, Btu /h-ft 2_eF temp, F temp, F ,_ temp, F

% TD mils kW/ft _

Y 1483 3126 4849

• 96,5 12.45 18.9 965 Y

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Table 3.2-3. Ef fects of Fuel Densification on DNBR and Power Margin at 114% of 2563 MWe Densified active Densified active length leegth and power ar,ike Axial DNBR Za Ia DNBR Za IA power shape (W-3) DNS gin (W-3) DNB Martin Outlet pean with +11.8%

core offset 1.50 -2.8 -1.3 1.48 -4.4 -2.1

!.OTE B&W topical report BAW-100?4 describes the TAFY computer

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program. The code has been used as described in the -

analysis of fuel densification except for the following:

1 The option in the code for no restructuring of fuel has been used in the analysis pre-sented here in accordance with DOL's interim evaluation of TAFY. l l

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Far,ute 3.2-1. Watsum Fuel Terperature Vs Linear Heat Rate I f m

$200 utsting Temperatu.e e BOL UO7 E I

5000 - ,,oys 251 Reduction in Cap Coeff.

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2. No Fuel Restructuring 4600 -

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1200 1000 I l I I I I I 800 I O 2 4 6 8 10 12 14 15 18 20 22 24 Linear Heat Rate. LE/ft ,

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Figure 3.2-2. Average Fuel Temperature.Vs Linear He.1t Eate a  ! I ASSUMPil0NS 3600 -1. 25% Reduction in Gap Coeff.

2. No Fuel Restructuring 3400 -

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Figure 3.2-3. Cap Coefficient Vs Linear Heat Rate 1700 -

1600 1500 _ ASSUMPil0NS .

1. 25% Reduction in Cap Coeff.

2. No Fuel Restructuring 1300 -

1200 -

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3.3. Nuclear Analysis 3.3.1. Reactor Protaction System The safe operation of a reactor core requires an extensive analysis of power distributions resulting from the various modes of plant operation. ne primary considerations and results of this analysis are as follows:

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1. Assurance that th'ermal criteria are not exceeded; i.e.,

specified minimum DNBRs and centerline fuel temperatures may not be violated.

2. Definition of imbalance limits to prevent adverse power peaks that would exceed the foregoing criteria.

3. Definition of core operational limits and recommended operating procedures to pr.*. vent unnecessary reactor trips.

The complete maneuvering study entails a combined nuclear-therreal analysis of the power distributions. This section describes the methods and criteria used in developing the RPS setpoints and in modify-ing the setpoints required to account for postulated densification ef- _

fects.

3.3.2. Analysis of Power Distributions Before bensification The three-dimensional PDQ07 code with thermal feedback el-fccts is used to analyze power distributions. This analysis determines power distributions 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. The -

fuel cycle and transient analyses determine power distributions for nor-mal equilibrium and transient conditions, respectively. The extremes of core operation, such as control rod bank insertion beyond normal limits  !

and maloperation of axial power shaping rods, are also examined. The l

extreme control rod bank conditions define the limito for the imbalance l protection system.

3.3.2.1. Correlation of Power Peaks to Thermal Design Criteria The power peaks from PDQ cases are corrected for calculational uncertainty and are w.alyzed to determine the margin to the 3-11 Babcock &Wilcox l

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thermal criteria: centerline fuel melt and departure f rom nucleate boil-ing (DNB). The margin to centerline fuel melting is defined as I

j Fuel melt margin =Max ( Max allowable calculated peakpeakf

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• The maximum allowable peak is defined as the pointwise power that yields centerline fuel melting:

22.2 kW/ft I

Max allowable peak = 5.66 kW/f t = 1.014

• FOP where 22.2 kW/ft = fuel nelt limit.

I 5.66 kW/f t = average heat rate at 2568 Wt.

1.014 = hot channel f actor.

FDP = fraction of power.

The naximum calculated peak is the largest total peak from the PDQ power mps increased by a f actor of 1.075 to account for calculational uncertainty.

The determination of DNB margin requires a more complex analysis. DNBR is a function of peak location, magnitude of the power peak component parts (radial and axial), and other core parameters.

To crrive at true DNB conditions, each power distribution is analyzed explicitly. From the PDQ power distribution, the maximum calculated total peak is obtained and adjusted for uncertainty. The DNB margin is then defined as ,

I Allowable total peak DNB margin = ( Max calculated total peak

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The basis for the allowable 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 most limiting.

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l 3.3.2.2. Offset-Margin Relationship Core offset, a measure of the axial power in-balance, is defined as the fraction of total core power in the top half of the core minus the fraction of total core power in the bottom half of the core:

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Offset = Power (topp + power (botton) he relationship between hot channel power peaks (i.e., thermal margins) and core offset defines the protection syst.eu setpoints. Power imbalance is the primary signal to the protection system for flux shape protection.

De maneuvering analysis defines the relationship between cora imbalance and thermal margin.

Limiting of fsets a i determined to prevent the violation of thermal criteria for all operating conditions and power levels.

To yield the imbalance trip envelope, the limiting offset values are cor-rected i'or potential instrumentation errors, imbalance detection bias, and calibration.

The imbalance trip envelope defi. -s the range of allow-able operational imbalance and ensures that 1.3 DNBR and/or the central f uel melting limit will not be exceeded. Figure 3.3-1 presents the trip setpoints based on these criteria. The overpower trip setpoint shown in Figure 3.3-1 is controlling for overpower transients, whereas, the solid horizontal line is the trip for loss of flow transients.

3.3.3. Analysis of Power Distributions .

With Densification Effects 3.3.3.1. RPS Considerations Provision for possible fuel densification re-quires modification of the imbalance trip system for two reasons: (1) the fuel seit (kW/f t) criterion change, armi (2) an additional power spike is included in the reactor power distributions. Since the power spike factor is a function of axial position, the appropriate power spike factor is used to increase each PDQ peak to account for potential densification.

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I The modified offset limits with fuel densifica-tion effects included are presented in Figure 3.3-2 and are compared with the previous offset licits. The primary dif ferences between the two sets of calculated limits are as follows:

1. The DNER loss of -4.4% results in a peaking margin loss of

~ 2.12.

2. The central fuel melting limit changes f rom 22.2 kW/f t before densification to 20.15 kW/f t.

3. A 4.1-inch decrease in fuel colums length increases the nomi- I nal heat rate at 2568 2dWt f rom 5.66kW/f t before densification to 5.82 kW/ft after densification.

4. The local power spike f actor is applie.d to the calculated power ,

distributions.

i. The overpowe limit in the imbalance protection system is re-defined as 1122 of 2568 S't. The ef fect of the reduced overpower limit is one-to-one for local heat rate and approximately two-to-one for DNBR. g  ;

g i The trip setpoints are obtained from the calculated of fset limits by adjusting for potential electronic errors and offset measure-ment bias by the out-of-core detectors. The error-adjusted limits for densified fuel are shown in Figure 3. 3- 3. The ic6alance trip points g and overpower trip provide operating flexibility with assurance that ther- 5 mal criteria are not exceeded. Furthermore, potential relaxation of these limits may be realized as B&W obtains operational data and experience with Oconee 1 and 2.

3.3.3.2. ECCS Considerations ECCS calculations have resulted in an axial- E dependent kW/ft limit as shown in Figure 3.3-4. (See section 3.4.2.2 5 for f urther infornition.)

1he maximum operating heat rates are maintained lower than this limit ty imposing restrictions on certain core operat-ing parameters. The maximum allowable heat rate and the maximum ex-pceted heat rate for Oconee 3 are compared in Figure 3.3-4.

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1he derivatian of the operating restrictions is fully described in BAW-10C79, which includes consideration of the fol-lowing operating parameters:

1. Fuel depletion.

2. Control rod position.

3. Azial powei imbalance.

4. Transient menon.

5. Quadrant power tilt.

Appropriate controls will be provided to ensure that the LOCA heat rate limits are not exceeded during plant operation.

3.3.4. Suwmary Fuel densification and associated design limit changes have required modifications to the technical specifications. The power peaking margin loss of 2.1% f rom the DNB analysis, the lower fuel melt-ing limits, and the additional power spike factor have been compensated by a 22 reduction in design overpower and by more stringent of fset lim-its. The revised te.:hnical specifications allow operation at 100%

power with assurance that thermal criteria, with all densification ef-fccts included, are not exceeded. The modifications are summarized and i

compared with the previous system in Table 3.3-1.

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I Tal,1= 'J.3-3. It,latications to Rasctor Protection System l B

_,,_Secpoints and Design Parameters A. linb al ang s t ein Previa 2s Hodified system Paveseter _ system _ (densification)

Fuel a.est 11mst, bW/ft 22.2 20.15 1.

2. DNb peek 's.4 ma. gin penalty, g

- 2.1 ikin ana.1 1.e a t rate, kW/fL 5.66 5.82 1.

114 112 4 Overpowec, I of 2568 MWt

5. 8sfi.et t .e l t s .s t rsted power

+49 +34

a. Positise ot'fset

b. Ngt se offset -56 -36

6. Trip acteoints at rated power

a. Pw oeiee in. hat.,nce +22 +15

b. r% ,.cIve s .Aa t a.u.. -33 -15

7. Spake r . t a. None 1.00 to 1.101

8. Nuc sear 5 .e pcab sincertainty 1.075 1.075

d. yo ft _1;p y n; C...a a ul -1.OCA kW/ft Limit A series of operating restrictions as gives in BAW-10079 has a been ic: posed on plant operation to limit the peak. linear heat g rate to less than the axially dependent IDCA kW/f t limit.

These will be factored into the technical specifications as l was done for the Oconee 1 and 2 applicatics.

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Figure 3.3-1. Trip Setpoints Vs Axial Imbalance Without tensification Effects 128

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I Figure 3.3-2. Calculated of f set Limits Vs Prun Prior to Densification i10 / 1

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.I I Figure 3.3-3. Trip Setpoints Vs Axial Imbalance With Densification Effects 110

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-40 -20 0 20 40 Core labalance, 5 3-19 BabC0CM & WilC0K u

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I FIE ure 3.3-4. Envelope of Maximum Operating Linear Heat g Rates as Function of Axial Position 5 18 ~

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Manisus Operatsag _ l I

Linear Heat Rate 10 l

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Babcock & Wilcox 3-20 I

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3.4. Safety Analysis 3.4.1. General Safety Analysis _

3.4.1.1. Introduction The significant effects of fuel densification are an increase in maximum fuel temperature and a slight increase in average heat flux due to shrinkage of the pellet stack length, in addi-tion, spikes in the neutron power can occur due to gaps in the fuel.

These combined effects will lead to a slightly decreased initial DNBR for the accident calculations presented in the Oconee 3 FSA1. For over-power transients such as rod withdrawal, the ef fects are of fset by a re-duction in the overpower trip setpoint. The parameters used in the analysis are the same as those used in the FSAR 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 maintain conservatism the original values were used. All calculations were made for BOL conditions.

3.4.1.2. Reactivity Insertion Transients The rod withdrawal was not recalculated since for all combinations of parameters, including the simultaneous with-drawal of all rods in the core, the peak thermal power attained during the transient is always less than the 112% design thermal power level; therefore, the 1.3 limit on DNBR is maintained for this transient.

The startup of an inactive loop was not consid-ered in the analysis since the maximum thermal power achieved during the I 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 f will perturb the flux shape sufficiently to exceed design conditions at The l 112%, such occurrences still do not present any thermal problems.

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 power level. There-fore, the 1.3 limit on DNER is maintained.

3-21 Babc0Ck 4.Wilcox

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The ejection of the u.aximum technical specifica-taon value of rod worth (0.65%) f rcca the core considering the ef fects of fuel densification. has been analyzed. The basic assumptions for the c.alculations of the plant parameters are the same as presented in the Oconee 3 FSAR. Figure 3.4-1 shows the neutron power fraction, pressure, and core average heat flux fraction for the ejection of a 0.65% ok/k control rod at beginning cf core life. The neutron power teaches about 710% prior to inward rod motion dich occurs at about 0.6 second after which the power decays to a value of about 302. The pressure increases to .about 2465 psia due to the increased energy transfer to the coolant.

then dec reases later on in the transient. Table 3.4-1 shows the impor-t .a n t assumptions for the thermal analysis. Figure 3.4-2 shoo the axial g pwer distribution used for the thermal analysis. Figure 3.4-3 shows N the fuel and cladding tenperature at the point of maximum temperature Juring the transient. It is seec that the fuel ten:perature reaches centerline celting at about 0.8 second after the peak neutron power.

The gap coefficient used was 669 Bru/h-ft 2 *F; this is an effective gap value chosen to match the TAFY steady-state fuel temperature. Figure l

! 3.4-3 also shows the cladding te:nperature, c1md-to-moderator heat trans-fer co. 'ficient and IEE ratio as a function of time. The IEB ratio reached 1.3 at about 0.4 second after which the maximum c1mading temper-ature reached was 1560F, a value well below the assumed limit of 2300F.

1 It can be seen from the plot of film coefficient versus time that the film boiling heat transfer coefficient reaches a low value of 450 l Btu /h-f t 2 *F at about 0.35 second and remains low for several seconds; .

however, the clad temperature decreases af ter about 2.2 seconds due to the decreased neutron power. A parameter study was performed to de-termine the percentage of fuel pins that would experience a DNBR less l than or equal to 1.3. It was determined that for the rod worth analyzed I (0.65 ak/k), about 28: of the pins would exhibit a DNBR of 1.3 or lower.

The - i-n- hot spot fuel enthairy was found to be about 147 cal /gs.

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3-22 Babcock s,Wilcox l

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Secondary system accidents resulting in a power increase occur at or near end of life (EOL) when a highly negative mod-erator coefficient exists. Since more DNB margin exists at EOL, these cecondary accidents, such as a steam line break, are not expected to cause thermal limits that are more severe than those presented in the FSAR. We FSAR analysis of secondary system accidents, such as stea e gen-erator tube ruptures and loss of electric power, is unchanged since the thermal power remains the same or decreases during the transients and, theref ore, 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 ana-lyzed 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 as shown in Figure 1.4-2. The other parameters normally considered in the coastdown calcu-lations remain unchanged from the FSAR values. Figure 3.4-4 shows power, flow, and the calculated core average heat flux fractions for a four-pump coastdown initiated from 102%. Figure 3.4-5 shous the calculated DNBR and film coef ficient as a function of time. he gap conductance used for this calculation was 669 Btu /h-f t 2 *F. he fuel and cladding temperature is not shown since there was no increase in these parameters, because the DNBR for this accident did not go below the criterion value of 1.3. It is therefore concluded that no fuel damage vill occur.

An analysis has been performed for the locked rotor accident with the assumptions presented in Table 3.4-1. The power distribution was assumed to be a 1.5 cosine with a power spike located as shown in Figure 3.4-2. Figure 3.4-6 shows the power, flow, and cal-culated core average heat flux fractions. We pressure was ass [metd to be constant at 2135 psig. We initial power level for this accident was 102% of 2568 MVt. Trip occur:. at about 0.9 second. Figure 3.4-7 shows i the maximum fuel temperature versus time. %e fuel temperature is af-fected very littic since the power rises only slightly. Figure 3.4-7 I also shows the seximum cladding temperature and the DNS ratio. It is I' seen that the DNBR reaches the criterion value of 1.3 at about 0.9 second af ter which the cladding temperature increases to a value of 139&F which cccurs 4.0 seconds af ter the initiation of the accident.

3-23

~

3.4.2. LOCA Analysis 3.4.2.1. Introduction The maximu:a al.levable linear heat generation rate for a typical B&W rodded plant acccu= ting for fuel densificacion is established in previous fuel densification reports and in BAW-10G79,

" Operational Para =eters for B&W Rodded Plants," which forms the basis for this section of the report.

The ef fectiveness of the emergency core cooling system (ECCS) for B&W's 177-fuel assembly, vent valve plants during a postulated LOCA was evaluated as specifie-d in Part 4, Appendix A of the AEC Interim Policy Statement. Calculations were made by using the CRAFI computer code during the blowdown period, the REFLOOD code during the vessel refill portion of the transient, and the THETA 1-B code for g

the fuel rod heatup. The results of these analyses and the general as l methods and assumptions used in B&W's evaluation model are reported in E

l topical report BAW-10034 Rev 3, and in the respective applicant's E FdARs. Both analyses were performed without assumed fuel densiiication effects._

3.4.2.2. Ef fects of Fuel Densification The LOCA analyses established the 8.5!-f t2 split in the cold leg pipe at the pump dishcarge as the break size and lo-g cation resulting in the highest cal $:ulated cladding temperature. The E l 1

l consequences of this design basis accident (DBA) with the added restric-tions imposed by the postulated fuel densification phenomena have been

)

, investigated. Three of the riost influential restrictions are as follows:

1. Power spikes assumed to occur in gaps between fuel pellets.

2. Increase in the average linear heat rate l due to the assumed reduction in the fuel i pellet stack height.

l

3. A 25% reduction in B&W's fuel pellet gap conductance model as specified by the AEC's preliminary evaluation of the analytical method.

1 3-24 Babcock s.Wilcox g l

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These restrictions, when incorporated in the B&W evaluation acdel, increase the core average fuel temperature at the otart of the LOCA analysis; however, in the earlier analysis (BAW-10034,

' Revision 3), for conservative purposes, a higher initial core tempera-ture was used rather than the value that resulted from fuel densification.

The limiting break size and location does not change due to fuel densi-fication eifects.

When the cladding temperature response for the DBA was calculated, the restrictions due to fuel densification were incorporat ed into B&W's evaluation model, and a maximum linear heat

~

rate was calculated for which a peak cladding temperature of 2300F re-culted. Initially, the flux shape, resulting from the design power maneuver for each plant, was used to establish the maximum allowable heat rate. This transient had the largest peaking factors at any time i

in life. In this analysis, an equivalent radial multiplier was applied l over the entire length of the pin instead of imposing a powcr spike only at the location of the peak axial power. This procedure leads to a conservative evaluation of the peak cladding temperature.

However, the results presented in the fual den-sification reports before preparation of the Crystal River 3 report, were calculated by assuming a negative moderator coefficient. Con-sistent with the ana' lyses and method presented in BAW-10079, this repor:

uses a zero moderator coef ficient. The sensitivity of the naximum al-lowable heat rate (LOCA limit) to this parameter was studied in BAW-10079, for Oconee 2, which is very similar to Oconee 3, and is presented in Figure 3.4-8. (For additional information, see SAW-10079, section i

1 2.2.)

To further demonstrate the safe full-power operation of B&W nuclear plants, the sensitivity of the LOCA limit to i

the axial position of the power peak was also investigated in BAU-10079.

l This study utilized a zero moderator coefficient and an axial power l peaking factor of 1.7 at various points from an elevation of 4 to 10 feet. This peaking factor was conservative due to operating restric-tions placed on B&W reactors, which preclude the existence of peaking factors of this magnitude. For additional conservatism, the most con-servative dimensions were used to deternine the stored energy values .

3-25 Babcock s,Wilcox

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used in L:.e calculations. The results of this analysis are shown graph-ically in Figure 3.4-9. The calculations showed that the allowable heat rate is essentially constart up to the 8-foot elevation. Beyond this elevation, a gradual decrease is observed owing to the degrace., beat transfer during the reflood portion of the LOCA.

The locus of points generated by this analysis defines the allouable heat rate versus axial position at rated power for Oconee 2 and ensures tha t the LOCA criteria specified in the in-terim policy statement are met.

Calculations conducted for Oconee 3 ensure that the LOCA limits in Figure 3.4-9 (generated for Oconee 2) are both adequate and conservative for Oconee 3.

I Table 3.4-1 7 enal Data Input for Safety Analysis Active fuel length, in. 139.9 Fuel pellet diameter, in. 0.365 Fuel cladding thickne.:9, in. 0.0265 Cap coefficient, Btu /h-ft 2 _.F 669 Filo coefficient Variable (a) llot channel factors Overall power factor (F q) 1.0107 Local heat flux factor (F") 1.0137 0.98 Flow area reduction factor Assu= sed DNB 1.30 DNB correlation used W-3 Errors T - inlet , F +2 I Fressure, psi -65 Flux trip setpoint, % +6.5 (a)After a DNBR of 1.3, the Bishop, Sandburg, Tong correlations were used for both transition and film boiling.

B B

3-26 , Babcock ta Wilcox s _.-

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Figure 3.4-1. Pressure. Power, and Flux Vs Time for Densified Fuc1, Rod "*?ction Accident 8.0 Susten Flav = Constant 7.0 -l 2380 E

l 3= 2360 s

- 6.0 . e -

2340

- _=_

2320 m

<= 2 Pressure

] 5.0 .

2300 .S 2

2 2280 ._.a j g Poser 4.0 - -

2260 e =

n 3 Initial Core Avg *a

, Heat Flux = 176,470 2240 yf l

C 3.0 Btu /hr-ft 2 -

2220 *5 0 -

2200 E3

= 0>

2.0 -

2180 $

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[

Co.e Avg Heat Flur 2160 2I40 E f ,

1.0 2120 2100 0 1 2 3 4 5 Time. Seconds 3-27 Babcock a VIllcox

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=

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W  ::

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I 8 8 9 g g.

8

• R .

. It!IV(yg)

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3., -- g

Figure 3.4-3. DhBR, Fuel and Cladding Temperatures, and Film Coefficient Vs Time for Rod F.joction Accident

• 1600 Fuel !! cit ON 1400 - 60.000

[ BR N 1.8 -

5000

\  %

• I g j

/ I L p

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U f = R i.s _ . 4:00

\ / ctaa iturtutunt n a y  :

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=.

= / m O  % -

1000 " - 40.000 g

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= = .

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e l

3

/ \ _ i00j_ o.0003

- =

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i 1.2 -

j4400 -

600 =

G FILN COEFF @ NAllNUM CLAD TEMPERATURE -

500 -

500 W i i i i i g i.0 4200 I 400

, ,. 0 0.4 0.. i. i.. 2.0 2.4 Tiet. Sec M

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I Figure 3.4-4. Power, Flow, and Flux Vs Time for Densi-i fled Fuel, Four-Pump Coastdown I  !

! 1.0 initial Core Avg O.9 Flow -

3

-HeatFluxe102Powg 3

= 180,000 Btu /ht-ft Power 1.00 e 0.8 .._,'~~%

e  %

3 Core Avg Heat Flux ' N ,

O 0.7 -

\ -

.98 l E i

C .E g j 0.6 .

.96 j

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% 5 5 0.5 -

\ -

.94 C

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\ I 0.4 \- .92 I

0.3 -

- ,g0 0 0.4 0.8 1.2 1.6 2.0 Time, Seconds I

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l Figure 3.4-5. DNE2 and Film Coef ficient Vs Time for Densified j Fuel. Four-Pump Ceastdown 98.065 _

88 Cl:0 -

l t

i e,

~ i r 78.002 - ,

=

$ / - 1.s l

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4 1.8 g 58.CCD -

084 8 FILE

( ~

CEIFICIENT l

$ - i.7 l -

2 SD.CDD _

1.8 "

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" 40.000 _ g,4 l

=

=

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=

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3_31 Babcock & Wilcox

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I Figure 3.4-6. Power, Flow, and Flux Vs Time for Densified Fuel, Locked Ector Accident 1.2 -

Initial Core Avg Heat Flux e 102 Poser

= 180,000 Btu /hr.ft 2 g 1. 0 ,_

j _

,N f oreC Avg Heat Flux 0.8 -

pga, s l 3 N

. 0. 6 -

y _

Power E

30.4 -

E -

l 0.2 -

0.0 e i . . .

0 1.0 2.0 3.0 4.0 5.0 Time, Seconds I

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3-32 Babcock a,Wilcox

_ m -s -__ . ._y .- -y

Figure 3.4-7. Cladding and Fuel Temperatures and DNBR Vs Time for Densified Fuel, Locked Rotor Accident 4500

1. "~~

E a \

! # \

h h \ -

4300 s . N 5  %

=

g

\ [ Clad Temperaturs - 4100 3 1400 -

^

3  % _

y 1.s - \ .

O I.7 - 1200 g """

\

\ - 3900 *-

{

\  ;;

\ Fuel Temperaturs 1.8 N

I g -

3780 j 1.5 ' 1000 -

\

1.4 - N

\

m \ -

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' S' l.3 - 800 -

\

b \

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9 1 1 1

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=

a rb- c o % E % a' -

s -

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Figure 3.4-9. Axially Dependent Linver Heat Rate 20 18 k .

2 w a 16 5

=c 2

$ 14 E

12 6 8 10 12 E' 0 2 4 Fr Axial I.ocation of Peak Power f rom Bottom of Core, f t g

P h

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3.5, Mechanical Analysis of Oconee 3 Fuel 3.5.1. Cladding Collapse Results

1. Predicted time-to-collapse [ ] efph.

3.5.2. Cladding Stress Results

1. Table 3.5-1 lists maximum cladding circumferential stress cal-culated at various times in life. In no case does stress ex- g cced yield. 5

2. Cumulative fatigue damage after three cycles <0.9.

3.5.3. Fuel Pellet Irradiation Svelling [

Results (,

1. Circumferential plastic strain is less than 1% at EOL. 5  ;

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Table 3.5-1. cladding Circumferential Seress p p Densified Yield Ultinate ut , int' total, strength, strength, Tclad, F psi Casa psia pais pei psi Beginning of life preop- ,

orational hot standby - 2200 460 532 -22,500 48,000 57,000 0% power 2500 460 532 -27,600 48,000 57,000 Beginning of life-void section of cladding - 2200 580 650 -20,800 45,000 50,000 100% power 2500 $80 650 -25,700 45,000 50,000 t

Beginning of life-void section of cladding - 2200 600 650 -20,500 45,000 50,000 114% power 2500 600 650 -25,400 45,000 50,000 i

[ Beginning of life-fueled

" section of cladding - 2200 580 723 -24,900 42,000 44,000 100% pwar 2500 580 723 -30,000 42,000 . 44,000 Beginning of life-fueled section of cladding - 2200 600 733 -75,100 41,500 43,500 114% power 2500 600 733 -10,200 41,500 43,500 End of life-hot standby - 2200 460 532 -22,500 48,000 57,000 ,

0% power 2500 460 532 -27,600 48,000 57,000 End of life-fueled section 2200 580 704 -23,800 ' 43,000 46,000 of cladding-100% power 2500 580 704 -28,900 43,000 46,000 N 711 -23,900 43,000 46,000 End of life-fueled section 2200 600 46,000 of cladding - 114% power 2500 600 711 -28,900 43,000 x End of life-Immediately 2200 460 535 -22,800 48,000 57,000 8' af ter shutdown 2500 460 535 -27,800 48,000 57,000

$ End of life, cladding temp g of 425F 1725 400 425 -16,300 50,000 62,500 g .

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APPENDIX Design Parameters for Oconee Unit 3 i

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1 A-1 BabC0Ck 8. WilCOX l

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1. Core Operating Conditions

a. The reactor vessel inlet temperatures are as follows*

E Nominal, F_ Maximum, F_

100% power 554* 556 1141 power 550.6 552.6 i b. The nominal outlet pressure is 2200* psia, and the minimum out-let pressure is 2135 psia.

2. Core Design

! 2.1. Fuel Assembly Information

a. There are 177* fuel ass,emblies in the core.

b. There are 208* fuel rods per assembly with an outside di- g ameter of 0.430* inch and an inside diameter of 0.377* 3 inch.

1 i c. There are l'6* control rod guide tubes per assembly with dimensions of 0.530* inch OD = 0.016* inch wall thickness and one instrument tube per assembly with dimensions of 0.493* Inch OD = 0.441* inch ID.

d. The fuei rod pitch is 0.568* Anch.

2.2. Fuel The undensified active fuel length is 144* inches.

a.

b. The, active length of the fuel with densification is 139.94 I l inches.

c. The cladding is Zircaloy-4 (cold worked) with a thickness of 0.0265 inch.

The undensified pellet is 0.370* inch diameter and 0.700 f d.

inch long.

f *

e. Unit 3 core 1 is 92.5 of theoretical density (specified).

• Values given in the Oconee 3 FSAR.

I babcock & WilC01 A-2 I

l

3. Power Distribution

a. The design cora radial power map is shown in Figure A-1.

b. he maximum fuel assembly local rod power peaking distcibution is shown in Figure A-2.

c. The percentage of power generated in the fuel is 97.31*.

J. ne percentage of power generated in non-fuel regions is 2.71*.

4. Fluid Flow

a. Coolant Flows and Mass Velocities: l

)

ona l Vent vent i valves valve j closed open  ;

Total reactor vessel coolant flow. 131.32* 132.60 106 lbm/h Ef fective core coolant flow 106 124.23* 118 52 lbmih Average mass velocity at core 2.53 2.41 intet. 106 1bm/h-ft2 Inlet nass velocity to hot 2.235 2.13 assembly, 106 lbm/h-ft 2

b. De core flow area (effective for heat transfer) is 49.19* ft2,

5. Hot Channel Factorr.

a. ne hot channel factor on average pin power (F ) is 1.011.* It is applied on the enthalpy rise for the entire ch el. The hot channel factor on local surface heat flux (F") is 1.014.* This 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.* n is value is applied over the entire length of the channel.

i

• Values given in the Oconee 3 FSAR.

1 A-3 Babccck 8,Wilcox

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c. Flow is reduced in the hot bundle by a flow maldistribution factor, which is 951* of the nominal isothermal bt.adle flow.

d. The energy mixing coefficient (a) is 0.02.*

6. Core Peakinst Conditions ne 1.5 cosine, synenetrical axial power shape of the reference design was used as a ~oase case to determine whether other axial power shapes in any way magnified the variation in DNBR. A 1.78 radial-local nuclear peak-ing f actor (Fah) associated with a 1.5 cosine axial flux shape establishes the maximum desistn condition resulting in the 1.71 DNBR at 114% of 2568 The results indicate that outlet peaks with the spike show an overall latger degradation in DNBR than does the densified 1.5 cosine axial power shape and its associated power spike. B&W utilized a conservative 1.83 l

(P/P) outlet axial power shape in conjunction with a 1.49 (P/F) radial-local peak to maintain the rcference design IEBR of 1.55 at 114% of g 2568 HWt. 5 This set of peaking conditions maximizes the DNBR penalty associ-ated with fuel densification and prevents the need to re-evaluate all DNBR data for the power / imbalance / flow trip system. The penalty de-termined in this manner was used to modify the power / imbalance / flow sys-tem as indicated in section 3.3.4. The 1.83 (P[P) outlet axial power shape enc,vn in Figure A-4 is precluded during normal operation as de-scribed in the technical specifications and as such is not a design cri-terion.

! The 1.5 axial power shape, in conjunction with a 1.783 radial shape peaking combination, is used for transient and accident analyses.

nis particular shape results in a more conservstive DNBR than any other shape existing during normal operation. This shape is shown in Figure A-3.

For LOCA analysis, the design basis axial power shape was a 1.816 peaking at a distance of 1.0 feet below the core midplane. This shape and peak in conjunction with the calculated radial factor, are most conservative for the 14CA peak cladding temperature analysis aad Values given in the Oconee 3 FSAR.

A-4 Babcock 8 WilCOX

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could occur acaentarily during the period of xanon undershoot following o design basis (100-30-1002) transient. He peaking factor and the associated radial factor are within tha DNB2 limiting criteria statement given in the previous paragraph. The reason is that in IDCA analysis, the laportant parameter is peak cladding temperature, whereas for DNBR protection, the important parameter is not only heat flux and flux shape, but also the integration of heat input up the channel and the resultant enthalpy rise.

The non-densified DNBR at design overpower is 1.55 With densifi-cation and the spike utilizing the 1.83 axial power shape, the DNBR is 1.48. ne reduction in overpower limit given in section 3.3.4 increased the 1.48 DNBR to the design value of 1.55.

7. Heat Flux Conditions The following data are based on the peaking conditions above so that a meaningful comparison between non-densified and denaified fuel can be made.

7.1. Non-Densified Conditions

a. The heat transfer surface area per fuel pin is 1.3509 f t2, I

b. The average heat flux (q")* is 171,470 Btu /h-ft2 ,  ;

l

c. The maximum heat flux at minimum DNBR is 457,774 Btu /h-ft2,

[q" (MDNBR) = { x 1.55 x 1.49 x 1.14 x 1.014].

Axial (P/F) at MDNBR = 1.55.

(P/P) radial-local = 1.49.

Max overpower = 114% of 2568 MWt*.

Hot channel factor on local surface heat flux = 1.014*.

d. Le average power density in the core is 83.38 kW/ liter, and the average linear heat rate is 5.66 kW/f t.

I e. The naximum surfa :e temperature at the exterior of the

! cladding at 100% power is 650F for a pressure of 2135 psia.

e Values given in the Oconee 3 FSAR.

3_3 Babcock & Wilcox l

I 7.2. Densified conditions

a. The hast transfer surface area per fuel pin is 1.3128 f t 2,

b. The average heat flux is 176,446 Etu/h-ft2, .

c. The maximus; heat flux at utnf == DEBR is 483.213 Btu /h-ft2:

x 1.59 x 1,47 x 1.14 x 1.014].

[q," (MDNER) = q,'

Axial (P[P) at HDNBR with power spike = 1.59.

(P/P) radial local = 1.49.

Max overpower = 114% of 2568 Wt .

Hot channel factor on local surface heat flux = 1.014*.

I

d. Average volumetric power density in the core la 83.38 W/

liter, and the average linear heat rate is 5.82 W/f t. This E l

assumes that all fuel pins have the densified active length, which is conservative. 5 a.. The maximum surface temperature at the exterior of the clad-ding at 100% power is 650F for a pressure of 2135 psia.

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