ML19322C164

Cycle 3 Reload Rept

ML19322C164
Person / Time
Site:	Oconee
Issue date:	08/31/1977
From:	BABCOCK & WILCOX CO.
To:
References
BAW-1453, NUDOCS 8001090573
Download: ML19322C164 (50)

Text

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G* NEE *5*! 1 CYCLE 3

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D.ETUi.:I Ta ".'c'.'.

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' BA3CCC 1 WILCCX Fower Centration Group Nuclar Pcver U nerr.t ton Div 'sion i

P. c. Box 1260 Lynchburg,..' a rginia 24505 Babcock s. Wilcox 8001090 D 3 f

i CONTENTS Page 1-1 1.

1*.TF'IRC11oM AND SMARY.

~..............

l 2.

>Pr.A 1:;4 itI STORY 2-1 3.

SEP.AL *]ESCRIFT.~A 3-1 4

TrEL SYST M DES!CN.

4-1 4.1.

ch.1 Asscably Mechanical Design.

4-1 4.2.

1%.1 Rod Destan.

4-1 4.2.1.

Cladding Collapse 4-1 4.2.2.

Cladding Stress.

4-2 4.2.3.

Cladding Strain.

4-2 4.3.

Thernal Design 4-2 4.3.1.

Power Sptke Model.

4-2

4. 3. 2.

Fuel Temperature Analysis.

4-1 a.4 Material Design.

-)

4-3 4.).

Operating Experience 5.

Nt' CLEAR DESIGN.

5-1 5.1.

Physics Characteristics.

5-1 5-2 3.2.

Analytical input

......~...............

5.3.

Changes in Nuclear Design.

5-2 n.

THERMAL-KYDRAL*LIC OES1CN.

6-1 1

1 e.l.

Evaluation 6-1 e.2.

DSBR Analysts.

5-1 6.3.

Pressure-Te=perature LLait An'alysis.

6-3 o.a.

Flux / Flow Trip Setpoint Analysis 6-3 7

ACCICENT AND TRANSIENT ANALYSIS 7-1 7-1 i

7.1.

General Safety Analysis.

7.2.

Rod 'Jithdrawal Accidents 7-1 I

7. 3.

Moderator Dilution Accident 7-2 7.4 Cold Water Accident 7-3 i

7.5.

Loss ot' Coolant Flow 7-3 7.6 Stuck-Out. Stuck-In, or Dropped Rad Ac-ident 7-4 7.7.

Loss of Electric Power 7-4 7.3.

Secas Line Tailure 7-5 7.9.

Steam Generator Tube Failu're 7-5 7.10.

Fuct llandling Accident 7-5 7.11.

Rod Election Accident 7-5 i

7.12.

Maximum Hypothetical Accident 7-6 g

g w-a

n1N TIN -x icent*J)

Paac 7-9 7.13.

Vaste cas Tank Rupture 7-6 7.14 LCCA Analysis..............,........

8-1 6.

PROPOSED MCDiFICATIONS TO TECHNICAL SPECITICATIONS.

9-1 9.

START *JP PROGRAM - FHYSICS TES!!NG A-1 RETF.RENCES............................

Li,t of Tables Table 4-4 4-1.

Fuct Design Parameters and Dimensions.

4-5 4-2.

Fuel Thermal Analysis Parameters 5-3 5-1.

Oconee 3. Cycle 2 and 3 Phystes Parameters 5-4 5-2.

Shutdown Margin Calculation for Geonee 3. Cycle 3...

5-3.

Comparison or Fuel Melt Margins t or Selectively Loaded 5-5 Assenblies and Limiting Assembly in Core 6-4 6-1.

Cycle 2 and 3 Maxicus D= sign Cendit tons.

7-7 7-1.

Cocparison of Key Parameters ter Accident Analysis L!st of Fleures Figure 3-2 3-1.

Core Leading Diagram f or Oconee 3. Cycle 3 3-2.

Enrichsent and Burnup Distribution for oconee 3. Cycle 3 3-3 3-3.

Control Red Locations fer Oconee 3. Cycle 3..........

3-4

-1.

Maximum Cap Size Vs Axial Posi:ien - Ocence 3. Cycle 3 4-6 4-7 4-2.

Power Spike Factor Vs Axial Position -- Oconee 3. Cycle 3 5-1.

DOC (4 EFPD) Cycle 3 Two-Divensional Relative Pcwer Distribution -. Full Power. Equilibrium Xenon Normal

$-o Rod Positions. Groups 7 and 8 Inserted S-1.

Core Protection Safety Limits -. 0conee 3. Cycle 3.......

8-2 8-2.

Core Protection Safety LLsits - Oconee 3. Cycle 3 8-3 8-3.

Core Protection Safety Limits - Geocee 3. Cycle 3 8-4 6-4.

Protective Systes Maximus Allowable Setpoints -

8-5 Oconee 3. Cycle 3.......................

S-5.

Protective System Maximum Allowable Setpoint s --

8-6 oc on e e 3. C y c le 3.......................

s Babcock s Wilcox

I Figurce (Cont'd)

F i gu re-Page 5-6.

Rod Post t ion Limits for Four-Pizzy Operat ion from 0 to 100 : 10 EFPD - Oconee 3. Cycle 3 d-7 8-7.

Rod Posi tion Limits f or Four-Pump Operation From 100

10 to 235 : 10 EFPD - Oconee 3. Cvele 3...

8-8 3-8 Rod Position Limits for Four-Pusp Operation Af ter 235 : 10 EFPD - oconee 3. Cycle 3..

6-9 4-9.

Rod Pomition LLmits for Two-. sad Three-Pump Operation From 0 to 100 : 10 EFPD - Oconee 3. Cycle 3....

'.- 10 3-10.

Rod Posi tion Ltsits f or Tvo-and Three-Pump Operation Fres 100 : 10 to 235 : 10 EFPD - Ocence 3. Cycle -3 3-11 9-11.

Rod Position Limits for Two-and Three-Pump Operation Af ter 235 10 EFPD.

5-12 8-12 Operational Fower Imbalance Envelope for Operation s'rne o to 100 : 10 EFPD - Oconee 3. Cycle 3.....

A-13 8-13.

Operational Power Imbalance Envelope for Operation From 100 10 to 235 : 10 EFPD - Oconee 3. Cycle 3 5-14 s-14 Operational Power Isbalance Envelope for Operation After 235 2 10 EFFD - Oconee 3. Cycle 3..

8-15 S-15.

APSR Position LLatts for Operation Frse O to 100 : 10 EFPD - Oconee 3. Cycle 3 8-16 8-16.

APSR Position Limits f or Operation From 100 10 to 235 : 10 EFPD -- Oconee 3. Cycle 3....

8-17 3-17.

APSR Position LLatts for Operation Af ter 235 : 10 EFPD - Oconee 3. C/cle 3 8-16' l

i

-v-Babcock & Wilcox i

1.

INTRODUCTION AND

SUMMARY

This report justif ies the operation of the third cycle of Oconee Nuclear Sta-tion Unit 3 at the rated core power of 2508 MWt.

Included are the required inalyses as outlined in the USNRC document " Guidance for Proposed Licenne I

Asendnents Relating to Refueling," June 1975.

To support cycle 3 operation of oconee Unit 3 this report employs analytical techniques and design bases established in reports that were previously sub-nitted and accepted by the USSRC and its predecessor (see ref erences).

A brief summary of cycle 2 and 3 reactor parameters related to power capability is included in nect ic 1 5 of this report. All of the accidents analyzed in the FSAR have been reviewed for cycle 3 operation.

In t hose cases where cycle 3 j

characteristics proved to be conservative with respect to those analyzed for evele 2. no new analy 4es were perf orted.

The !cchnical Specifications have been reviewed, and the nodifications required for cyei: 3 operation are justified in this report.

Based on the analyses perforsed, which take into account the postulated ef fects of fuel densification and the Final Acceptance Criteria for Emergency Core Cool-

]

ing Systess, it has been concluded that Oconee Unic 3. Cycle 3 can be safely I

operated at the rated ~over level of 2568 MWt.

Babcock s.Wilcox g.g

2.

optga ;NG HISTORY

• he ro t erence f uct cycle tot the n aclear and ther.al-hydraulic analyses Jf 2hc Oc..nce Nuclear Station, Unit 3. is the currently opetating cycle 2.

Cycle a was terainated 4f ter 478 ETPD at operation. Cycle 2 Jehieved initial crit-icality on Novesber 7, 1976, and power escalation connenced on November 10, 1976. The 100: power level of 2568 MWs was reached on November 21. 1976.

The fuel eycle design length is 282 EFFD. No operating anomalies occurred during cycle 2 operation that would adversely af f ect fuel performance in cy-cle 3.

1

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i BLANK FRAME FOR PROPER PAGINATION

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GE*.~ ERA 1. DESCR1PTICN

't e Oconce Lnit ) reactor core is described ir Jetati in Cnaster 3 of the Unit 1 FSAR.i The cycle 3 core consist s of 17 7 f uel a m.cso lics, eacn of knten is a 15 by 15 arts < cent.itnant 209 fuel rods. 1% contr31 raf guide tubes. -and one i n e n r e, instru= ant guide tube.

• he f uel rod cladair.4 1s co:J-worked Zircalcy-a with an 03 of 0.430 inch and a uall thickness of 0.026) incn. The f uel con-

.a ts of diabed-end, cyt tadrical pellets of uran tum Jaotide swteh are 0.170 inch in 4tameter.

(See Ta51e 4-1 for additional 1sta.)

All feel assemblies in cycle ) a41ntain 4 constant nostnal fuel loading sf 463.9 k4 of uranium.

The undensified nominal active fuel lengths and enecretical densities vary be-tween batches. however. and these values are given in Table a-1.

Figure 3-1 is the core loading diagram for Oconee ),

cycle 3.

All of the bet en 2 assenolles will be discharged ~4t the end of sycle 2.

Five once-i>urneu hatch 1 assenolles. with an initial enrichment or 2.01 we : 2 Ji'J. will be re-toaded inta t he central portien of the core. Batches 3 ", and 4A - with int-e t41 enrichsents of 3.00. 2.53. and 2.64 we

• 2 35. respectively will be shuffled to new loca t ion s.

Batch 5. with an initial enrichment of 3.02 we *

• 5 '. will occupy primartly the core pertphery and etaht interior locattuns.

L Figure 12 is an eighth-core map snowing the assensly burnup and enrichment

~

distrit ution 4t the beginning of cycle 3.

Reactivity control is supplied by 61 full-length A4-in-Cd centrol rods and soluble boron shim.

In addition to the f ull-length control reds. etaht par-t tal-length axial power shaping rods (A?SRs) are provided for additional con-i t rol of ax141 power dist** 5ution. The cycle 3 locations of the 69 control rods and the group designations are indicated in Figure 3-3.

The core loca-i tions of the total pattern (69 centrol rods) for cycle 3 are identical to i

those of the reference cycle indicated in Chapter 3 of tne ISAR.1 Hewever.

i the group designations dif fer between evele 3 and the reference evele to anni-Once power peaking. Neither control rod interchange nor burnable poison rods are necessary for cycle ).

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. 01 3.00 2.C1 3.00 3.02 3.30 3.00 3.02 3

16.215 26.640 14.215 23.eS3 0

20.423 23.483 0

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18.391 4.718 27.527 7.557 3.00 2.53 2.53 3.A 3.02 3.a2 1.

20.626 6.194 11.043 21.)$8 0

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4 FtT.I. SYSTEM DESIr.3 L.I.

Fine ' Asse-hiy Mechanical Destatn The types of ft.et assemblies and pertinent feel design parameters and di en-sions f or Oconce 3, cycle 3 are 114ted in Table.-l.

The f resh f uel assen-blies (batch 5) incorporate sinor des.c. sodif icationr. to the spacer 4rit corner cells to reduce spacer Arid interactic.t during handling. In addition, lasroved test methods (dynamic i= pact testing) snow that the spacer < rids have a higher. seismic capability and thus an increased saf ety sargin over the val-use reported in reference 3.

All other results, references, and identified conservatisna presented in the 2 (section 4.1) are applicable to the cycle 3 previcus Oconce 3 reload report reload core.

6.2.

Fuel Rod Design

. 2.1.

Cladding Collapse Creep collapse analyses were performed for three-cycle ass ssly power his-tories.

Batches 3 and 4 were analyzed using aa-Nilt data.

The baten 3 f uel is more limiting for cladding collapse due to.ts previous incore exposure i

time.

The assembly power history for the most limiting assembly was used to calcu-late the f ast neutron flux level for the energy ran;;e above 1 MeV.

.h col-lap se t ime f or the :ros t 11 string assembly was coeservatively decemined to be more than 30,000 EFPrt (effective full-power hours), which is longer enan ene saxistas three-cycle design lives (Table 4-1).

M creep collapse analyses were performed based on the conditions set forts :.n references 2 and 4 i

n.s...... u:e..

. 2.2.

Claddine Stress 3 stress parameters are enveloped bv a conservative f uel rod stress The 0cence be less For design evaluation, the prtmary membrane stress must analysia.

than two-thirds of the statsum specified unitradiated yield strencth, and all be less than the mint =um specified unitradiated yield strength.

stresses must In all cases, the marctn is in excess of 3G. The following conservatisms to Ocence 3 f ac1 were used in the analysis:

with respect 1.

A lower post-denstf ication internal pressure.

2.

lower initial pel;et density.

3.

A hiAner system pressure.

4 A hig5er thermal gradient across the cladding.

a.2.1.

Claddine Strain The f uel design criteria specify a LLait of 1.0% on cladding plastic circum-forential strain. The pellet design is established for plastic cladding strain of less than 12 at values ei saximum design local pellet burnup and heat generation rate which are considerably Sighee than the values the Oconee 3 f uel is expected to see.

This will reu It in an even greater margin than the analysis demonstrated. The strain analysis is also based on the maximum Specif ication value f or the f uel pellet diameter and density and the lowest permitted Specificatten tolerance for the cladding ID.

4.3.

Thermal Design All f uel assemblies in this core are thermally similar. The fresh baten 5 differences in fael inserted for cycle 3 operation introduces no significant f uel thermal performance relative to the other fuel remaining in the core.

as shown The design minimum linear heat rate (LHK) capability is 20.15 kW/f t, in Table 4-2.

LER capabilities are based on centerline fuel melt and were established using the TAFT-3 code 5 with f uel densification to 96.5% of theo-retical density.

Power Spike Model (Densif'ication)_

4.3.1.

i The power spike model used for cycle 3 analysis is the same as that used for i

cycle 2.2 Figures 4-1 and 4-2 show the maximum gap size and power spike fac-tor, respectively, versus 4xial position. The power spike f actor and gap size were cased on unitradiated batch 4 and 5 fuel (94.0 TD) with an assumed Rahr.ock & Wilcox 1

enrichment of 3.0 wt : 2 15U.

These values are conser.atively high f or batch I and 3 tue!.

4.1.2.

Fue! Temperature Analvsis T;-rmal analvsis at the f uel rods assumed in-reactor densification to 96.5%

theoretical density. The analytical methods utilized are the same as those documented in references 2 and 6 fcr cycle 2.

The average :uel temperatures ahown in Table 4-2 are taken t' rom the analyses used to det ine the LHR cana-bility for the iuel.- ** These analyses were based on the lower tolerance 11=-

it of the spect:teation f uel density and

..;ned isotropic diametral shr takage and anisotropia axial shrinkage tconsis s t.: reference 7) resulting from fuel densification.

4.4 M.teri.nl Design The ba tch 5 f uel 45 sblies are not new in concept, nor do they utilize dif-ferent component materials. Therefore, the chemical compatibility of all possible f uel-cladding-coolant assembly interactions for the batch 5 fuel as-semblies are identical to those of the present fuel.

4.5.

operating Empertence B&W's operating experience with the Mark S,15 by 15 fuel assembly design has verified the adequacy of this design. As of April 30, 1977, the following op-erating experience has been accumulated for the seven B&W 177-fuel assembly plants using the Mark 5 fuel assembly:

Max assembly Cumulative Current burnup,

net electrical Reactor evele Y.*d/stU output, MVh Oconee 1 3

25,400 16,742.549 Oconee 2 2

25,900 12,919,680 Oconee 3 2

23,400 12,130,627 TMI-1 2

26,200 13,306.C85 Arkansas One 2

20,700 9,826,476 Rancho Seco 1

15,400 6.040,979 Crystal River 3 1

1,000 575.364 Babcock s.Wilcox ta

4 Table 4-1.

ruel Design Para =eters and Di=ensions Twice-Once-burned FAs Fresh burned

FAs, Batch

FAS, batch 3 Batch 1 Batch 4 4A batch 5 FA type

. Mark 53 mrk B3 N rk B4 Mark B4 mrk B4 No. of FAs 60 5

52 4

56 Fuel rod OD. in.

0.430 0.430 0.430 0.430 0.430 Fuel rod 10, in.

0.377 0.377 0.377 9.377 C.377 Flex. spacers, type Spring SprLng Spring Spring Spring Rigid spacers, type Ir-4 2r-4 Zr-4 Ir-4 2r-4

'Jndensif active fuel 142.0 142.0 142.23 142.23 142.25 length (nominal), in.

I I

Fuel pellet initial 95.5 *I 95.5 "

94.0 94.0 94.0 density (nom),

• TD U

O.3680(*) 0.3695 0.3695 0.3695 Fuel pellet CD (mean 0.3680 specif), in.

Initial fuel enrich, 3.00 2.01 2.53 2.64 3.02 wt 2 85 :t SOC burnup (avg),

21,766 14,320 7,881 7,043 0

WJ/ntU Cladding collapse

>30,000

>30,000

>30,000

>30,000

>30,000 time, EFPH Design life, EFFH 24,888 18,120 20,928 20,928

>21.144 I*I Nominal values af ter resintering.

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Ta51e 4-2.

Fuel ~heru.a1 Analysis Para eters a

5.treh 1 Batch 3 Batch 4 Esteh 5 No. of assemblies 5

60 56 36 Initial density : *3 95.3 95.5 'I 94.0 94.0 I

Pellet diameter, in.

0.3682 0.3680 0.3695 0.3695 Stack height, in.

141.0 141.0(*}

142.2 142.2 Densified Fuel Para =eters' I j

Pellet diasetet, in.

0.3649 0.3e49 0.3646 0.3646 Fuel stack height, in.

140.2 140.3 140.5 1 0.5 Neminal LHR at 2568 5.80 5.80 5.80 5.80 MWt. kW/ft Avg f uel temp at nomi-1310 1305 1120 1320 na l LHR, F LHR capability (centet-20.15 20.15 20.15 20.15 line fuel melt), kW/ft (a) Nominal values af ter cesincering.

(b)Densification to 9e.

TD assumed.

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5.

NL' CLEAR OESICN 5.1.

Physics Cha ra c t e r i s t ic s Table 5-1 compares the core physics parameters of eve'c. 2 and 3; the values for both cycles were generated using PDQ07 S.ince the core has not vet reached an equilibrium cycle, dif ferences in core physics para =eters are to be expected between the cycles. The shorter cycle 2 will produce a smaller cycle dif fer-en:tal burnup than that for cycle 3.

The accumulated average core burnup will be higher in cycle 3 than in cycle 2 because of the presence of tne once-burned batch 1,

3. 4 and 4A fuel.

Figure 5-1 illustrates a representative relative power distribution for the beginning of the third cycle at full powce with equilibrium xenon and normal rod positions.

The critical boron concentrati)ns for cycle 3 are higher than for cy' le 2 be-c cause of a higher feed enrichment. different radial power distribution, etc.

As indteated in Table 5-2 the control rod worths are suf ficient to maintain the required shutdown margin. However, due to changes in isotopics and the radial flux distribution, the 50C hot, full-power control rod worths are gen-erally less than those for cycle 2.

The cycle 3 ejected rod worths are lower than those in cycle 2 for the same number of regulating banks inser:ed. It is difficult to compare values betw.en cycles or between rod patterns since nsitner the rod patterns from which the C/A is assumed to be ejected nor the icotopic distributions are identical. Calculated ejected rod worths and their 1

edherence to criteria are considered at all times in lif e and at all power levels in the development of the rod insertion limits presented in section 8.

Ths eaximum stuck rod worths for cycle 3 are less than those in cycle 2.

The adequacy of the shutdown nargin with cycle 3 stuck rod wortns is demonstrated in Table 5-2.

The following conservatisss were applied for the shutdown cal-culations:

1.

Poison asteria) depletion allowance.

2.

10% uncertainty on net rod wurth.

3.

Flux redistribution penalty.

o h,a,. ura...

Flux redistributton was accounted for since the shutdown analysis was cal a-lated usine a two-disensional nodel. The shutdown calculation at the enJ of evele 3 ts analyzed at approxisately 235 EFPD.

This is the latest time (; 10 days) in core life at wnich the transient bank is nearly fully inserted. After 235 EFP3 the transient bank will be almost fully withdrawn. thus increasing the available shutdown margin. The reference fuel cycle shutdown margin is presented in ref erence 2. Table 5-1.

• he cycle 3 power deficits f rom hot zero power to hot full power are sizilar to but slichtly higher than those for cycle 2.

Doppler coef ficients, coderat-or coef ficients, and xenon worths are similar for the two cycles. The differ-ential boron worths for cycle 3 are lower than for cycle 2 due to depletion of the f uel and the associatej buildup of fission products. The effective delayed neutron fractions for born cycles show a decrease with burnup.

5.2.

Ana'veical Input The cycle 3 incore measurement calculation constants used to compute core power d.istributions were prepared in the same manOer as for the reference cycle.

)

5. 3.

Changes in Nuc1 car Design The same calculational =ethods and design information were used to obtain the

=portant nuclear design parameters for cycles 2 and 3.

In addition, there are no significant operational procedure change s f rom the reference cycle with regard to axial or radial power shape control. xenon control, or tilt control.

The operational limits (Technical Specification changes) for the reload cycle are shown in secticn 8.

A fuel seit limit of 20.15 kW/f t has been employed in calculating the reactor protection system setpoints and is tt.e same as in cycles 1 and 2.

The batch 5 f t 1 assemblies will be loaded as ahown in Figure 3-1.

Tuo batch 5 assemblies 1

have been assigned a maximum linear power rating of 19.74 kW/f t based on as-built data.

These assemblies will be placed in non-limiting locations during 1

their entire core residence. For cycle 3 investigation has determined that if these assemblies are placed in locations M-14 and E-2, they will not exper-1ence linear power rates higher than 19.15 kW/f t.

Thus, as shown in Table 5-3 for various tLmes during the nominal fuel cycle, the margis to fuel =elt will always be greater in these assemblics than in the most limiting assembly in the core.

I w

• h i.* 5-1.

Oconee ). Cvele 2 and 1 Phvstcs Para =eters

.s gele 2 Cycle 3 Cvcle length. EFP3 265 277 Cycle burnup..T4d/=tt 4293 8668 Average core burnup. EOC. M'4distU 18.160 13.921 initial core loading, stU B2.1 J2.1 Crit ical boron - BOC. p,na (no Xe)

I 1251 'I 1261 HZP. s;raup 8 inserted HZP. 4roups 7 and 8 inserted 1108 1138 HFP. groups 7 and 8 inserted 931 1000 Critical boron - EOC. ppe (eq Xe)

1. I HZP. group 8 37.5% wd. eq Xe 328 288 HFP. group d 37.5% vd. eq Xe 29 35 Control rod worths - HTP. 50C. I ak/k Group o 1.18 1.04 Croup 7 0.97 0.77 Croup 8 37.5 wd 0.54 0.40 Control rod worths - HFP. 235 EFP3

• ik/k Group 7 1.29 *I 1.05 f

Group 8 37.5% wd 0.49('I 0.44 Max ejected red worth - HIP Sk/k BOC groups 5-8 inserted 0.66(*)

0.73 235 EFPD. groups 5-8 inserted 0.60 0.$1 Max stuck rod worth - HZP. : ik/k BOC 2.30 2.54 I

235 EFPD 2.16 'I 2.24 Power deficit. H2P to HFP. I ak/k BOC 1.65 1.55 l

I'I l

235 ETPD 2.! t 1.98 Doppler coef f - Boc.10-5 (ak/k/*F) 100% power (0 Xe)

-1.54

-1.43 Doppler coeff - EOC, 10-5 (ak/k/*F) 100 peteer (eq Xe)

-1.54

-1.56 Moderator coeff - HTP. 10" (ak/k/*F) 80C (0 Xe. 1150 ppe. group 8 inserted)

-1.06 *I

-0.55 I

E0C (eq Xe. 17 ppe. group 3 inserted)

-2.39

-2.55 Boron worth - HFP. pps/ ak/k BOC (1050 ppm) 107(*)

105 EOC (17 pra) 101 95 i

l 1able 5-1.

(Cont'd)

Cycle 2 C y.- le 1 Xe n on wo r t h -- H TP, *. 4k / k BOC (4 davs) 2.64 2.66 EOC (equilibrium) 2.08 2.75 Ef f ective delayed neutron f ractien - HFP 30C 0.00585 0.00544 LJ2 0.00520 0.00522

'a)for eendition applicable to these values. refer to ENJ-1432.*.

Tt51e 5-2.

Shutdown Margin Calculation f or Ocence 3. cvele 3 50C.

EOC.

• Ak/k

• ak/k Avs11able Rod 'Jorth Total rod worth. HIP 8.74 8.70

'Jo r t h r e d ' n d ue to burnup of poison

-0.24

-0.31 Marfaum stuc4 rod. K2P

-2.54

-2.24 Net worth 5.96 6.15 Less 301 tacertainty

~0.60

-0.62 Total available worth 5.36 5.53 Required Red '* orth l

Power deficit. HTP to HZT 1.55 1.98 Mas allowable inserted rod worth 1.06 1.31 Flux redistribution 0.45 0.77 l

Total required worth 3.06 4.06 l

Shutdown Martin Total avail worth - total req'd worth 2.30 1.47 Required shutdown margin - 1.00: ak/k (4)For shutdown margin calculations, this is defined as A u

235 EFPD. the latest tLae in core lif e at which the transient l

bank is nearly fully inserted.

9D

D*1e 5-3.

f.coparison of fuel Melt.% rgics for Selectively leaded Asse=blies and Limitin.t Ass.cbly in Core selectively lasded Most l imit t in g ETP3 Loc 4 tion Marrin, 7.

Location _

Maratn, 4

M14 32.23 L14 29..?

100 M14 38.23 L14 35.23 200 M14 41.93 K9 36.la 277 M14 48.69 K9 38.31 I

l l

l l

l 1

1

\\

i l

l Rehench a Wileny

1. Leare 3-1.

EX ( *.

~~0i Cvele 1 Tve-31s nsional :clative Pc er

?i s t r i 1t :.. - Tu11 :'oser. Eq a t i t br t e Xenon. Oiarn.41 Fad P.,*sti.~

. Cr.)ups 7 4:1d 5 Inserted 9

4 10 11 12 Il 14 15 8

7 H

0.802 1.0C4

0. ii i 0.989

1. 07

0. d 4 7 0.439 3.e75 3

i K,

1.004 1.350 0.9e7 1.054 0.960 0.935 0.770 0.7%7 l

I

\\'l e

NI 0.'l54 0.9%7 0.eJa l

1.111 0.174 0.992 1.268 0.751 C.991 i

1.056 1..

1.337

1. 301

1. CSS 1.139 t

.I

1. 09

0. 462 0.978

1. 307 1.3s1

.1 7 0.813 i

i a

0.568 0.957 0.93 1.087 1.148 0.4;;

l 7

,o

0. 40 0.771 1.269

• 140 0.514 l

l l

A 0.676 0.7ta 0.751 0

l laserted FA Croup so.

0.000

' islative Power Denetty i

A 1

l l

6.

DiER.%1.-if?O7M1.1C DE5!'.N E v a l ua t ian The tr e rna l-hvd raul ie denten evaluation in suppor. of cycle 1 Ortratian uti-Itzed t.me methods and s dels described in ref erences 1. 2.

an d 9.

Cycle 3 an.nlysea r. ave been based on 109.5; of the (first sare) design reacter coolaa-.

'RCI system t iew rate.

Cycle 2 analyses; used 107.et of desien flew based oa a measured f low vilue of 110.00.

The reduced flow rate has been selected f ar vcle 3 analyses to provide consisteacy wita 0conee units 1 and 2.'**

The de-t'reases in RC f low used f ar these a,41yses are char ges in calculational parcs-eters only and do not represent changes in operat

• an of the plant.

Th. care conf igurat tar. f or cycle 3 differs slight v from tnat of eycle 2 in tnat the baten 2 f uel reaoved at the end of cvcit 2 is the Mark 31 fuel as-sen51v design, end the f resh baten 5 fuel insert. ' for cvele 3 is the Mark B-a s s eso l v design. Mark Sa assemblies dif fer f ren cae Marn S3 prisar tly in the design of the end fitting, which results in a slight reduction la flew re-sistance for the 54 design. No credit was taken in the analyses f or the

.n-creased slow to the Mark Se assemblies. located in the hottest core lecations, as a resnle of slight changes in the core flow dtstrt5ution or for the in-crease in system flow resulting f rom the reductice in t otal core pressure drop.

l n.2.

D43R Analysis

• he BL*-2 CHT correlation was used f or thermal-tvdraulic analysis et cycle Thia correlation, which has been reviewed and ap sroved for use with tr.e Mark i f uel assembly design.M has been used previously f or licensing of cycle 2 of the Ccocee 3 core.2 The ef f ect of f uel denstfication on minisus DNSR is pri=arily a resulc of the reductien in active :uel length, which increase.n the average he.st flux.

The cycle 3 DN5R analysis was based on a cold densified active length of 1.J.2 irches, a value selected to apply generically to 4 number of 394 plants.

This is a conservative nethod of applying the densif teat ten ef f ect minee all the

4 4

fuel assemblics in evcle I have longer densified lengths (Tabic 4-2) and be-cause no credit is taken fer ax:al ther=al expansion of the fuel colu=n.

This analysis differs frem that 0: evele 2 in two respects:

First, the effect of the densification pewer spike is no longer considered for DNER analysts based on iniornation presented in reterences 11, 12, and 13.

Second, the densified active !cngth is incorporated directly into the DNSR analysis, resulting in a calculated minimum DNBR of 1.40* at 112 power (Table 6-1). The cycle 2 anal-ysis had been based on a la.-inch active length with r.he effect of a reduced active length and the der.sification power spike calculated separately.

The' potential ef fect of fuel rod bow on DNBR can be considered by incorporat-ing suitable margins into DN3-1Lmited core aafety limits and RPS setpoints.

The sasimum rod bow magnitude would be calculated f rom the equation cb = 11.5

+ 0. 00 4 53, where 3 is tne rod bow magnitude (in sils) and BU is the burnup

( in F.*d. s t L*). The resultant DNBA penalty based on the maximum predicted as-l sembly burnup at the end of cycle 3 is approximately 6.C2.

However, since NRC review of this bow model had not been completed before the design of this reload core, the maximum rod bcw magnitude was calculated using the NRC intet-is model, 'C/Co = 0.065 + 0.001'49 /Bd. where SC is the rod bow magnitude (in r

mils) and Co is the initial gap.

The resultant DNBR penalty, based on the saxt:nn predicted assembly burnup at EOC 3 is 11.22.

I A t he rmal margin credit equivalent to 1: DNBR is available as a result of the flow area (pitch) reduction f actor included in all the thernal-hydraulic anal-yses to partially offset the projected fuel rod bow penalty. For the flux / flow trip setpoint analysis, an additional thermal margin credit equivalent to 2%

excess flow has been applied. The NRC Statf has accepted, en a plant-specific basis. the use of thermal margin credits resulting from RC system flow rates in excess of that assused for safety analyses.3 The 2% flow credit is claimed en the basis that 106.5% of design RC flow was used for safety analysis and an RC flew of 110% of design has been proven in the plant. For those analyses performed for previous cycles that are applicable to cycle 3 or future cycles,

(

credtt will be taken (as appropriate) for the removal of the densification i

power spike penalty. A note specific discussion of thermal margin credita is i

provided in sections 6.3 and 6.4.

6. 1.

Pressure-Temperature Limit Analvsis The pressure-temperature limit ce ves shown in Figure 3-3 provide the basis for the variable low-pressure tr1 > setpoint. The curves shown f or four-and three-pump operation each represent a locus cf points for which the calcu-lated minimum DNBR is equal to 1.30 (SrJ-2) plus the margin required to off-set an 11.2: DNBR reduction due to rod bow.

The specific credits used in this analysis to account for rod bow are as follows:

DNBR credit Credit for rod bow penaltv already 10 * **

=

included in analysis Credit for flow area reduction 1.0

=

factor in analysis credit for plant excess flow (3.5 None clat=ed available) l Total 11.2 6.4.

Flux / Flow Trip Setpoint Analvgis The flux / flow trip setpoint was determined by analyzing an assumed two-pump Loastdown scirting f rom an initial indicated power level of 102: plus flux seasurement and heat balance errors (equal to 103 f ull power in core). The analytical method was the sa=e as that used for licensing of cycle 22 with the following exceptions: (1) The initial system flow on which this analysis is based was reduced from 107.6% of the design flow rate to 106.5%.

(2) The densification power spike penalty was deleted from the analysip.

(3) Suitable sargin was included for an 11.2% DNBR reduction due to rod bow.

The specific credits used in this analysis to account for rod bow are as fo11ews:

DN3R credit Credit for rod bow penalty already

..$

• g included in analysis Credit for flow area reduction 1.0

=

factor in analysis Credit for 2* excess RC ficw 4*4 (3.5 available)

Total 11.2 o.s

.,. w:...

T.i b l e ai-1.

Cycle 2 and 3 Maximus Desien Conditions Cvele 2 Cycle 1 2esign power level %t 25o8 2508 System pressure, psia 2000 2200 Reactor coolant flev. 2 desica 107.6 106.5 Vessel inlet / outlet coolant temp 555.9/602.2 555.6/602.4 at 120*. power. F Ref design radial-local power pe.sking factor 1.78 1.73 Ref design.ixial riux shape 1.5 cosine 1.5 cosine Het channel f.ietors: Enthalpy rise 1.011 1.011 F. cat flux 1.014 1.014 Flev area 0.98 0.98 Active fuel length, in.

140.2 140.2 2I*

Avg heat flux at 100; pwer. Bru/h-ft 175,640 175.427 Max reat flux at 100% power. Stu/h-ft2 f.68.959 468.391 OtF correlation RAW-2 BAW-2 Min DN3R (~ power)(#'

1.86 (112) 1.90 (112)

(a)Cvele 2 heat flux was based on batch 3 densified length. Cycle 3 uses batch I. and 5 densified len;;th (located in hottest core location).

(b) Based on average heat flux with reference peaking.

(c) Cycle 2 DNBR includeo ef fects of densification power spike; cycle 3 does not.

i I

o.s.

.a.. w:..

I.

7.

ACCIDENT AND TRANSIENT ANALYSIS

!.1.

Generai Saferv Analysis Each FSAR1 accident analysis has been examined with r2spect to changes in cycle 3 parameters to determine the ef f ects of the cycle 3 reload and to ensure that thermal perf ormance is not degraded during hypothetical transients.

The core thermal parameters used in the PSAR accident analysis were design op-erating values based on calculated values plus uncertainties. Cycle 1 values (FSAR values) of core thermal parameters are compared with those used in the cycle 3 analysis in Table 6-1.

These parameters are common to all of the acci-dent analyses presented herein. For each accident of the FSAR a discussion and the key parameters are provided. A comparison of the key parameters (see Tat te 7-1) fros the FSAR and the present cycle 3 is provided with the accident dis-cussion to show that.he initial conditions of the transient are bounded by t he FSAR analysis.

The ef fects cf fuel densification on the FSAR accident results have been eval-uated and are repor*-d in BAV-1399.6 Since cycle 3 reload fuel assemblies con-tain f uel rods whc

neoretical density is higher than those considered in ref erence 6. the conclualons derived in that reference are still valid.

Calculation 41 techniques and cethods for cycle 3 analyses remain consistent with those used for the FSAK. Additional DNBR margin is shown for cycle 3 be-cause the S&V-2 CRF correlation was used instead of the V-3.

I

(

No new dose calculations were performed for the reload report. The dose con-siderations in the FSAR were based on maximum peaking and burnup for all core cycles; therefore, the dose considerations are independent of the reload batch.

'.2.

Rod Withdrawal Accidents This accident is defined as an uncontrolled reactivity addition to the core cu2 to withdrawal of control rods during startup conditions or f rom rated power Rabenck & \\Milens

condittens. !ct h types o: incident s were analyced in tne FSAR. - The inportant parameters during a rod withdrawal accidant are Doppler coef fietent.

>>f era t or te=perature coerficient. and the rate at which reactivity is added to the core.

Only high-pressure and high-flux trips are accownted for in the FSAR analysis, which ignores cultiple alar s.

interlocks, and trips that normally preclude this ty;e of incident. For positive reactivity additions indicative of these events. the most severe results occur ior 50L eenditions. The FSAR values of the key parameters for BOL conditions were -1.17

• 10' (ak/k/*F) for the Dop-pler coefficient. 0.5 = 10" ak/k f or the moderator tenperature coef ficient and red creup worths up to and including a ICT ak/k rod bank wort.5 Comparable cycle 3 paranetric values are -1,43 10 (;k!k/*F) for the Doppler coet t i-cient. -0.33 10 (ak/k/*F) for the moderater temperature coef ficient, and a saximum rod bank wort h of 8. 74 2k/k. Therefore, cycle 3 parameters are bo mded by design values assumed for the FSAR analysis. Thus, f or the rod wit hdrawal transients, the consequences will be no more severe than those presented in the FSAR. For the rod withdrawal t ros rated power the transient consequences are al+o less severe than those presented in the densification report.6

7. 3.

%3derater Dilution Accident Boren en the form of boric acid is utilized to control excess reactivity.

The l

boren centent of the reactor coolant is periodically reduced to compensate for f uel burnup and transient xenon ef fects with dilution water supplied by the makeup and purification systes. The moderator dilutica transients considered l

are the pumping of water with zero boron concentration f rom the makeup tank to the RCS under conditions of full-power operation, hot shutdown, and refueling.

The key parameters in this analysis are the initial boron concentration, boron reactivity worth, and moderator temperature coef ficient for power cases.

For positive reactivity additions of this type, the most severe results occur for 50L conditions. The FSAR values of the key parameters for 50L conditions were 1400 pro for the initial boron concentration. 75 pps/12 (ak/k) boron re-activity worth and +0.94 = 10 ak/k/*F for the moderator temperature coef fi-cient.

l Comparasle cycle 3 values are1000 pro for the initial boron cencentration. 60 ppe/1 (li/k) boron reactivity worth and -0.53 = 10 ' (ak/k)/*F f or the modera-l tor temperature coef ficient. The FSAR shows that the core and RCS 4re adequate-l ly protectea during this event.

Sufficient time for operator action to o.s

..s.. ser.s..

l,

terminate this transient is also shewn in the FSAR, even with =aximum dilution and mirocas shutdown margin. The predicted cycle 3 parametric values of in-portance ta the moderator dilution transient are bounded by the FSAR design values; t hus, the analysis in the FSAR is valid.

7.4 Cold Water (Pump Startup) Accident There are no check or isolation valves in the reactor coolant piping; there-fcre, the classic cold water accident is not possible. However, when the re-actor is operated with one or more pu=ps not running, and then these are turned on, the increased flow rate will cruse the average core te=perature to dec-ease.

If the moderator temperature coef ficient is negative, then reactivity will be added to the core and a power rise will occur.

Protective interlocka and procedures prevent starting idle pumps if the reac-tor power is above 223. However, these restrictions were ignored, and two-pump startup f rom 50T power was analyzed as the most severe transient.

To maximize reactivity addition, the FSAR analysis assumed the sont negative

~

soderator tesi.erature coefficient of -3.0 = 10 ' (ik/k)/*F and the least nega-

-5 tive Doppler aefficient of -1.30 = 10 ak/k. The corresponding most negative sederator te-*erature coef ficient and least negative Doppler Coefficient pre-

~5 Jicted for cycle 2 are -2. 55 = 10 ' and -1.56 = 10 (ak/k)/*F. respectively.

Since the predicted cycle 2 moderator temperature coefficient is less negative and the Deppler coef ficient is more negative than the values used in the FSAR.-

t he transient results would be less severe than those reported in the FSA1.

7.5.

Loss of Coolant Flow The reactor ecolant flow rate decreases if one or more of the reactor coolant pumps fail.

. pumping failure can be caused by mechanical f ailure or loss of electrical powet. With four independent pumps available, a sechanical failure in ene pump vill not affect the operation of others. With the reactor at power, the ef fect of loss of coolant flow 14 a rapid increase in coolant temperature dus to the reduction of heat removal capability. This increase could result in DNS if corrective action were not taken immediately. The key parameters for four-pu=p coastdown or a locked-rotor incident are the flow rate, flow coastdown characteristics, Doppler coef ficient. =oderator te=perature coeffi-cient, and hot channel DSB peaking f actors. The most conservative initial conditions were assumed for the denstfication reporti: FSAR values of flow l

10 ' tak/k)/*F Doppler eceificient. +0.5 10 (ak/h)/

~

and coastdown. -1.17

• F moderator temperature coefficient, with densified f uel power spike and peak-ing.

The results shewed that the DSSR remaired above 1.3 C.*-3) for the four-ptmp coastJcvn, and the fuel eladding temperature renained below criteria limits for tne lecked-rotor transiet.t.

The predicted parametric values for cycle 3 are -&.

3 10' (ak/k)/*F Doppler coefficient. -0.53 10 (ak/k)/* F moderator t emper iture coef ficient. and peaking f actors as shown in Table 6-1.

Since the predicted cycle 3 values are beunded by those used in the densification report. the results of that a nalysis represent the most severe consequences frra a loss-of-flow incident.

7.6.

Stuck-out, Stuck-In, or Dropped Control Rod

!1 a control rod were dropped into the core while it was operating, a rapid decrease in neutron power would occur, accompanied by a decrease in the core average coolant temperature. The power distributien might be distorted due to a new control red pattern, under which conditions a. turn to full power might lead to localized power densities and heat fluxes in scess of design limita-tions.

The key parameters for this transient are moderator

=perature coefficient.

dropped rod worth, and local peaking factors. The FdaR analysis was based on 0.46 and 0.36* ak/k rod worths with a moderator temperature coefficient of

-3.0 a 10 (ak/k)/*F. For cycle 3. the maximum worth rod at power is 0.20*

ak/k and a moderator temperature coetficient of -2.55 10 (ak/k)/*F. Since the predicted rod worth is less positive and the moderatcr temperature coef fi-cient is more positive. the consequencer of this transient are less severe than the results presented in the FSAR.

7. ?

Loss of Electric Power Two types of power losses were considered in the FSAi-(1) a loss-of-load condition caused by seoaration of tha unit from the transmission system and (2) a hypothetical condition resulting in a complete loes of all systen and unit power except that from the unit batceries.

Tne FSAR analysis evaluated the loss of load with and without turbine runback.

When there is no runback, a reactor trip occurs on high reactor coolant pres-sure or temperature. This case results in a non-limiting accident. The largest offsite dose occurs for the second case, i.e., loss of all electrical a..........

power except unit batteries. Jr.d assumana operation witn talled f ael.and steam 4enerator tube leakage. These results are independent ot core loading; there-fore, the results of the FSAA are applicable for.any reload.

7. M. St eas 1.ine Failure A steam line failure is de,f ined as a rupture of any of the steas lines from the steam generators. L*ren initiation of the rupture, both steam generators start to blow Jown, causing a sudden decrease in the primary system tempera-ture, pressure, and pressurizer level. The temperature reduction leads to positive reactivity insertion. and the reactor trips en high flux or low RC pressure. The ISAR has identitled a double-ended rupture of the steam line between the steam generator ar.d s* tan stop valve as the worst-case situation at end-of-tite conditions.

The key parameter for the core response is the moderator temperature coef fi-

[

cient, which was assumed in the FSAR to be -3.0 10 ~ (ak/k)/*F. The cycle

~

l 3 predicted value of moderator temperature coef ficient is -2.55 10 (ak/h)/

• F.

This value is bounded by those used in the FSAR analysis; hence. the re-suits in the FSAR represent the worst situation.

7. 9.

Stean Generatar Tube Failure A rupture or leak in a steam generator tube allows reactor coolant and associ-ated activity to pass to the second,ary system. The F3AR analysis is based on complete severance of a steam generator tube.

The primary concern for this incident is the potential radiological release, which is independent of core i

loading. Hence, the FSAR results are applicable to this reload.

l 7.10.

Fuel Handling Accident Th2 mechanical damage accident is considered the maximum potential source of activity release during fuel handling activities. The primary concern is radiological releases that are independent of core loadtag; therefore, the FSAR reeults are applicable to all reloads.

7.11.

Rod Efection Accident For reactivity to be added to the core more rapidly than by uncontrolled red withdrawal, physical f ailure of a pressure barrier component in the control rod drive assembly must occur. Such a failure could cause a pressure dif feren-tist to act on a control rod assembly and rapidly eject the assembly from the n.6..........

w core region.

Ihis incide t represent s t he most rapid reactivity insertion that CJn be reJMonab1:. postulated. The values used in t bc F SAR and Jensi:1eation re; ort at BOL conditions. -1.17 10" (ik/k)/*F Doppler ec+:ficient. +0.5 a 10 ' (.*.k/k)/*F sederator tecperature coefficient, and an efected rod worth of

~

J.s5% *k/k represent the maxtrum possible transient. The corresponding cycic 3 parametric values of -1.43 a 10-t;k/k)/*F Doppler. -d 53 - 10 ' t;k/kJ/*F

~

soderator temperature cocificient (both more negative than t50sc used La refer-ence 5), and a =axt=um predicted ejected rod worth of 0.**

A/k ensure that I and the den-the results will be less severe than those presented in the FSAR sification reporth.

.12.

Maxicum Hvpotherical Accident There is no postulated =cchaniam whereby this acetdent can occur since it would require a multitud* of failures in the engineered saf eguards. The hypothetical

(

accident is based solely on a gross release of radioactivity to the reactor building. The consequences af this accident are independ en t of core loading; nence, the results reported in the FSAR are applicable for all reloads.

7.11.

-ante Cas Tank Rupture i

The vaste gas tank was assumed to contain the gaseous activity evolved from degassing all of the reactor coolant following operatien with 12 defective f uel.

Rupture of the tank would result in the release of its radioactive contents to the plant eentilation systen and to the atmosphere through the unit vent.

The consequences of this incident are independent of core loading; therefore, l

the results reported in the FSAR are applicable to any relsad.

l 7.14.

LOCA Analvsis l

A generic LOCA analysis has been perfor=ed for the SFJ 177-FA. lowered-loop NSS using the Final Acceptance Criteria ECC5 evaluation model.I' The analysis is generic since the limiting values of key parameters for all plants in this category were used.

Furthermore, the combination of average fuel temperature as a function of linear heat rate and the lifetime pin pressure data used in the reterence 14 LOCA lisits analysis are conservative compared to those calcu-lated ror this reload. Thus, the analysis and the LOCA limits reported in ref-la provide conservative results for the operation of oconee 3. cycle 3 l

erence i

fuel.

The following tabulation shows the boundi:.6 values for allowable LOCA l

peak LHRs for oconee 3. cycle 3 fuel.

Babcock s Wilcox

,.c

l A;iovable peak linear Core elev.ttion, ft heat rite, Fa*/ : t j

15.5 4

16.o 6

18.0 9

17.0 10 16.0 Table 7-1.

Comparison of Key Para =eters for Accident Analysis FSAR. densi.* 4ed Pr ed ic ted Pa r sme t er value cycle 3 value 50L Doppler coeff. 10" (a k/ k) /

• F

-1.17

-1.43

~

EOL Doppler coeff. 10 (ak/k)/*F

-1.33

.56 Sol moderator coeff. 10 ' (ak/k)/*F

+0.5(b)

-0.

i EOL soderator coef f.10 (ak/k)/*F

-3.0

-2.

411 rod bank worth (H2P). I ak/k 10.0 3

Init ial boron cone (HFP). ppm 1400 10' Soron reactivity worth (7CF). ppe/1:

75

'd ak/k F.ax e j ec t ed rad wor t h (H FP).

• ak /k O.65 0.44 Lrepped rod worth (HTP).

ak/k O.46 0.20 I'#(-1.2 10'[' Ik/k/F) was used for stesa ILne failure at.alysis; t - 1. 3 10 ak/k/F) was used f ar cold water analysis.

D'(+0.94 = 10 ik/k/F) was used for the moderator dilution accident.

i e

O G

BLANK FRAME FOR PROPER PAGINATI-ON l

1 l

l

l l

l 1

8.

PROPOSED EIFICATIO*.S TO TECGICAI. SPECIFICATICNS The Technic.31 3pecif icatacos have been revised for cycle 3 operation. Changes were the results of the fellowing:

1.

Specifying APSR postti:o limits in addition to the usual regulating con-trol rod and isbalance limits f or ECCS. The APSR positiwi limits will provide additional control of pcwer peaking and assurance that 1.DCA kW/ft limits are not exceeded.

2.

I' sing 106.5% of desi4s flow rather than 107.6% as discussed in section 6.1.

3.

The FI.AME computer cede used in setting the Tec. nical Specification lis-i t s. 3 5

• I'*

6.

Tlie Technical Specif teation limits based on DN31 and L4R criteria include a; propriate allowanees for projected fuel rod bow penalties, i.e..

poten-tial redactico in DN31 and increase in power peaks. A statistical combt-teation of the nuclear.::2 certainty factor, engineering hot enar.ne l f actor, and rod bow peaking pe_.alty was used in evaluating 1.11R criteria, as ap-proved in reference 17.

l S.

Per reference 13. the power spike penalty due to fuel densification was not used in sett;: g the DNBR-and ECCS-dependent Technical Specification limits.

Sased on the Technical Spe:ifications derived f rom the analyses presented in enis report, the, Final A.:ceptance Criteria ECCS limits will not be exceeded, j

nor will the thersal destgm criteria be violated. Figures 8-1 through 8-1;.

illustrate revisions to previous Technical Specification Itzits: Figures 3-15 througn 8-17 illustrate ltries not previously included in the Technical Spe-eificaeions.

l l

m ow :.*r.r w-s* - ~

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. 129

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Operational Power Imbalance Envele x for Operation Froo O to 100 : 10 ETP3 -

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Poser. 5 of 2568 NWt RESTRICTED REGION

-13.73,102

^ 8.66.102

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Tigure 3-13.

Operatienal Power Imb tace Envelope for Operatien From 100 : IJ to 2 35 : 10 EFPD

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RESTRICTED REGION

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l f-9.

STARTUP PROGRAM - PHYSICS TESTING The planned startup testing associated with core performance is outlined be-low.

These tests verify that core performance is within the assumptions of the safety analysis and provide the necessary data for centinued safe plant operation.

Pre-Critical Tests.

1.

Control rod drop tLae Zero Power Tests 1.

Critical boron concentration 2.

Temperatura reactivity coef ficient 3.

Control rod group worth 4

Ejected rod worth Pewer Tests 1.

Core power distribution verification at approximately 40, 75, and 100: FP, normal control rod group configuration.

2.

Incore/out-of-core detector imbalance correlatica verification' at approxi-nately 75: FP.

3.

Power Doppler reactivity coefficient at approximately 100: FP.

4.

Temperature reactivity coefficient at approximately 100: FP.

e

• e g

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I

REFERENCES I Oconee Nuclear Station, Units 1, 2 and 3. Final Safety Analysis Report, Docket Nos. 50-269, 50-270, and 50-287.

2 Oconee Unit 3, Cycle 2 Reload Report, BAW-1432, Babcock & Wilcox, June 1976.

3 Fuel Assembly Stress and Deflection Analysis f or Loss-of-Coolant Accident and Seismic Excitation, BAW-10035, Babcock & Wilcox, June 1970.

" Program to Determine In-Reactor Performance of B&W Fuels -- Cladding Creep Collapse. BAW-10084, Rev. 1 Babcock & Wilcox, Nove=ber 1976.

5 C. D. Morgan and H. S. Kao, TAFY - Fuel Pin Temperature and Gas Pressure Analysis, BAW-10044 Sabcock & Wilcox, May 1972.

6.0conee 3 Fuel Densification Report. BAW-1399, Sabcock & Wilcox, November 1973.

7 B. J. Buescher and J. W. Pegram, Babcock & Wilcox Model for Predicting In-Reactor Densification, BAW-10083P. Rev. 1. Eabcock & Wilecx, November 1976.

4 Oconee Unit 1. Cycle 4 Reload Report, BAW-1447, Babcock & Wilcox, March 1977.

• Oconee Unit 2, Cycle 3 Reload Report, BAW-1452, Babcock & Wilcox, April 1977.

l 13 Correlation of Critical Heat Flux in a Bundla Cooled by Pressurized Water, J

BAW-10000A, Babcock & Wilcox, June 1976.

• 1 K. W. Hill, et al., " Effects on Critical Heat Flux of Local Heat Flux Spike or 1.ocal Fluw Blockage in PWR Rod Bundles " 74-WA/HT-54 ASME Winter Annual Meeting, New York, November ~ 1974.

12 CHF -- Critical Heat Flux Correlation for CE FA With Standard Spacer Crid --

Part 2. " Nonuniform Axial Power Distribution," CENPD-207. Combustion En-

)

i gineering, June 1976.

w BLANK FRAME FOR PROPER PAGINATION

3 9

13 " Core Physics Methods Data Used as Input to IDCA Analysis " XN-75-l.2, August 1975 and letter, D. A. Sixel, Constssers Power, to R. A. Purple. April 5.

1976.

I '* ECCS Analysis of B&W's 177-Fuel Ass,sably, Lowered-Loop NSS. BAW-10103, Bab-cock & Wilcox, Lynchburg, Virginia, June 1975.

15 FLAME - Three-Dimensional Noded Code for Calculating Reactivity and Pow.

Distributions, BAW-1012'A, Babcock & Wilcox, Lynchburg, Virginia August

1976, 16 C. W. Mays, Verification of Three-Disensional F1.AME Code, BAW-10125. Bab-cock & Wilcox, Lynchburg, Virginia, August 1976.

17 S. A. Varga (NRC) to J. H. Taylor (B&W), Letter, "Connents on 5&W's Submit-tal on Combination of Peaking Factors," May 13, 1977.

I* K. E. Suhrke (B&W) to S. A. Varga (NRC), Letter "Densification Power Spike," December 6, 1976.

s.

l l

Babcocks Wilecx s_,

SUPPLD:r*,7 TO bah *-14 4 7 OCONEE I, CYCLE 4 RELOAD REPORT SEPTEMSER, 1977 l

1 1

e I

i 1

l

Int r. % tJ. n and ju.arv hi? repert suppl ants the R enee 1. C=cle a heloa.! h; or t ( P.NJ-14 4 7. .o r s h I' M i to account for a c.oditisd Cycle 4 core loadine. The 2.stif led core loading censist s of t he leading of four onec burned l'.atch 2 fuel asse-blics

.a t core locations D-4 D-12. N-4 and N-12. origin.slly intended to be loaded with f our Eatch 4 f uel assenblics. The replacernent of the four Batch 4 fuct assemblics with the four Batch 2 a.ssemblies was necessitated because of =cchanical da. ace to one Batch 4 asse-bir while it was examined in the spent fuct pool.

The codified core loadinc does not affcct the results of t he cor e saf ety analys ts or limit inc condit ions of operat ien, as documented in F.N4-1447.

The f ollowine paragraphs describe the four replaceeent fuel assemblies and provide evaluat ions of t he impact of the modified core loadins,upon ti e prevsoua analyses.

Nuricar Desirn The codifled core loading consists of replacinc. 4 twice burned. 3.2 vt I tatch a fuel asse-blics with 4 ence burned. 2.1 wt batch 2 fuel assemblies.

The batch 2 assemblics were selected to natch the reactivity of the replaced batch 4 assee5lics as closelv as possible.

Ik'C radial power dist ribution results f or t he revised core loading show that all fuel assembly pevers are within 1*. of the original Ovele 4 core leading. except for the replacenent locat ten W12. which has about 3.5: less pner in ti.e revised core loading.

A Cvcle 4 l'DQ deplet ion analysis was perf or cd which showed that these differences beco e proeressively smaller as the evele is depleted. Ficure I shows the revised core loading diagram f or Oconce 1 Cycle 4 and Figure 2 is an eigt.th-core cap showing the burnup and enrichment distribution at the 1

beginning of Cyclt-4 with the revised core loading.

Mechanical Design The replace =ent fuel assemblies have been evaluated for mechanical design adequacy and found to meet the criteria for allowable cladding strain and irradiation swelling specified in the Reload Report. The creep collapse tice was determined to be > 30.000 EFPH. which as greater than the n.sx!=um design life of 14.232 EFPH.

Pertinent fuel design parameters for the a Batch 2 fuel assemblics are given in Table 1.

Therra1-Hydraulic Design The revised core loading has been evaluated f or thermal-hydraulic design considerations and found not to affect the design presented in the Reload Report. The Mark B-2 replacement assemblies have a higher resistance to flow than the Mark B-3 assemblics being replaced. However, the hot assembly during Cycle 4 is either a Mark B-3 or B-4 assembly that is not in a core location where the replace:ent asse=blies will be placed. Since the inser-tion of the higher resistance Mark B-2 asseeblies will cause an increase in flow in the hot assembly, the thermal-hydraulir design presented in Section 6 of the Reload Report is conservative.

' e f.- t y An i l,v s i..

The systet par. meters it:por t an t to safety analv~ta, are not a f f e r t. ! t v t h.-

sevised core Inad in.:. and the safety evaluatson prtsented in t h+

F elo. d Report therefore renains valid.

7??h"_I 3I SP?f1I_lf 3.t_1 ons F

The power peaking limits and the ejected and shutdown rod insertion licits have been verifled and found to be lew Il=1 tint than the 11 it s calculated previcusly.

h.esed on these results, t he Technical pec if icat ions previous 1v suleitted are valid for the m.wlified loading and ilo not require revision.

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END i

M CR0 PHOTOGRAPHER _#-,______

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MICROFILM SECTION 1

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8UIL(MPdG ISP 7 W ASHINGTG4 NAVY YARD l

WA$HsNGTON D C. 20374 1

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