ML19351A419
| ML19351A419 | |
| Person / Time | |
|---|---|
| Site: | Mcguire, Catawba, McGuire, 05000000 |
| Issue date: | 09/30/1989 |
| From: | FRAMATOME COGEMA FUELS (FORMERLY B&W FUEL CO.) |
| To: | |
| Shared Package | |
| ML19351A418 | List: |
| References | |
| BAW-10174, NUDOCS 8910230185 | |
| Download: ML19351A419 (151) | |
Text
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<,A RAW-10174 l
Topical Report September 1989 i
j Mark-BW Reload LOCA Analysis l
e for the Catawba and McGuire Units i
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I Babcock & Wilcox Fuel Comp &tay P.
O. Box 10935 2
Lynchburg, Virginia 24506-0935
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G910230185 G70117 PDR ADOCK 050s0369 F'
FDC
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Babcock & Wilcox Fuel Company P. O. Box 10935 F
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Lynchburg, Virginia 24506-0935 l
Topical Report BAW-10174 l
september 1989 1
Mark-BW Reload LOCA Analysis l
t for the j
Catawba and McGuire Units I
i Kev Wordar Largp Break. IDCA. Transient. Water Reactors l
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i The B&M Fuel Company will be delivering reload fuel to the Duke Power Catawba and McGuire Units beginning in 1991.
This report pre sents a complete LOCA evaluation.for operation of the Catawba and McGuire nuclear units with Mark-BW reload fuel.
Compliance l
with the criteria of 10 CFR 50.46 is demonstrated.
Operation of
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the units while in transition from Westinghouse-supplied OFA fuel to B&W-supplied Mark-BW fuel is also justified.
Other B&W
[
topical reports describe the Mark-BW fuel assembly design; the mechanical,
- nuclear, and thermal-hydraulics methods supporting j
the design; and ECCS codes and methods.
The analyses and eval sations prssented in this report serve, in conjunction with the other topical reports, as a reference for future reload safety evaluations applicable to cores with BWFC-supplied fuel assemblics.
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ACKNOWLEDGENELS i
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1 The B&W Fuel Company wishes to acknowledge the efforts put forth
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by J.
R.
- Biller, J. J.
- Cudlin, B. M.
- Dunn, J.
A.
Flingenfus, R.
l J.
- Zowe, C.
K.
Nithianandan, H.
H.
- Shah, and K.
C.
Shieh in preparing and documenting the material contained in this report.
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Tonical Renert Revision Record l
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Documentation l
Revision Description j
0 Original Issue i
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TABLE OF CONTENTS f
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1.
Introduction 1.1 2.
Summary and Conclusions.
2.1 l
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3.
Plant Description 3.1 f
3.1 Physical Description.
3.1 3.2 Description of Emergency Core Cooling System.
3.4 3.3 Plant Parameters.
3.5 l
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Analysis Inputs and Assumptions.
4.1 4.1 Computer Codes and Methodt 4.1 l
4.2 Inputs and Assumptions.
4.1 l
4.2.1 RELAP5/ MOD 2-B&W Modeling 4.2 l
4.2.2 REFLOD3B Modeling.
4.7 l
4.2.3 FRAP-T6-B&W Modeling 4.9
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4.2.4 BEACH Modeling 4.9 4.3 Comparison of Plant Model with McGuire & Catawba 4.10 I
5.
Evaluation Model Changes 5.1
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5.1 Revisions to the BWFC LOCA Evaluation Model.
5.2
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5.2 Effects of Evaluation Model Revisions 5.4 i
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6.
Sensitivity Studies.
6.1 6.1 Evaluation Model Generic Studies.
6.1 6.2 Confirmable Sensitivity Studies 6.5 l
6.3 Break Location.
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7.
Plant-Sp6Jific Studies and Spectrum Analysis 7.1 7.1 Base Case 7.1 f
7.2 Accumulator Configuration 7.2 I
7.3 Break Spectrum Analysis 7.3 j
7.4 Break Type.
7.5 7.5 Maximum ECCS Analysis 7.7 f
8.
LOCA Limits.
8.1
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.S.1 LOCA Limits Dependencies.
8.1 8.2 LOCA Limits Calculation Results 8.2 l
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8.3 Compliance to 10 CFR 50.40 8.6 h
9.
Whole-Core Oxidation and Hydrogen Generation.
9.1 e
10.1 10.
Core Geometry.
P 11.
Long-Term Cooling.
11.1 11.1 Initial Cladding Cooldown 11.1 11.2 Extended Coolant Supply 11.2 11.3 Boric Acid Concentration.
11.2 11.4 Adherence to Long-Term Cooling Criteria 11.4 1-12.
Small Break LOCA 12.1 l
12.1 12.2 Fuel Design Effects 12.4 l
12.3 Current FSAR Results.
12,7 12.4 Compliance with Acceptance Criteria 12.7 13.
References 13.1 l
Appendix A.
Evaluation of Transition Cores A.1 A.1 OFA and Mark-BW Design Differences A.1 A.2 Assessment of Impact on Peak Cladding Temperatures A.3 A.3 Conclusions.
A.6
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l fs List of Tahles k_
i Table Page 2-1 Summary of Retalts (LOCA Limit Runs) 2.3 3-1 Plant Parameters and Operating Conditions.
3.6 j
4-1 Comparison ot LOCA Model Geometric Values j
to Plant FSAR Data.
4.14 l
1 7-1 sisctrum and Break Type Comparison 7.9 8-1 LOCA Limits Results.
8.7 A-1 OFA / Mark-BW Design Differenceu A.7 l
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I List of Fiaures Figure Page 4i Large Break Analysis Code Interface 4.16 4-2a RELAP5/ MOD 2-B&W LBLOCA Noding Diagram Reactor Coolant Loops 4.17 i
4-2b RELAP5/ MOD 2-B&W LBLOCA Noding Diagram
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Reactor Core.
4.18 4-3 REFLOD3B Noding Diagram 4.19 4-4 BEACH and FRAP-T6-B&W Noding Diagram j
f for Mr %-BW Fuel Assembly 4.20 5-1 Revision 0 Evaluation Model Code Interfaces i
for Large Break LOCA.
5.9
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6-1 Core Mass Flux for EM Spectrum Study Versus McGuire/ Catawba Base Model 6.13 6-2 TACO 3 Fuel Temperatures as a Function of Burnup 6.13 7-1 Plant-specific Studies Analysis Diagram 7.10 7-2 Sensitivity Study - Base Model System Pressure During Blowdown.
7.11 7-3 Sensitivity Study - nase Model Mass Flux During Bicwdown at Peak Power Location.
7.11 7-4 Sensitivity Study - Base Model Reflooding Rate 7.12 7-5 Sensitivity Study - Base Model Heat Transfer Coefficient at Peak Power Location.
7.12 7-6 Sensitivity Study - Base Model Peak Cladding 7.13 Temperature 6
7-7 Sensitivity Study - Base Model Cladding Temperature at Rupture Location 7.13 7-8 Sear,itivity Study - Base Model Cladding Temperature in Adjacent Grid Span 7.14 7-9a St.isitivity Study - Base Model Fluid Temperature at PCT Location 7.14 7-9b Sensitivity Stud) - Base Model Fluid L
Temperature at PCT Location 7.15
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List of Floures (Con't)
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Figure Page j
7-10 Accumulator Study - McGuire Vs Catawba
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Accumulator Flow Rates.
7.15 7-11 Accumulator Study - McGuire Vs Catawba Peak Cladding Temperatures.
7.16 7-12 Discharge Coefficient Study - Cd = 1.0 System Pressure During Blowdown.
7.16 1
7-13 Discharge Coefficient Study - Cd = 1.0 j
Mass Flux During Blowdown at Peak Power Location.
7.17
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7-14 Discharge Coefficient Study - Cd = 1.0 Reflooding Rate 7.17 1
7-15 Dischargs coefficient Study - Cd = 1.0 Heat Transfer coefficient at Peak Power Location-7.18 7-16 Discharge Coefficient Study - Cd = 1.0 Peak Cladding Temperature.
7.18 7-17 Discharge coefficient Study - Cd = 1.0 Cladding
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Temperature at Rupture Location 7.19 7-18 Discharge Coefficient Study - Cd = 1.0 Cladding Temperature in Adjacent Grid Span 7.19 l
7-19a Discharge Coefficient Study - Cd = 1.0 Fluid Temperature at PCT Location 7.20 l
7-19b Discharge Coefficient Study - Cd = 1.0 Fluid Temperature at PCT Location 7.20 7-20 Discharge Coefficient Study - Cd n 0.8 System l
Pressurs During Blowdown.
7.21 7-21 Discharga coefficient Study - Cd = 0.8 l
Mase Flux During Blowdown at Peak Power Location 7.21 7-22 Discharge. coefficient Study - Cd = 0.8 Reflooding Rate 7.22 7-23 Discharge coefficient Study - Cd = 0.8 Heat Transfer Coefficient at Peak Power Location 7.22 ix -
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i List of Ficures (Con't) f Figure Page f
7-24 Dischkrge Coefficient Study - Cd = 0.8 Peak l
[/gy Cladding Temperature.
7.23 f
7-25 Dischat.'ge Coefficient Study - Cd = 0.8 Cladding Temperature at Rupture Location 7.23 f
7-26 Discharge Coefficient Study - Cd = 0.8 l
Cladding Temperature in Adjacent Grid Span.
7.24 i
7-27a Discharge. coefficient Study - Cd = 0.8 Fluid f
Temperature at PCT Location 7.24 7-27b Discharge Coefficient Study - Cd = 0.8 Fluid f
Temperature at PCT Location 7.25 7-23 Discharge Coefficient Study - Cd = 0.6 System Pressure During Blowdown.
7.25 7-29 Discharge Coefficient Study - Cd = 0.6 Mass Flux Duritag Blowdown I
at Peak Power Location.
7.26 7-30 Discharge Coefficient Study - Cd = 0.6 h,
Reflooding Rate 7.26 7-31 Discharge Cecfficient Study - Cd = 0.6 Heat Transfer Coefficient at Teak Pow 1r Location 7.27 7-32 Discharge Coefficient Study - Cd = 0.6 Peak f
Cladding Temperature.
7.27 j
7-33 Discharge Coefficient Study - Cd = 0.6 Cladding Temperature at Rupture Location.
7.28 7-34 Discharge Co'ef ficient Study - Cd = 0. 6 Cladding Temperature at Adjacent Grid Span.
7.28 7-35a Discharge Coefficient Study - Cd
- 0.6 Fluid Temperature at PCT Locati,on 7.29 7-35b Discharge Coefficien. Study - Cd = 0.6 Fluid Temperature at PCT Location 7.29 7-36 Braek Type Study - Split, Cd = 1.0 System Pressure During Blowdown.
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I List of Figures (Con't) f g~w
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Figure Page 7-37 Break Type Study - Split, Cd = 1.0 l
1 Mass Flux During Blowdown
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at Peak Power Location..
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7-38 Break Type Study - Split, Cd = 1.0 j
Reflooding Rate 7.31 j
7-39.
Break Type Study - Split, Cd = 1.0 Heat i
l Transfer coefficient at Peak Power Location 7.31 7-40 Break Type Study - Split, Cd = 1.0 Peak l
Cladding Temperature.
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7-41 Break Type Study - Split, Cd = 1.0 Cladding Temperature a: Rupture Location 7.33 l
7-42 Break Type Study - Split, Cd = 1.0 Cladding l
Temperature in Adjacent Grid Span 7.33 7-43a Break Type Study - Split, Cd = 1.0 Fluid l
l Temperature at PCT Location 7.33 7-43b Break Type Study - Split, Cd = 1.0 Fluid l
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Temperature at PCT Location 7.34 7-44 Maximum ECCS Study - DFOLB, Cd = 1.0 Minimum ECC Pi.mped Injection Containment Pressure.
7.34 7-45 Maxirxm ECCS Study - DECLB, Cd = 1.0 System Prassure During Blowdown.
7.35 7-46 Maximum ECCS Study - DECLB, Cd sa 1.0 Mass Flux During Blowdow' at Peak Power Location 7.35
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7-47 Maximum ECCS Study - DECLB, Cd = 1.0 Maximum ECC Pumped Injection Containment Pressure.
7.16 7-48 Maximum ECCS Study - DECLB, Cd = 1.0 Pumped ECC Injection F*ow Rate.
7.36 7-49 Maximum ECCS Study -
dCLB, Cd = 1.0 l
Dowl' ceder Water Lev.1 7.37 7-50 Maximum ECCS Study - DECLB, Cd = 1.0 Reflooding Rate 7.37
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List of Flaures (Con't\\
Figure Page f
7-51 Maximum ECCS Study - DECLB, Cd = 1.0 Heat Transfer coefficient at PCT Location 7.38 7-52 Maxistin ECCS Study - DECLB, Cd = 1.0
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Heat Transfer Coefficient at Rupture Location 7.?8 l
7-53 Maximum ECCS Study - DECLB, CC = 1.0 Heat Transfer coefficient in Adjacent Grid Span 7.39 f
7-54 Maximum ECCS Study - DECLB, Cd = 1.0 l
Peak Cladding Temperacure 7.39 7-5$
Max!aum ECCS Study - DECLB, Cd = 1. 0 i
Cladding Tempcrature at Rupturo Location.
7.40 7-56 Maximu:e ECCS Study - DECLB, Cd = 1.0 Cladding Temperature in Adjacent Grid.2 pan.
7.40 7-57a Maximum ECCS Study - DECLB, Cd = 1.0 Fluid Temperature at PCT Location 7.41 7-57b Maximum ECCS Study - DECLB, Cd = 1.0 Fluid Temperature at PCT Locdtlen 7.41 7-58a Maximum ECCS Study - DECLB, Cd = 1.0 Fluid Temperature at Rupture Location 7.42 7-58b Maximum ECCS Study - DECLB, Cd = 1.0 Fluid Temperature at Rupture Location 7.42 7-59a Maximum ECCS Study - DECLB, Cd = 1.0 Fluid l
Temperature in Adjacent Grid Span 7.43 7-59b Maximum ECCS Study - DECLB, Cd = 1.0 Fluid Temperature in Adjacent Grid Span 7.43 8-1 Axial Dependence of Alleved Total Peaking
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Factor Large Break LOCA Mark-BW 8.8 8-2 Normalized Local Power Burnup Dependency j
Factor.
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8-3 LOCA Limits Study - Axial Power Shapus 8.9 8-4 LOCA Limits Study - 2.9 Foot Case Mass Flux
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During Blowdown at Peak Power Location.
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4 List of Fluures (Con't) 73b Figure Page 8-5 LOCA Limits Study -~2.9 Foot Case cladding Temperatures 8.10 8-6 LOCA Limits Study - 2.9 Foot Case Heat Transfer Coefficient at PCT Locatien.
8.10 8-7 LOCA Limits Study - 2.9 Foot Case Local Oxidation 8.11 8-8 LOCA Limits Study - 4.6 Foot Case Ma5? Flux During Blowdown at Peak Power Locatign...
8.11 8-9 LOCA Limits Study - 4.6 Foot Case cladding Te7peratures 8.12 8-10 LOCA Limits Study - 4.6 Foot Case Heat Transfer Coefficient at PCT Location.
8.12 8-11 LOCA Limits Study - 4.6 Foot Case socal Oxidation 8.13 8-12 LOCA Limits Study - 6.3 Foot Case Mass Flux During Blowdown at Peak Power Location.
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S-13 LOCA Limits Study - 6.3 Foot Case Cladding Temperatures 8.14 8-14 LOCA Limits Study - 6.3 Foot Case Heat Transfer Coefficient at PCT Location.
8.14 8-15 LOCA Limits Study - 6.3 Foot case Local Oxidation 8.15 8-16 LOCA Limits Study - 8.0 Foot Cese Mass Flux During Blowdown at Peak Power Location.
8.15 8-17 LOCA Limits Study - 8.0 Foot Case Cladding Temperatures 8.16 8-18 LOCA Limits Study - 8.0 Foot Case Heat J
Transfer Coefficient at PCT Location.
8.16 8-19 LOCA Limits Study - 8.0 Foot Case Local Oxidation 8.17 8-20 LOCA Limits Study - 9.7 Foot Case Mass Flux During Blowdown at Peak Power Location.
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List o.? Fiaures (Con't)
Figuro Page 8-21 IhCA Limits Study - 9.7 Few case i
Cladd'. ag *fssperatures 8.18 t
8-22 LOCA Limits Study - 9.7 Foot Case Heat Transfer Coefficient at PCT Location.
8.18 i
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8-23 IDCA Li: nits Study - 9.7 Foot Cass Local Oxidation 8.19
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1.0 Introduction
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The B&W Fuel Company (BWFC) will be delivering reload fuel to the Duke Power Catawba and McGuire units beginning in 1991.
The Mark-l BW reload fuel will be similar in design and performance to l'uel assemblies already licensed and operating in these plants.
In accordance with the requirements of 10 CFR 50.46 and 10 CFR 50, Appendix K,
an evaluation of the Emergency Core Cooling System (ECCS) performance has been performed for BWFC reload fuel for the Catawba and McGuire units.
In presenting that analysis, this report complements other BWFC topical reports that describe the Mark-BW fuel design; the mechanical,
- nuclear, and thermal-hydraulics methods supporting the design; dnd ECCS codes and methods.
The analyses and ovaluations presented in this report are intended to serve, in conjunction with these other *)pical reports.
l as a reference for future reload safety evaluations of the Catawba l
and McGuire units applicable to cores with BWFC-supplied fuel I
assemblies.
The McGuire and Catawba nuclear power plants use nuclear steam i
supply systems designed by Westinghouse that are representative of l
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the standard Westinghouse four-loop, 3411 Mwt design, with the j
exception of certain /.nternal structures.
The ECCS provided for the plant consists of the conventional combination of high pressure pumped injection, pressurized water storage ttnks, and low pressure l
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pumped injection, all connected into the reactor coolant piping l
just upstream of the reactor vessel.
The McGuire and Catawba t
containments are of the ice condenser type, limiting building l
press res during a postulated loss of coolant accident (LOCA) to near atmospheric values.
The results of calculated predictions of LOCAs mus.t meet thc criteria imposed by 10 CFR 50.46.
At the time of initial i
operation, the McGuire and Catawba plants were fueled with Westinghouse-supplied fuel and compliance was demonstrated by
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I calculations performed by Westinghouse and Duke Power.
This report i
documents compliance to 10 CFR 50.46 when the plants are fueled by BWFC-supplied fuel.
In this report, possible LOCAs are divided j
into two groups depending on the assumed break size.
For breaks Ir.cger ther. 1. 0 ft', compliance is demonstrated by calculations and analyses performed in accordance with the BWTC large break LOCA evaluation model for recirculating steam generator plants, BAU-I 10168 (Reference 1).
For breaks smaller than 1.0 ft, compliance is shown by valf dsting that the calculations performed in support of the plant prior to the loading of the Mark-BW fuel remain l
applicable when the Mark-BW is in use.
f A summary of the results of the annlyses is presented in Chapter 2.
Chapter 3 provides a general description of ths Catawba and McGuire plants including their major differences.
The analysis parameters used for the large break calet.lations are discussed in Chapter 4, evaluation model changes incorporated for this analysis are discussed in Chapter 5, and system nors.tivity studies in Chapter 6.
The large break spectrum analysis to determine the most limiting break is documented in Chapter 7.
Chapter 8 presents the LOCA limit calculations which confirm adherence to the first two criteria of 10 CFR 50.46, The evaluation of marinum hydrogen f
generation, coolabl's geonstry, and long-term cooling are presented in Chapters 9,
10 and 11, respectively.
Validation of the applicability of the earlier small break LOCA studies is provided in Chapter 12.
During the transition from Westinghouse fuel to the Mark-BW g
assembly, the core will for some time consist vf a mix of the two fuel assembly types.
For sucn cyc3 es, Appendix A shows that the mixing of the assemblies does not alter the IOCA performance of either fuel assembly to any degree approaching the criteria of 10 CFR 50.46.
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2.0 su===rv and conclusion l
i 10 CFR 50.46 specifies that the emergency core cooling system f
(ECCS) for a
commercial nuclear power plant must meet five eW uations documented in this criteria.
The calculations and 1
report demonstrate that the Duke Pwet company McGuire and Catawba l
units continue to meet these criterin when operated with Mark-BW fuel.
Large break LOCA calculations performed in con:urrence with j
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approved evaluation model (BAW-10168 and revisions) demonstrate compliance for a full Mark-BW core for breaks up to and including l
the double-ended severance of the largest primary coolant pipe.
The small break 14CA calculations used to license the plant f
operation during previous fuel cycles are shown to be unaffected by j
the change in fuel design, and therefore, demonstrate that the plant meets 10 CFR 50.46 for small breaks when loaded with Mark-BW fuel.
The coexistence of the Mark-BW assembly and the Westinghouse OFA assembly in the same fuel cycle is shown to be inconsequential and does not cause the calculated temperatures for either assembly l
(N to approach the limits of 10 CFR 50.46.
Specifically, this reporv, demonstrates that when the McGuire and Catawba plants are operated with Mark-BW fuels 1.
The calculated peak cladding temperatures for the limiting i
cases are lesr than 2200 F, (Chapter 8).
2.
The ma :imum calculated local cladding oxidation is less than i
17.0 %,
(Chapter 8).
3.
The maximum amount of core-wide oxidation does not exceed L_Q
% of the fuel claddina, (Chapter 9).
4.
The cladding remains amenabl9 to coolina, (Chapter 10).
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.Lona-term coolina is estr.blished and maintained after the IDCA, (Chcpter 11).
i The results of the large break sensitivity studies and th? break i
spectrum studies performed with the BWFC evalcation model show that i
the double-ended guillotir.e break at the pump discharge with a discharge coefficient of 1.0 and maximum ECCS is the most limitiny f
case.
Table 2-1 shows the results of this accident on the Mark-BW l
fual design when the assumed axial location of peak power is varied along the length of the pin.
As the local power for the Mark-BW l,
assemblies is controlled such that it can not exceed the Mark-BW I4CA limits curve, this table lists the results of the calculations which demonstrato compliance with the first four critaria of 10 CFR 50.46.
Compliance with the long-term cooling criterion (as described in Chapter 11), is through the w e of a pumped t.njection system that can be recirculated, drawing water from the reactor building through a heat exchanger, to provide extended energy removal.
The concentration of boric acid is held below its solubil'ity limit by starting hot leg injection with 15 hours1.736111e-4 days <br />0.00417 hours <br />2.480159e-5 weeks <br />5.7075e-6 months <br /> of the accident.
i During the transition from the Wostinghouse OFA to the Mark-BW, both fuel assemblies will reside in the core simultaneously for several fral cycles.
Appendix A demonstrates that the results and conclusions presented above are also applicable to the Mark-BW ass mblies in the trancition core.
Similarly, Appendix A
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demonstrates that the insertion of the Mark-BW fuel with the OFA j
does not advarsely affect the cooling of the OFA.
- Thus, the original calculations that showed that the OFA met the criteria of 10 CFR 50.46 temain valid for the OFA assemblies through the j
transition period.
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g-Table 2-1 summary of I asults (LOCA Limit Runs)
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. Core Peak Cladding Maximum oxidetion,%
l Elevation, ft Tamparature F
Lts,31 Whole Core j
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2.9 1816 3.4 0.25 l
ii d.6 1963 5.2 0.41 l
i 6.3 1873 4.8 0.40 j
l 8.0 1930 4.7 0.32 h
9.7 1823 3.7 0.29 l
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3.0 Plant Descrintion The McGuire and Catawba nuclear power plants use nuclear steam supply systems (NSS) designed by Westinghouse that are j
representative of the standard Westinghouse four-loop, 3411 MWt j
design except for certain internal structures provided for the j
upper head injection system.
The ECCS provided for the plant
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consists of the conventional combination of high pressure pumped i
injection, pressurized water storage tanks, and low pressure pumped I
injection all connected into the reactor coolant piping just upstream of the reactor vessel.
Both the McGuire and Catawba plants use ice condenser containment systems.
L,1 Physical Descrintion The reactor coolant system is enclosed entirely within a
containment and in arranged into four heat transport loops, each of which has one recirculating steam generator and one reactor coolant pump.
The reactor coolant is directed through the nuclear core within the reactor vessel, transported to the steam generators via j
four pipes (hot legs), cooled within the stcan generator tubes, and returned to the reactor vessel through f1ur cold leg pipes.
Flow
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through the system is driven by four reactor coolant pumps, one I
per coolant loop.
System pressure is maintained by a pressurizer
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l connected to one of the coolant loops in the hot leg.
l Reactor Vessel t
The reactor vessel is basically a
cylindrical shell with a
hemispherical bottom head and a removable hemispherical upper head.
Major regi.as of the reactor vessel are the inlet and outlet I
- nozzles, the downcomer, the lower plenum, the core, the upper plenum, and the upper head.
Coolant enters the vessel through one of four inlet nozzles and passes downward through the dcwncomer to the lower plenum.
From the lower plenum, coolant ia directed
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t upward passing either through the core or the baffle bypass region to the upper p1enum.
Within the upper plenum, the coolant mixes f
with a small amount of flow which was bypassed directly fam the f
'downconer to the upper head a,nd exits the reactor vessel through l
the hot leg nozzles, j
t In the original design, the McGuire units baffle bypass region was not connected in paral 3e1 with the core, but instead, connected to the top and bottom of the downcomer.
During the 1980s, this design was shown to lead to damage of the core perimeter fuel assemblies l
by small, horizontal jets of fluid pasuing around the edges of the plates which form the baffle regior..
Altering the baffle region to connect between the inlet plenum and the outlet plenum substantially reduces the pressure drop between the core and the baffle region, and eliminates the fluid jets.
In 1987, Duke Power l
contracted for the alteration of the McGuire units so that the baffle bypass region would be connected between the lower plenum and the outlet plenum.
As this work is planned prior to the i
l introduction of the Mark-BW fuel assembly into either of the l
McGuire units, this report considers both McGuire and Catawba l
plants to be of the upflow type within the baffle bypass region.
Reactor Core and Fuel Assembly The reactor core is comprised of 193 fuel assemblies, with each fuel assembly consisting of 264 fuel rods, 24 guide thimbles, and one instrument tube.
Each fuel rod consists of stacked fuel l
pellets contained in a Zr-4 fuel rod with a gap between the fuel pellet and the fuel rod.
Fifty-three of the fuel assemblies have rod cluei.er control assemblies (RCCAs) used for power control and shutdown capability.
McGuire Unit I has silver-indium-cadmium control rods, while the other units have BC control rods.
The 4
Catawba and McGuire plants will be replacing the Westinghouse OFA fuel assembly with the Mark-BW fuel assembly.
Both the Mark-BW and the OFA are 17 x 17 fuel rod arrays with an active length of
- 3.2 -
g approximately 12 feet.
A comparison of fuel rod geometries for I
both fuel types is provided in Appendix A.
Reactor Coolant imens The coolant loop piping is connected to the reactor vessel through eight nozzles, all of which are located at the same elevation, approximately six feet above the top of the core.
The outlet piping (hot legs) runs from the reactor vessel in a hocizontal plane and undergoes an upward bend as it attaches to the steam generator inlet plenum.
The stsam generators are of the recirculating or U-tube type with vertical tubes and inlet and outlet plenums at a common elevation.
The steam generator outlet pipe is bent to vertically downward at the plenum and continues downward for about 10 feet.
At this point, the piping is bent through a 180 degree turn to vertical upward and rises to meet the reactor coolant pump casing.
Discharge from the reactor coolant pump is horizontal and at the same elevation as the reactor vessel inlet nozzles, making the run of piping' from the reactor coolant pump to the reactor vessel horizontal.
Steam Generators The Catawba and McGuire steam genetators are of the recirculating or U-tube design.
Catawba Unit 1
and McGuire Unit 2
have Westinghouse model D3 preheat recirculating steam generators, McGuire Unit 1 has model D2 prehest steam generators, and Catawba Unit 2 operates with model D5 preheat steam gen:,rators.
The operation of the steam generators for all four units is essentially the same.
Most of the feedwater enters the steam generator on the cold side of the U-tubes, then travels op the tube nest to mix with fluid from the hot side of the U-tubes.
The tuo-phase mixture from the tube nest then enters the separators wherein the steam is allowed to proceed to the steam generator upper dome, and the liquid is recirculated to the downcomer and back to the tube nest.
- 3.3 -
[
The generators differ only in the way they achieve feedwater
(
preheating and in the positioning of the primary separators.
L M
Descriotion of Emeraancy Core Coolina Svstam l
\\
The ECCS provided for the plants consists of the conventional combination of high pressure pumped injection, pressurized water
{
storage tanks, and low pressure pumped injection all connected into i
the reactor coolant piping just upstream of the reactor vessel.
I originally, the ECCS also included a direct upper head injection l
system.
In 1987 the upper head injection system was shown to be unnecessary for accident mitigation and an encumbrance to plant maintenance that adversely affected plant operability.
It has, l
therefore, been removed from the active plant systems and is no j
longer considered in plant licensing.
i l
The. high pressure injection capability of the plants is achieved through two systems: the Centrifugal Chr rging System (CC) and the Safety Injection System (SI).
The Centrifugal Charging System is j
the highest pressure syste.m of the ECCS, capable of injecting above i
normal operating system pressure, and is part of the makeup and
)
purification system during normal operation.
The system embodies sufficient redundancy such that one full train remains operative under the assumption of a
single active failure.
Emergency
]
l operation is activated automatically after receiving a safety Injection Systems "S" signal, indicating low reactor coolant system pressure or high containment pressure.
The Safety Injection System l'
operates in the middle pressure range, capable of injecting up to i
about 1300 psia.
It has two separate pumping sources with 1
I sufficient redundancy in the number of components to provide the required flow rate assuming a single active failure.
The system is also actuated by the Safety Injection Systems "S" signal.
l The accumulator system consists of four tanks, each containing about a thousand cubic feet of borated water and four hundred cubic
- 3.4 -
1
Q
\\.
I,j feet of nitrogen pressurized to 600 psi.
The tanks are connected to the reactor coolant system at the reactor coolant pump discharge via pipes.
Reverse flow during normal operation is prevented by in-line check valves.
The system is, therefore, self-contained, self-actuating and passive.
Flow into the RCS occurs whenever the reactor coolant system pressure falls below the tank pressure, f
Low pressure injection is achieved with the Residual Heat Removal System (RNR).
Normally used for cooling when the reactor is not operating, the system also serves the low pressure ECC injection function by providing borated water through the four accumulator injection nozzles.
In emergency operation, the RHR pumps initially inject water from the Refueling Water Storage Tank (RWST).
When the RWST is exhausted, the RHR pumps are aligned to take suction i
from the containment sump.
During recirculation, the injectior.
l flow is passed through a heat exchanger before being returned to the reactor coolant system.
The system contains sufficient l
redundancy such that one full train is available under a single
(
active failure.
Actuation is by the Sr.foty Injection Systems "S"
signal on low reactor coolant system pressure or high containment pressure.
In its recirculation mode, the RHR injection system provides for long-term core cooling.
In the recirculation r-i, only the RHR pump is capable of taking suction from the containment sump.
However long-term, high pressure cooling is possible because both the Centrifugal Charging pumps and the Safety Injection pumps can take suction from the RHR pump discharge and deliver coolant through their cold leg connections to the RCS.
t 2_ 1 E.11pt Parameters t
The major plant parameters and operating conditions are presented in Table 3-1.
- 3.5 -
i G!
I Table 3-1 Plant Parameters and Operating conditions i
Reactor Power 3411 Mwt Operating Pressure 2250 psia
)
Highest Allowable ?6tal Peaking (F )
2.32 1
g e
System Flow 142.6 x 10' lbm/hr
)
Core Heat Transfer Area 59973 ft f
I Average Linear Heat Rate 5.44 F.w/ft Fuel Assembly Mark-BW, 17 x 17 array i
1 Fuel Pin OD 0.374" 1
Hot Leg Temperature 622 F
{
Cold Leg Temperature 562 F Steam Generator Pressure 990 psia i
I I
I
)
I l
l 1
1 hl
- 3.6 -
i I
4.
Analysis Inouts and__Assumotions j
The McGuire and Catawba plant evaluation performed for this report j
was conducted in accordance with the B&W Fuel Company recirculating steam generator IDCA evaluation model (Reference 1) using generic j
inputs which are conservatively applicable to both facilities.
This chapter provides a brief discussion of the computer codes used in the evaluation, a
discussiors of plant parameters and assumptions, and an outline of the applicability of the generic inputs to the two facilities, M
Comouter Codes and Methods j
For the evaluation of cladding temperature transients and local oxidation, the B&W Fuel Company LOCA evaluation model consists of j
several computer codes.
Figure 4-1 illustrates the interrelation
]
of the computer codes used for the large break analyses.
The
]
REIAPS/ MOD 2-B&W code calculates system thermal-hydraulics and core i
power generation during blowdown.
The thermal-hydraulic transient calculations are continued with the REFLOD3D code to determine i
refill time a.nd core refloeding rates for the remainder of the I
The FRAP-T6-B&W code is used to determine the hot pin temperatui's response during blowdown and refill.
The BEACH code is
)
l used to determine the hot pin cladding temperature response during reflood with core flooding rates from the REFLOD3B outputs.
M Inouts and Assumptions l
The major plant operating parameters used in the LOCA codes are:
The plant is assumed to be operating in 1.
Power level steady-state at 3479 Mwt (102% of 3411 Mwt).
6 2.
Total System Flow - The initial RCS flow is 142.6 x 10 lbm/hr.
O
- 4.1 -
)
i The initial fuel pin parameters are 3.
Fuel Parameters taken from TACO 3 runs performed for the fuel assembly Darnup which produces the highest peak cladding temperature.
Sensitivity tudies discussed in Chapter 6 i
show that fuel conditions at the beginning of life are j
the most severe for large break LOCA.
j l
l 4.
ECCS - The ECCS flows are based on the worst case between the assumption of a
single active failure and the assumption of no failure.
Sensitivity studies discussed in Chapter 6 show that the condition of maximum ECCS is the most severe assumption.
l 5.
Total Peaking Factor (F,)
The maximum total peaking factor assumed by this analysis is 2.32.
)
i 6.
The moderator density reactivity coefficient is based on beginning-of-cycle conditions to minimize negative reactivity.
l 7.
The cladding rupture model is based on NUREG-0630.
l l
4.2.1 RELAP5/ MOD 2-B&W Modeling The RELAPL/h0D2-B&W computer code is used to analyze RCS thermal-
]
hydraulic behavior during the blowdown phase of a
LOCA.
RELAPS/ MOD 2-B&W, a modified version of the RELAP5/ MOD 2 code, is documented in BAW-10164 (Reference 2).
RELAP5 permits the user o
select model representation that results in a suitable finite difference model for the fluid system being analyzed.
The nodalization for the plant evaluation, shown in Figure 4-2, was e
developed in accordance with the EWFC LOCA evaluation model (Reference 1).
- 4.2 -
The control volume inputs generally consist of volume ejeometry J
l (area and height), flow-related parameters (resistance, hydraulic V
i diameter, surface roughness), primary metal heat data, and initial conditions (pressure, temperature and flow).
The non-equilibrium, non-homogenous option is used throughout, except for the core region, where the equilibrium, homogeneous option is elected in order to generate blowdown thermal-hydraulic data consi: stent with the formulation of the FRAP-T6 code.
Flow paths are defined between control volume geometric centers.
The B&W-developed SAVER f
computer code (Reference 3) is used to determine the initial pressure drops and flow distribution.
RELAP5/ MOD 2-B&W is run in l
steady-state to assure proper initialization.
%KA As shown on Figure 4-2b, the reactor core model consists of an average channel (nodes 316-326), a hot channel (nodes 328-338), a core bypass region (nodes 340-344), and a baffle gap region (nodes 346-350).
The hot channel consists of one assembly and the A
remaining 192 assemblies are modeled in the average channel.
Axially, the powered regions are represented by six equal-length segments and the unpowered regions by three equal-length segments.
Crossflow is allowed between the average core and the hot assembly using crossflow junctions.
The racistance for these junctions is developed from the experimental correlation given in Section 2.2.7 of BAW-10092 (Reference 4).
For the sensitivity studies and the break spectrum, the power distribution in the core is based on a symmetric chopped cosine with an axial peak of 1.5.
For the LOCA limits studies, where the position of the axial peak varies with the case, the power shapes and peaking correspond to the case (refer to Chapter 8).
Initial fuel temperatures, pressure, gas composition, and dimensions are based on beginning-of-life TACO 3 calculations.
- 4.3
.O U
)
Reactor power during e-LOCA transient is calculated using the RELAP5/ MOD 2-B&W kinetics option.
The rate of heat generation within " the ' fuel is computed by the code and is the sum of the fission power and fission product decay power.
The point kinetics model used in the code accounts for changes in reactivity due to the effects of' fuel temperss.no and coolant density.
The action of I
control or safety rods is 1.ot credited by the BWFC IOCA evaltation model until after the end of blowdown.
The Doppler coefficient is g
developed from and-of-life reactor physic, calculations, so that j
fuel temperature decreases during ICA maximize power generation.
c Reactor Vessel The reactor vessel model consists of a downcomer (nodes 300-308),
I
' the lower head (node 310), the core inlet plenum (nodes 312 and 314), the core outlet plenum (nodes 352 and 354), the flow W support columns (node 356), and the upper head (nodes 358 to 2..
In the dcyncomer, node 300 is centered at the core inlet nozzic,.
In the upper plenum, nodes 352 and 354 ere centered at the vessel outlet nozzles.
Reactor coolant bypass between the downcomer and upper head is taken into account by connecting node 300 to node 360.
The upper plenum model differs from that of the non-upper
- head injection plant model in 53AW-10168 due to the upper head injection-(UHI) internals structure.
The UHI flow columns and RCCA guide tubes'are modeled with a separate volume (node 356), and the i
upper plenum is divided into two volumes (nodes 352 and 354).
Reactor Coolant Locos l
The loop noding scheme is a result of the 1.p aoding and break L
noding sensitivity studies performed in Appendix A of BAW-10168.
L The four RCS loops are medeled as two by combining the three loops which will v.ot contain the break simulation together.
The 100 series nodes model the unbroken loops, and the 200 series models the broken loop.
Within each loop, the hot legs are oeparated into
- 4.4 -
g L
W1 ave--
v-- - - + -
~..,,. - - -
+ -, -,
e a-m----e
~,-.,m-i-
.\\
l s
m
- i i
4 nodes; the RSG inlet plenum (nodes 120 and 220) and RSG out.let c
i' plenum (nodes 100 and 230) are single nodes.
The RSG tubes (nodes 125 and 225) are separated into sixteen segments.
The tube flow area is based on the assumption that 10 percent of the tubes have A
been plugged on the primary side and removed from service.
The
)
cold leg reactor coolant pump suction consists of 5 nodes, and the reactor coolant pump (no6es 160 and 260) is a single node.
The.
I cold leg from the reactor coolant pump to the reactor vessel is modeled as four nodes for the broken loop and as two volumes for the unbroken loop.
Reacter Coolant PMEE.E The' reactor coolant pump performance is developed from homologous relationships adjusted for two-phase degradation based on the data j
in Table 2.1.5-2 of BAW-10164.
This is the same degradation data j
in NUREG/CR-4312 (Raference 5).
In accordance with the LOCA 3.
evaluation model, the reactor coolant pumps are assumed to trip at
/
the time of a LOCA.
y.
pressurizer l
The pressurizer model consists of three parts: the surge line (node 400), an eight-section prescurizer (node 410), and a valve model (junction 415 and node 420).
The initial condition for the pressurizer is saturated steam over saturated liquid with a void fraction specified in the interface node.
The initial inventory for the pressurizer is set to approximate the normal operating level.
'ecirculatina Steam Generator In agreement with the loop noding arrangement, the three steam generators associated with the unbrokon loops are modeled as a combined region (700 node series) with the broken loop cons.aining a
- 4.5 -
)
i i
I single ' generator (600 node series).
The preheat region 15 modeled i
l-l cas the two lower volumes of the cold side tube riser (nodes 633 and l
733)..
The five-volume cold side tube riser section (nodes 633 and 735) and four-volume hot side tube riser section (nodes 630 and 730) span the height of the generator tubes.
Two volumes (nodes 635, 640 and 735, 740) are modeled below the separators.
Nodes 650 l
and 750 are separator volumes.
The steam dome is modeled by two nodes (660, 670 and
- 760, 770).
The separator component in REIAP5/ MOD 2 acts as a steate separator and dryer. with two-phase J
fluid entering from the bottom, steam ex3 ting upward, and saturated
[
fluid going back to the downcomer through nodes 655 and 755.
Nodes p
625' through 665 and 725 through 765 form the steam generator l;
downcomer.
Auxiliary feedwater as well as 13 percent of main l,
feedwater is' supplied at nodes 645 and 745.
Nodes 675 to 680 and l
775 to 780 provide simulation of the steam lines and the safety 1
valves.
7 l
The heat structuras which model the steam generator tubes are reduced in heat transfer area by about 5000 square feet to simulate 10 percent tube plugging.
Heat structures of characteristic volume and surface' area are also included for the shell, the downcomer walls, and separatcc components.
Draak Characteristics l
l.
The four-node break location configuration is a result of the break noding sensitivity study in Appendix A of BAW-10168.
Referring to Figure 4-2a, a double-ended guillotine break is nodeled with leak paths from both nodes 270 and 275 to the containment.
For a split-type break, the leak is :nodeled as a single path that leaves node 275.
FF the double-ended break, no flow is permitted between the
' leak nodes following the break.
The switching criterion from subcooled (Extended Henry-Fauske) to two phase (Moody) discharge models is based on a leak node quality of 0.1 percent.
- 4.6 -
1 4
b r
--,,n, c
-,w, w
---a
F i
(
3; 7
Primary Metal Heat Model h
ILv)
All major components within the reactor vessel, loops and steam generators are considered.
Within a specific region, primary metals are grouped together based on similar thickness and geometry.
The exterior surfaces of the primary and secondary systems' pressure boundaries are assumed adiabatic to maximize l
- orad energy in metal slabs..
i E.CCE The accumulators, centrifugal
- charging, safety injection, and rebidual heat removal systems are modeled by nodes 900 through 915.
Injection is allowed only.'.nto the unbroken loops at the cold leg.
The injection which would take place in the broken loop is assumed to flow directly into the containment.
The accumulator is a passive system, check valve controlled, and activates automatically when the primary pressure falls below the tank pressure.
The
(
pumped systems are activated by the Safety Injection Systems signal of the ' Engineered Safety Features Activatio? System.
Appropriate
, time delays for signal generation, electrical supply startup, and injection pump startup are accounted for in the initiation of the pumped injection systems.
4.2.2 REFLOD3B Modelina 6
The REFLOD3B code simulates the thermal-hydraulic behavior of the primary system during the core refill and reflood phases of the LOCA.
The noding, shown in Figure 4-3, is consistent with the LOCA evaluation model and consists of reactor vessel and loop models.
The reactor vessel is represented by a four fixed-nodes model; nodes 1R and 2R are volumes above and below the steam-water interface in the inner vessel region, and nodes 4R and 3R represent steam and liquid volumes, respectively, in the downcomer region including the lower plenum.
The primary cyctem piping is
(~
- 4.7 -
.u a
1 j
Qi:
I g
represented by two loops similar.to the RELAPS blowdown model with I
- a. reduced number of volumes.
Values for volume geometry and flow
' path hydraulic parameters are developed from the RELAPS model.
(
?
n RELAP5 results at the and of blowdown (EOB) define the starting point for the REFLOD3B calculations.
The core and system initial conditions for REFI4D3B are derived from those of the associated RELAPS run at EOB with appropriate accounting given for the differing noding details.
In defining initial liquid inventory for REFLOD3B,. the liquid remaining in the reactor vessel at EOB in RELAP5 is placed in the lower plenum of REFLOD3B.
The initial flow rates,- gas volumes, liquid inventories, and pressures of the accumulator tanks are taken directly from RELAP5.
The reactor vessel (RV) steam volumes (1R and 4R) are initialized with saturated steam. corresponding to containment pressure, and the
' loops contain superheated steam corresponding to the containment pressure and fluid temperature of the secondary sides.
The primary metal heat structures are also taken from thu RELAPS model with the initial conditions in REFLOD3B matching those in t
RELAP5 at the end of blowdown.
The secondary side metal structures include a representation of the shell material.
In mostly stagnate regions, such as the downcomer or lower head, the heat transfer
- coefficient is based on pool boiling or natural convection to vapor.
In regions with flow, such as the hot legs or the steam generator, the~ heat transfer coefficient is set to 1000 Btu /hr-F for both vapor and liquid.
These selections insure that tha fluid leaving the steam generator is continuously dry steam superheated to the secondary side temperature.
'For a double-ended pump discharge
- break, two leak paths are modeled, one from the RV upper cawncomer (node 4R) and the other from the pump side of the break (node 28).
The pump rotor resistance is based on the locked rotor condition.
A 0.85 psi pressure drop is imposed on cold leg pipe junctions to account fot
- 4.8 -
8 e
i a
momentum losses due to steam-ECC water interaction during the n -.
()
' accumulator injection phase.
This value is reduced to 0.50 psi for the pumped injection once the accumulators have fully discharged.
The containment backpressures as a function of time from the Catawba FSAR are used in the REFLOD3B calculations.
A comparison of the calculated mass and energy releases with thoso in the l
Catawba FSAR'showet no significant differences between the BWFC and i
Westinghouse calculations.
Therefore, the containment pressures previously calculated by Westinghouse are acceptable for the REFLOD3B analysis.
4.2.3 FRAP-T6-3&W Modeling The FRAP-T6 code is used to predict thermal responses of the hot fuel rod for the blowdown and refill phases of the LOCA.
It is a boundary data driven code with inputs taken from RELAP5.
The fuel
~
rod axial ard radial nodalization is identical to that of the BEACH I
model.
Prior to transient calculations, the fuel rod is k
initialized, and the fuel temperature, fuel rod geometry and pin
('
pressure. are compared with the data p redicted by TACO 3.
The l
boundary data inputs from RELAP5 are core power, hot channel flows, hot channel enthalples, and system pressure.
FRAP-T6 is terminated l
l at the end of refill when water begins to enter the bottom of the coro for reflooding.
LM LEACH Modeling i
The BEACH code is used to detern.ine the hot fuel rod cladding l
temperature response during the reflooding phase of the LOCA.
The BEACH model consists of a not fuel rod and a flow channel with time-dependent inlet and outlet volumes to permit inputs of boundary data from the REFLOD3B calculations.
The fuel rod is axially divided into 20 segments, as shown in Figure 4-4, with variable nocal length such that each grid is located at the bottom
- 4.9 -
i if of a node and three nodes are used to cover a grid span.
The i
4' nodall' Ovn is basically that used in the code benchmarkt in J
Appendix C of the BEACH topical report, BAW-10166 (Reference 6).
Radially, the fuel pellet in divj led into 7 equally spaced mesh s'
points and two equally spaced mesh points for cladding.
The initial temperature distribution in the fuel rod and fuel pellet-clad gap conductance are obtair+d from the FRAP-T6 calculations at the beginning of reflooding.
The boundary data c
inputs from the REFLOD3B calculations are inlet and outlet pressures, flooding rate, inlet water temperature, and core 69 cay heat.
The initial to.mperature of the steam surrounding the fuel rod is set, equal to the cladding swface temperature.
If rupture occurs, BEACH is run in two passes.
First, the code is run to the time of rupture.
At this point, the cladding surface area at the location of rupture is increased, the blockage model applied, and BEACH restarted to the end of the analysis.
M' g,qmppli.gon of Plant Model with McGuire & Catawba The plant model upon which the evaluations herein are based is neither McGuire nor Catawba specific, but rather, a generic model incorporating the features of both plants in a way which assures a conservative LOCA evaluation for either plant.
For the most part, the McGuire and Catawba plants are identical at the level at which i
\\
t
.a LOCA evaluation can sense a meaningful difference in results.
Table 4-1 compares selected data from the McGuire and Catawba FSARs i
p to the values used in the generic model.
The reactor coolant system volume, as reported in the FSAR, differs between the two plants by 4 percent.
The model uses a value slightly lower than that reported for either plant.
The loop flow areas are identical L
-and the flow area through the steam generator tubes differs by only 2 percent.
Therefore, the pressure distribution around the system is essentially identical.
Furthermore, the positioning of the
- 4.10 -
s
't v
1
'[Y ;
major-components relative to each other is the same.
The Reactor m
Q' \\
Trip System and the Engineered Safety Features Actuation System of l
the two plants function the same.
Therefore, if there are differences in the plants which can affect the results of a LOCA, they will lie in the attached systems.
The secondary or auxiliary system which can affect the results of a LOCA evaluation are:
l-l The Containment System The Seconden Coolant System The High Pressure ECCS pumped systems L
The Accumulator System The Low Pressure ECCS pumped system h
The containment system can affect a LOCA result by imposing a l
different back pressure (containment pressure) for the reflooding calculations.
'With higher containment pressure - during reflood,
' (J reflooding rates are higher and reactor core cooling is enhanced.
The containment system for both facilities is of the ice condenser-type of the.same design and sizing.
Thus, for the same accident, the containment pressure will be th ' same for the two plants.
Furthermore, the containment pressure used in the LOCA calculation is already so low, because of the ice cendenser design, that it would take a major deviation between the plants to cause a change
}
in containment response significant to a LOCA evaluation.
i The secondary system can affect a LOCA calculation through a change in the safety valve set point.
During reflood, liquid carry-over from the core is vaporized in the steam generators and superheated to the saturation temperature of the secondary side.
The evaluations set the secondary conditions according to the safety valves.
If the temperature of the secondary coolant is reduced by a change in safety valve setting, the loop flow in REFLOD3B will be less superheated and its specific volume, lower.
This flow will,
- 4.11 -
.y
^
^^
^
,e
- t, 1
I
~therefore, be easier to rove arcund the ' loop to the
- break, resulting in h 3ower loop pressure drop, higher core flooding
- rates, and colder cladding temperatures.
The only substantial differenc6 between the steam generators of these four plants is in the manner of achieving feedwater preheating, a difference that will not influence a INCA result.
The sa!.ety valve set points for the two plants are the same.
i The high pressure puniped injection is not of substantial import for a large braak LOCA since the accumulators and low pressure systems provide more injection than can be effectively used in cooling the plant.
Notwithstanding, the high pressure pumped injection flows used in the analysis were si.ected as the minimum of the flows of all four units tor the minimum ECCS cases, and the maximum of the flows of all four units for the maximum ECCS cases.
"~
The plants do differ in tl.a accumulator system provided.
The McGuire accumulator liquid volume is lower than Catawba's, and tne pressurizing gas volume is higher.
Also the surge line resistance for McGuire is higher than for Catawba.
As the downcomer is full or nearly so before the accumulators empty, the different liquid volumes are unlikely to affect the LOCA results.
The higher gas volume and higher line resistance for'McGuire would tend to off-set each other, making it difficult to determine which configuration would produce the most conservative LOCA result.
Therefere, a sensitivity study, presented in Chapter 7,
was performed.
Both configurations were evaluated; the Catawba configuration produced slightly higher cladding temperatures than the McGuire j;
configuration.
The accumulator used in the evaluation was based on the Catawba plant.
The low pressure pumped injection systems for both plants supply I
water at rates above those that can be et~ i'ectively utilized in core cooling during LOCA.
This is directly observed in reflood where, under a
single failure assumption, the downcomer runs full L
- 4.12 -
l l
I continually with about 20 percent of the ECCS injection spilling e
from the downcomer through the break.
That water would be available to enter the core if higher flooding rates could be achieved, but the flooding is controlled by other plant features.
Nonetheless, the low pressure pumped injection flows used in the analysis were selected as the minimum of the flows of all four units for the minimum ECCS cases, and the maximum of the flows of all four units for the maximum ECCS cases.
As has been shown, the McGuire and Catawba plar.ts are essentially
~
identical to one another at the level of design parameters that can affect the results of a 14CA.
The single deviation which does exist between the plants has been the subject of a sensitivity study (Chapter 7) and shown to affect the cladding temperatures only slightly.
Nevertheless, the most conservative of the two accumu'.ator configurations has been used in these evaluations.
Therefore, the studies and conclusions presented in this report are conservative for application to t.ll four units of the Duke Power Company's McGuire and Catawba facilities.
- 4.13 -
1 I:
V TABLE 4-1 COMPARISON OF IOCA MODEL GEOMETRIC VALUES l
TO PLANT FSAR DATA OVERALL' SYSTEM Paranciar Plant Model Catawba McGuire l
Total System Volume Including Pressurizer, ft 11,900
- 12,516 12,040 3
1 L
. Total System Liquid Volump Including Pressurizer, ft 11,100,
11,774 11,298 1
L REACTOR VESSEL i
Parameter Plant Model Catawba McGuire j
L, RV I.D.
at Flange, in 167.0 167.0 167.0 0
l RV I.D.
of Lower
- Shell, i'.1 -
173,0 273.0 173.0 1
[
RV Inlet Nozzle L
I.D.,
in 27.5 27.5 27.5
.RV Outlet Nozzle 1
I;D., in 29.0 29.0 29.0 J
Overall Height of Vessel and Closure Head, ft NA 43.8 43.8 REACTOR COOLANT LOOP COMPARISON Parameter Plant Model Catawba McGuire Hot Leg Pipe I.L.,
in 29.0 29.0 29.0 Cold Leg Pump Suction Pipe I.D.,
in 31.0 31.0 31.0 Cold Leg Pump Discharge Pipe I.D.,
in 27.5 27.5 27.5 3
Pump Volume, ft 78.6 78.6 57.0 i
i
^
- 4.14
t TABLE 4-1 COMPARISON OF LOCA MODEL GEOMETRIC VALUES l
7s x--)
TO PLANT FSAR DATA (continued)
't RECIRLULATING STEAM GENERATOR Parameter Plant Model Catauba Et,Guire U-Tube Outer Diameter, ia 0.75 0.75 0.75 Tube Wall Thickness, in 0.043 0.043 0.043
[
Number of U-Tubes 4,568 4,568 4,674 Heat Transfer Area, f '.I 48,300 48,300 48,000 The valites provided for the table are without tube plugging j
The actual evaluation used reduced values to account for the assumed degree of tube plugging.
This value is not utilized in the plant modeling but has been supplied for the plants to illustrate the similarity netween the McGuire and Catawba units.
- 4.15 -
lE w
p n ;-
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- ' FIGURE 4-1 LARGE BREAK ANALYSIS CODE INTERFACE
-[NITIAL RC SYSTEM AND CORE PARAMETERS INITIAL CORE PARAMETERS Y
RELAPS/ MOD 2-B&W MASS AND
> ENERGY RELEASE O < TIME < EOS
~
l-CORE STORED ENERGY SYSTEM MASS AND ENERGY ACCUMULATN MASS AND ENERGY STEAM GENE 81ATOR MASS AND ENERGY CORE DRESSURE CORE ENTHALPY s
CORE ? MSS FLUX y
REFLOD3B CONTAINMM EOB < TIME : EOE MINIMUM BACKPRESSURE FLOODING RATES Y
Y CORE FRAP-TB-B&W PARAMETG4S BEACH AT EOAH 0 < TIME < EOAH
?
EOAH < TIME < EOE l
Y METAL WATER REACTION
> HOT PIN THERMAL RESPONSE PEAK CLADDING TEMPERATURE
- 4.16 -
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'l i
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'wI'
.' J FIGURE 4-2a RELAPS/ MOD 2-B&W'LDLOCA NODING DIAGRAM REACTOR COOLANT LOOFS
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CENTRIFICAL CNARc1No INTACT i_DDP
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-Md 915 l RESIIlUAL HEAT RENOVAL
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FIGURE 4-2b RELAP5/ MOD 2-B&W LBLOCA NODING DIAGRAM REACTOR CORE i
e 364 360 f
h 358 h
i 300 l
356 354 TD TD n
I Y
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> NDDE NDDE <
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=*
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- 4.18 -
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i FIGURE 4-4.
BEACH AND FRAP-T6-B&W NODING DIAGRAM
?'
FOR MARK-BW FUEL ASSEMBLY NODE ELEVATION 12.00' 20 11.28' 19 l
g 10.56' 18 9.99' 17 9.42' 16 4
8.85' 15 8.28' 14
=
7.71' 13 4
7.14' 6.57' 11 6.00' 10
- p 5.43' 9
4.86' t
8 l
4.29' 7
3.72' 6
3.15' 5
2.58' 4
- p 2.01' 3
1.34' 2
0.67' i
- ' GRID NODE 0.00' l,
L 1
~.
g L
.y..
i, L 71 5.
Evaluation Model Chances k
1-The original version of the BWFC large break LOCA evaluation model (Reference 1) was submitted to the NRC in July, 1988.
Following f
that submittal, it became evident that the original evaluation model code package produced excessive conservatism in calculating j
certain aspects of the retlooding heat transfer processes.
Consequently, improved methods were developed.
The changes were incorporated in two steps:
The first step involved changes to the hEACH code to refine the spacer grid effects model and to include rupture-blockage effects.
The'second step ontailed modifying BEACH
.to include cladding metal-water reaction and active calculation of clad swelling and rupture using the same cladding behavior modeling as RELAPS/ MOD 2-B&W and FRAP-T6-B&W.
This final modification thus allows BEACH to be used alone during the reflooding period to calculate peak cladding temperatures.
These new models have been benchmarked aghinst appropriate data and are being submitted to the i
l-NRC as revisions to the ECC3 evaluation model topical report and to the affected code tcpical reports.
L The analyses presented and discussed in this report were, of necessity, perf ormed previous to or in parallol with the modeling L
revisions.
The sensitivity studies described in Chapter 6.C were done using the original ECCS evaluation model as reported in Reference 1.
The plant-specific sensitivity and break spectrum analyses reported in Chapter 7.0 used the codes as applied in the original evaluation model but included the modifications to the DEACH spacer grid ar.d rupture blockage models.
Finally, the LOCA limits cases presented in Chapter 8.0 were executed with the final BEACH version, which included the clad swelling,
- rupture, and metal-water reaction models.
The remainder of this chapter delineates the differences among the methods used and discusses the effects of these differences upon the calculated results, bQ
- 5.1 -
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1 i
M Revisions to the BWFC LOCA Evaluation Model Oi Revision 0 of the BWFC LOCA evaluation model topical report (BAW-
)
10168) was submitted to the NRC in July, 1988.
Volume 1 of that report contained the large break I4CA model and Volume 2, the small break-model.
The model changes to be discussed in this section affect only the large break modeling.
The code package and interrelationships for the original moCal are shown in Figure 5-1.
RELAP5/ MOD 2-B&W is used to calculate system thermal-hydraulics and core power generation during blowdown.
REFLOD3B determines the l
refill tims and the core reflooding hydraulics for the remainder of the transient.
The BEACH code is used to determine cladding J
surface' heat transfer coefficients (HTC) with core reflooding hydraulics from REFLOD3B.
FRAP-T6-B&W completes the analysis by using RELAP5 and EEACH inputs to evaluate the hot rod cladding j
temperature for the entire transient.
This model is fully described and benchmarked in Revision 0 of the evaluation model topical report (Reference 1).
O' Initial Modifications to BEACH l
l The BEACH code version represented in the original evaluation model was found to contain two limitations.
- First, the fine drop vaporization model employed to predict the shattering of entrained
[
l, liquid into small droplets..at the spacer grids (and the controlling l
input parameter for-that model) was inaccurate--and overly l
conservative--in terms of sensitivity to the grid blockage ratio.
Second, the modeling of clad rupture effects was limited to those l
of increased cladding surface area at the rupture site.
The additional effects, most notably those beneficial to interphase heat. transfer and clad cooling above the
- rupture, were not represented.
The spacer grid effects model as originally formulated was found to be appropriate to account for the effects of simple spacer grids,
- 5.2 -
p l
i f
1 as had been used in the
- FLECHT, FLECHT-SEASET, and other.
l
. O).
experiments.
Applying. the model to grids with higher blockage f
ta ratios, particul'arly to the mixing vane grids used in the Mark-BW design',.was found to be both crude and severely conservative.
j 5
Simply put, the model was not capable of representing the mixing vanes other than ' as an increase in the t.hickness of the grid
. straps, that is, as a simple increase in the total blockage ratio.
In addition, in r:nsveloping the original model, it had been decided thLt ' the occurrence of rupture and the impact of the resultant channel blockage on intelphase heat transfer would not be modeled.
The first set of revisions to the evaluation model removed these limitations.
The grid fine-drop vaporization model was upgraded so that the droplet impact Webor number is calculated as a function of axial position within the grid, and a new correlation for the fine drop
. vaporization constant, c,, was developed.
Second, the procedure for running BEACH was altered such that a blockage effects model, similar to a grid model, could be user-initiated at the location of
- rupture when rupture occurred.
This required stopping BEACH near the time of rupture, e:tecuting FRAP-T6 with pre-rupture inputs from BEACH to determine.the time and location of rupture, and restarting BEACH, post rupture with the model and appropriate inputs active.
'the first set of revisions preserved the code interrelationships shown in Fjgure 5-1.
Final Revisions As work with the first revisions progressed, the results suggested that excessive conservatism remained in the model.
It was found that the cladding temperatures calculated in FRAP-T6 were higher than those calculated in BEACH, and that the difference was greater than that attributable to the metal-water reaction calculated by FRAP-T6 and not in BEAL 1.
The difference was primarily the result of the code interrelationships selected for the package gr'igure 5-
- 5'3 -
V
g t
i i
1).
FRAP-T6-B&W does not model axial conduction along the fuel pin.
With 200 F to 300 F cladding temperature differences between ruptured and unruptured locations, the effect of axial conduction on the solution can be significant.
The' final model change was to replace the FRAP-T6-B&W code during reflooding with the BEACH code.
This required the addition of metal-water reaction and clad strain a
and rapture models that were already available in RELAPS/ MOD 2-B&W (the parent code for BEACH).
In addition, the fine-mesh rezoning logic in the PEACH conduction solution was modified to use the maximum requested fine mesh near the quench front.
A constant heat transfer coefficient of 528 2
Btu /hr-ft -F below the quench front was adopted to stabilize vapor generation rates after quenching.
With these
- changes, the code applications and interface relationships became those of Figure 4-1.
FRAP-TG-B&W is run only through the refill period, and BEACH, initialized to FRAP-T6, is started at the beginning of reflooding and run to the end of the analysis.
If rupture occurs, BEACH is run in two passes.
- First, BEACH is run to the time of rupture.
At this point the cladding surface area at the location of rupture is increased, the blockage model applied, and BEACH is restarted to the end of the analysis.
L M
Effects of Evaluation Model Revisions
{
The ' changes to the evaluation model described in the foregoing section affect the calculated LOCA results only in the hot pin portion of the calculation and only during the reflood phase of the transient.
The RELAPS/ MOD 2-B&W and RET' LOD 3 B computer codes are completely unaffected by the changes, and the FRAP-T6-B&W analysis of the blowdown and refill phases is not impacted.
For the most part, the original evaluation model remains intact, and the trends and conclusions derived from it continue to be applicable.
- 5.4 1
\\
p L
r i
f
. Effects of.Scacer Grid and Ruoture Modifications
{
V Because the grid model cnange was to make the model more responsive to.the' type of grid and thus, generate a more accurate solution, the effect of the modification depends on the type of grid being
[
modeled.
The original grid model was directly tied to spacer grida f
of conventional design (simple grids).
Such grids have blockage factoru (fraction of the flow channel area blocked by the grid) of about 30 percent.
The revised model is formulated to be the same as the original' when applied to simple grids.
- Thus, results produced by BEACH in benchmarks of experiments that used plain or simple spacer grids are unaffected by the model changes.
For grid designs with higher blockage ratios (such as zjrcaloy grids or mixing vane spacer grids), the blockage factors range from 45 to 60 percent and the blockage is distributed along the direction of flow.
(For. example, there may be a step to 40 percent blockage followed by a step to 60 percent blockage.)
The effects of such grids on interphase heat transfer were found to be barely, if at all, distinguished in the original model from the effects of grids with. the 30 percent blockage factor.
It was obvious that the higher blockage grid should interact with more droplets and should cause a_ larger fine-drop effect than the lower blockage grid.
The revised model accounts for this, and the interphase heat transfer, particularly due to fine-drop ef fects, is substantially increased
(
for the higher blockage factor grids.
Depend'.ng on the particular grid design, the amount of interphase heat cransfer may increase l
considerably.
Similarly, if a grid does not have a higher blockage factor, the amount of interphase heat transfer will be redisced by the new modeling.
The immediate effect of this change in the calculation is observed l
in the " grid nodes" of the hot channel simulation.
At these locations, the vapor temperature can be reduced by several hundred l
- degrees, and that can be directly reflected in the cladding l
temperature.
- However, this observation would apply to any
- 5.5 -
l:
I 1
l l:
representative grid model, and the local effect is not the one of L
interest.
Because the interphase heat transfer has been improved h.
at the grid, the downstream locations within the span experience increased flow of the steam phase, and the temperature of the steam j.
is reduced.
Thus, heat transfer downstream of the sptcar grid is improved.
This interaction propagates downstream for several feet before it is ' essentially washed out.
By the time (length) the 1
effects of'one grid have diminished, the flow oncounters the next I
I grid.
The grid spans are typically 1.6 to 1.9 feet.
Thus, the l
effect of ' this model change on the results in this report (the MARK-BW fuel assembly uses high blockage grids) is an overall l.
. decrease in cladding temperatures.
This effect will pertain to the 1
L.
entire length of the fuel pin, but the timing may differ between the upper and lower elevations.
Trends observed with the original I
model should be preserved in the revised model, but overall, the 1
temperatures will be lower.
The effect of improving the interphase heat transfer at the location of rupture is nearly identical to that of the grid model.
Here, the. blockage factor is determined by the amount of swelling L
predicted by.the rupture model.
If that swelling is high, there l
will be a substantial' increase in the interphase heat transfer.
If L
the swelling is low, there will be very little change.
For the cases considered here, the blockage factors from rupture are on the order of 50 to 60 percent, and the amount of resulting interphase j'
heat transfer is significant.
As with the grid model, the effect is on the interphase heat transfer and can directly impact the ruptured node cladding temperature.
At other locations on the pin, the higher and colder steam flow from the ruptured location will provide greater cooling for nearly two feet downstream.
Beyond that, there is no disr2ernible effect.
l
'Near the time of rupture, the node of highest power that does not 1
j contain a spacer grid will tend te have the highest temperature and thus become the ruptured node.
With this model causing cooling in
- 5.6 -
I g
w
-w-y m-w.y
the rupture node and in the next one or two nodas, the model
()
creates the possibility that the rupture location and the location of peak cladding temperature will be separated by one grid span.
Previously, the peak cladding temperature was predicted to occur in one of the nodes adjacent to the ruptured node.
With the revised 1
.model, the node immediately above the ruptured node is provided with additional cooling to the extent that the node centered in the grid span above the location of peak power will produce a very
[
close, if not higher, cladding temperature.
Taken together, the grid and rupture model changes produce a generally increased cooling of the entire fuel pin, a more dramatic temperature reduction in the grid nodes and the ruptured node, and i
l l
allow a shift in the peak cladding temperature to a grid span l
removed from the location of peak power.
These changes do not l
necessarily affect trends produced by the original model.
Where l
the original trends pertained to behavior of the fuel pin globally, j.
they will remain applicable.
However, those trends that were local
(
effects, particularly near the rupture location, may be altered.
l Effects of the Final Chances l.
The effects of incorporating axial conduction into the hot channel calculation are substantial at the ruptured node and in the nodes i
adjacent to it.
Without axial conduction, the ruptured node can be 200 F to 300 F below the adjacent nodes.
With axial conduction, sufficient heat is transterred to the ruptured node that it runs only 100 F to 150 F below the adjacent nodes.
Those nodes in turn are up to 100 F to 150 F cooler than calculated in the original model.
The model also affects the grid nodes and adjacent nodes in the same way, though the effects are less marked.
That is simply because the temperature gradients between the grid nodes and
,O V
- 5.7 -
1 I
. rieighboring nodes are not so severe.
The net effect, though, is the same:
axial conduction reduces the temperature differences between adjacent heat structures.
]
The.other model changes in the final set were refinements to remove spurious oscillations, driven by switching in the logic. They are found to produce essentially no difference in peak cladding l
temperatures.
Post-peak cooldown is smocthed, by the use of 32-point fine-mesh rezoning, for nodes at and above the midplane.
The cumulative effect of the model changes on the calculated LOCA results is an overall reduction in predicted cladding temperatures.
Post-rupture temperatures at locations adjacent to the ruptured node are noticeably reduced.
With the additional cooling in the rupture node, and in the adjacent nodes through axial conduction, the model has the effect of separating the rupture location and the location of peak cladding temperature by one grid span.
The trends of sensitivity studies observed with the models still pertain if those trends were developed over the entire fuel pin.
Trends derived from specific locations on the fuel pin may need to be reevaluated for impact and may require confirmation.
- 5.8 -
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"V CODE INTERFACES FOR LARGE BREAK LOCA-i NITIAL RC SYSTEM j
AND i
CORE PARAMETERS I
INITIAL CORE I
i' PARAMETERS
]
U j
> ENERGY l
RELEASE 0 < TIME < EOS CORE STORED ENERGY SYSTEM MASS AND ENERGY ACCUMULATOR MASS AND ENEROY STEAM GENERATOR MASS AND ENERGY
.I)
CORE PRESSURE
(/
. CORE ENTHALPY CORE MASS FLUX j f REFLOD3B CONTAINMENT EOB < TIME < EOE MINIMUM BACKPRESSURE FLOODING RATES t
I Y
Y CORE FRAP-T6-B&W PARAMETERS BEACH 4
AT EOAH 0 < TIME < EOAH EOAH < TIME < EOE EOAH < TIME < EOE 4
i REFLOOD HEAT TRANSFER f
COEFFICIENTS METAL WATER REACTION HOT PIN THERMAL RESPONSE PEAK CLADDING TEMPERATURE
- 5.9 -
. ~.
r I
6.0 Sensitivity Studies
,, - ~
LOCA evaluations require that a substantial number of sensitivity studies be performed with the evaluation model in order to establish model convergence and conservatism.
Most of the studies upon which the evaluations in this report are based are generic and were documented in the evaluation model
- report, BAW-10168 (Reference 1).
Studies such as break spectrum and worst case ECCS configuration are considered plant-specific and are documented in this report.
This chapter provides a discussion of the generic sensitivity studies from the reference evaluation model report that have been applied.
The plant-specific sensitivity studies are presented in Chapter 7.
_al Evaluation Fgdel Generic Studies of the sensitivity studies presented in the original evaluation model topical report (Reference 1), the majority are generic and
(
would apply to any plant evaluated.
Those studies considered l
generic each demonstrate results that are characteristic of the i
evaluation model--the codes and interfaces--and that are not plant l
l dependent.
An example of this is the RELAPS/ MOD 2-B&W time step study, which demonstrated that the automatic time step selection in 1
l.
RELAP5/ MOD 2-B&W would produce converged results.
This l
I l
demonstration need not be repeated for plant-specific applications l
wherein the modeling techniques used are represented by those in l
the evaluation model studies.
The following is a listing of the l
sensitivity studies considered to be generic, with a discussion of why the conclusions of the study remain applicable for this i
applications report.
In particular, the continued applicability is discussed relative to the later modifications to portions of the l
evaluation model.
For convenience of review, each discussion is referred to the section in the evaluation model report where the study is, documented.
- 6.1
i RELAPS/ MOD 2-B&W Time Steo Study This study (BAW-10168, Appendix A, Section A.2.1) verified that for i
light water reactor geometry, the RELAPS time step controller governs the code solution sufficiently to assure converged results.
Alternate system designs within the group to be covered by the evaluation model will not change that result.
The differences I
between the evaluation package used for the study and the evaluation model used for this report do not affect the RELAPS/ MOD 2-B&W code nor the blowdown period of the evaluation.
Therefore, the study remains applicable.
RELAP5/ MOD 2-B&W Loon Nodina Study This study (BAW-10168, Appendix A,
Section A.2.2) verified the general noding requirements within the loop for recirculating steam generator plants.
In conjunction with the break noding study, the results can be applied to the separate regions of the hot leg, the steam generator, and the cold leg.
Alternate system designs within the group to be covered by the evaluation model will not change the noding requirements.
The differences between the evaluation package used for the study and the evaluation model used for this report do not affect the RELAPS/ MOD 2 B&W code nor the blowdown period of the evaluation.
Therefore, the study remains applicable.
RELAP5/ MOD 2-B&W Break Nodina Study
(
This study (BAW-10168, Appendix A,
Section A.3.1) verified that L
hydraulic stability is achieved by providing at least one control l
volume in the pipe between any adjacent component and the break l
node.
The break noding study is applicable to all plants covered L
by the evaluation model.
The differences between the evaluation package used for the study and the evaluation model used for this report do not affect the RELAPS/ MOD 2-B&W code nor the blowdown j
period of the evaluation.
Therefore, the study remains applicable.
- 6.2 -
l
.---.--r
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--...-y-,
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.dD--
RELAPS/ MOD 2-B&W Pressurizer Location Study Although - the assumption placing the pressurizer in one of the intact loops was somewhat conservative, this study (BAW-10168, j
I Appendix A, Section'A.3.2) showed that there is little difference in results when the pressurizer is modeled in the broken loop.
The lack of sensitivity to pressurizer location is expected to hold for all designs covered by the evaluation model and this study will not be repeated.
The differences between the evaluation package used for the study and the evaluation model used for this report do not
)
affect the RELAP5/ MOD 2-B&W code nor the blowdown period of the i
evaluation.
Therefore, the study remains applicable.
j RELAPS/ MOD 2-B&W Core Crossflow Study This study (BAW-10168, Appendix A,
Section A.3.4) verified that l
cross flow in a light water reactor is limited and does not alter O
the course of a LOCA evaluation substantially.
The study is l_
y dependent only on the very basic aspects of the fuel design, which are consistent across the range of designs considered by the evaluation model.
The differences between the evaluation package used for the study and the evaluation model used for this report do not affect the RELAP5/ MOD 2-B&W code nor the blowdown period of the l
evaluation.
Therefore, the study remains applicabic.
l-l l
RELAP5/ MOD 2-B&W Core Nodina Stud _y l
l l
In conjunction with the core cross flow study, this study (BAW-10168, Appendix A, Section A.3.5) verified that the modeling of a light water reactor core in six axial segments with a hot and an average channel provides sufficient spacial detailing for both model convergence and result r mracy.
As the basic core arrangement and fuel design an ae altered across the range of designs to be considered, the resu of the study are applicable
- 6.3
y i
to all plants considered by the evaluation model.
The differences i
between.the evaluation package. used for the study and the evaluation model.used.
for this report do not affect the i
RELAP5/ mod 2-B&W code nor the blowdown period of the evaluation.
Therefore,.the study remains applicable.
REFLOD3B Primary Coolant Pumn Rotor Resistance Study This study (BAW-10168, Appendix A,
Section A.2.4) showed a
considerable reduction in flooding under a locked rotor assumption.
+
The study affirms the generally accepted. data on loop resistance effects on reflooding rates and is applicable for all plant types covered by the evaluation model.
The differences between the evaluation package used for the study and the evaluation model used for this report do not affect the REFLOD3B code.
Therefore, the study remains applicable.
FRAP-T6-B&W Time Sten Study This study (BAW-10168, Appendix A, Section A.3.6) verified that the time step selection for FRAP-T6-B&W provided converged results for the spatial detail modeled in the base runs.
Because the spatial detail required for the FRAP-T6-B&W model is not be altered for the other' designs covered by the evaluation model, the study remains l'
' valid for all designs.
The differences between the evaluation L
. package used for the study and the evaluation model used for this report only shorten the period of the LOCA for which the FRAP-T6 code is applied to the blowdown and refill phases.
As the model i
l changes do not directly affect the FRAP-T6 code nor the blowdown and refill periods, the study remains applicable.
l-FRAP-T6-B&W Radial Fuel Seamentation Study This study (BAW-10168, Appendix A, Section A.3.7) verified that the number of solution points selected for radial representation of the
- 6.4 l
_ _ _ _ _, _ _. _ _... ~., _ _ _..
p i.,
' ~ ~
i fuel pin used by the base FRAP-T6-B&W model was adequate.
The i]
study is dependent only on the very basic aspects of the fuel design, which are consistent across the range of designs considered by the evaluation model.
The differences between the evaluation package used for the study and the evaluation model used for this report only shorten the period of the LOCA for which the FRAP-T6 code is applied to the blowdown and refill phases.
As the model I
changes do not directly affect the FRAP-T6 code nor the blowdown 1
and refill periods, the study remains applicable.
M Confirmable Sensitivity Studies In addition to the generically applicable studies, some of the studies performed for the evaluation model are considered confirmable.
These studies remain valid under most but not all j
circumstances.
The following is a listing of such sensitivity studies with a discussion of why the conclusions of the study can be applied to the McGuire and Catawba evaluations, why they remain j
valid in light of the alterations in the evaluation model, and a reference to the section in the evaluation model report where the study is documented.
RELAPS/ MOD 2-B&W Pumo Dearadation Study This study (BAW-10168, Appendix A,
Section A.3.3) established a most severe pump degradation multiplier by altering the pump effects on the core flow.
The study can be applied to all plants
)
which experience similar LOCA core flow histories durirg blowdown.
t 1
The core flow shown for the McGuire/ Catawba evaluation base case (a 1
guillotine break at the pump discharge with a Cd of 1.0) in Figure l
l-6-1 is of the same general form as the core flow for the reference 1
i case in the evaluation model report, also shown in Figure 6-1.
i l
Therefore, the applicable debilitating effect of the degradation has been preserved across the plant design and the results of the study are applicable to the McGuire and Catawba evaluation.
The
- 6.5 -
l l
differences between the evaluation package used for the study and the evaluation model used for this report do not affect the REIAPS/ MOD 2-B&W code nor the blowdown period of the evaluation.
Therefore, the study remains applicable.
REFLOD3B Loon Nodina Study This study (BAW-10168, Appendix A,
Section A.2.3) verified the noding detail used in the REFLOD3B code.
It is applicable to plants with a one-to-one correspondence of hot and cold legs, such as the McGuire and Catawba units.
A separate study is required 1
only for severely altered loop
- designs, such as the B&W or Combustion Engineering 2-by-4 designs.
The differences between the i
evaluation package used for the study and the evaluation model used for this report do not affect the REFLOD3B code.
Therefore, the study remains applicable.
Time-in-Life Study This study (BAW-10160, Appendix A,
Section A.3.9) showed the initial fuel temperature to be the dominant fuel burnup-related parameter affecting the LOCA results.
There are three fuel parameters that vary significantly with fuel burnup: initial fuel energy, internal pin pressure, and oxide thickness.
The initial stored energy of the fuel, given by the volume-average temperature at any specific local power level, changes with burnup.
The
)
pattern observed is that the fuel temperature at a specific local power may increase sharply during the first few days of operation, but it will then decrease until relatively high burnups, after
.l which a slow increase may occur.
Calculations with TACO 3, Figure 6-2, have shown that the initial rise in temperature takes place virtually at the start of fuel operation.
From that point, there remains a long burnup period with decreasing fuel temperatures.
The fuel pin internal pressure undergoes a continuous increase with
- 6.6 -
L 1
h burnup.
The third parameter, of somewhat secondary importance, is d
the oxide thickness, which increases with fuel burnup.
t While the interactions among these parameters in a I4CA analysis are complex--depending somewhat on the complexity of the model being used--the sensitivities of the results to each can be generalized:
The higher the average fuel temperature at any location on a fuel pin at the start of a LOCA, the higher the resultant cladding temperatures at that location during the LOCA.
t The higher the internal pin pressure at the initiation of the IhCA, the more likely or earlier the occurrence of rupture will be.
This may or may not have any bearing on the peak cladding temperature.
l The thicker the initial oxide layer of the fuel pin, the lower the cladding temperatures for cladding that has temperatures above about 1800 F.
The oxide effect results from metal-water reaction and is relatively inconsequential for cladding that is not heated
[
to 1800 F.
1 The effect of
- rupture, and
- thus, pin pressure on cladding temperature is model-dependent.
For the computer codes and methods with which the original evaluation model sensitivity studies were performed,.a rupture after the end of blowdown had essentially no effect on the cladding temperature at any location other than at the rupture.
At that location, the effect of rupture was to allow
(
the gap and the cladding surface area to increase, both of which act to decrease the cladding temperature.
Thus, the effect of rupture during reflooding was to reduce the temperature increase of i.
the cladding at a location that, clad rupture notwithstanding, 1
l would not have been the peak temperature location.
A rupture during blowdown was modeled to have one additional effect:
The blowdown rupture would cause additional flow resistance to be f
calculated in the hot bundle.
This would cause some flow diversion l
away from the hot assembly and result in poorer blowdown cooling i
for both the ruptured location and the rest of the hot pin.
The rupture location would, nonetheless, remain relatively cooler than
- 6.7 -
l
i I
other locations on the pin, but those other locations would have j
higher cladding temperatures because of the blowdown rupture.
These results and observations lead to the general conclusion that, for fuel pin temperatures and internal pressures that do not produce a blowdown rupture, the most restrictive burnup for the LOCA to be analyzed is the one with the highest initial fuel
]
temperature.
If blowdown ruptures are to be considered, a more detailed study is required to determine the most severe combination (s) of temperature, pin ' pressure, and oxide thickness.
This is, in fact, what was observed in the sensitivity study performed for the evaluation model report (Reference 1).
This i
conclusion is independent of the fual and plant designs and thus applies to any fuel or plant of reasonably similar design that is i
determined not to encounter a blowdown rupture.
These conditions, and thus the conclusion,-apply to the McGuire and Catawba units as will be further described.
l The foregoing discussion is derived from the codes and methods used in the sensitivity studies presented in the original evaluation model topical report.
The recent evaluation model changes do not directly affect the occurrence of rupture during blowdown.
The changes do,
- however, extend the calculated effects of rupture during reflood.
In the latest model, all of the previously modeled aspects of rupture are included, with additional effects.
The l.
blockage of the flow channel by the expanded clauding will interact l
with the entrained liquid--droplets--in the dispersed flow causing l
further droplet shattering.
This produces an increase in the l
calculated interphase heat transfer and thereby causes lower vapor temperatures and higher steam flows at the rupture location.
Both l
of these propagate up the bundle channel to promote better cooling in the adjacent nodes.
Additionally, because BEACH models axial
. conduction in the cladding, once the ruptured node expands to as much as twice the heat transfer area, the resulting cladding temperature gradients permit substantial transfer of heat from the
- 6.8
p I
w
' c
,q-unruptured locations on either side of the rupture to the ruptured
]
V location.
Both the increased interphase heat transfer and the axial conduction produce additional cooling in nodes beyond the immediate rupture location.
1 Considering the effects of cladding rupture in the latest model separately, it could be inferred that, because a predicted rupturn improves the cooling and reduces the temperature increase during reflood for a portion of.the cladding beyond the rupture site, l
earlier prediction of rupture would result in lower calculated peak cladding temperatures.
Taking this to be the case, it further develops that the relationship between pin internal pressure--the hoop stress leading to rupture--and fuel average temperature is L
such that the pin pressure is lowest at the burnup conditions producing the highest fuel average temperatures, namely at BOL conditions.
The coincidence of these conditions continues to l'
support the conclusions related to burnup drawn from the studies of l
the original evaluation model:
For fuel pin pressures and
(
I temperatures not producing rupture during blowdown, the limiting i
w/
I L
LOCA initial conditions are those at beginning-of-life.
l l
The analyses. presented in this report are done using TACO 3-calculated fuel inputs for BOL conditions.
To establish BOL as the l
worst
- case, the possibility of a
blowdown rupture must be L
precluded.
It is evident that a rupture during blowdown does not l
occur for the BOL case in the analyses presented in this report.
Confirmation of the same for end-of-life (EOL) conditions can be accomplished by comparing TACO 3 EOL fuel parameters to those used in the time-in-life study in the original evaluation model report.
The TACO 3 volume-averaged fuel temperature at 12.6 kw/ft at 60,000 mwd /mTU is slightly less than 1800 F.
The internal pin pressure is limited by technical specifications to the operating system pressure of 2280 psia.
The corresponding conditions for the sensitivity study at 60,000 mwd /mTU were 1911 F and 2280. psia, and
- 6.9 -
,w,-,
.,,e
-u
l i
the predicted rupture occurred well into the reflood phase.
As the i
evaluation model has not been altered for the blowdown and refill i
phases, these results continue to apply, and the new Taco 3 data will not produce a blowdown rupture for burnups as high as 60,000
)
mwd /mTU.
Figure 6-2 shows representative volume-averaged fuel temperatures j
as calculated by TACO 3 at three different local power levels as a l
function of pellet burnup.
The power level marked "high local power" is sorewhat above that used at the peak power locations in these analyses.
As shown, the volume-averaged fuel temperature decreases substantially for early burnups and continues to decrease to 30,000 Mwd /Mt after which it undergoes a slight increase.
By 60,000 mwd /mTU, the difference between the volume-averaged fuel temperature and the coolant temperature is still only 80 percent of that at the beginning of life.
Therefore, the worst time in life to perform a LOCA evaluation is at the beginning of life.
The time actually selected for the TACO 3 inputs for this report was 1 mwd /MTu.
The studies for burnup in the referenced evaluation model report have been shown to support a basis for the selection of the worst time in life to evaluate a LOCA given that a blowdown rupture cannot occur.
These same studies, when compared to TACO 3 fuel temperatures, assure that a blowdown rupture will not occur for the McGuire and Catawba analyses.
The evaluation model studies determine that the beginning of life is the most conservative time to assume for the LOCA analysis, and this conclusion applies for both the codes and methods of Reference 1
and the current evaluation model.
Thus, the time in life for the McGuire and Catawba LOCA evaluations and calculations is established as beginning-of-lifa.
t
- 6.10 -
l 1
y 1
1 e
i I
I M
Break Location In general, break spectrum studies are plant-specific in terms of applicability.
The break location studies, however, which have
-typically been included among the break spectrum cases can be j
considered generic.
There are three break locations to consider for the large break LOCA:
the hot leg piping, the cold leg piping at the pump suction, and the cold leg piping at the pump discharge-i 1
-between the pump and the reactor vessel.
Both the hot leg break and the pump suction break have been consistently shown to result in peak cladding temperatures below those predicted for pump discharge breaks.
(This is demonstrated in the Westinghouse RESAR for the 3411 Mwt class and is consistent with the numerous analyses reported by B&W for B&W-designed PWRs.)
The analysis in Appendix A of Reference 1 concluded that the blowdown and reflood cooling for the pump suction break were sufficient to dismiss the suction break as a potential limiting case strictly on the basis of the' blowdown and reflood system analyses.
iC An examination of the consequences of the break location explains these observations.
With the break at the pump discharge, one accumulator and approximately 25 percent of the available pumped injection will be bypassed directly to the containment, not providing blowdown cooling and lengthening adiabatic heatup.
Flow through the core during blowdown for the hot leg and the pump suction breaks will be more positive than for the pump discharge break because of the leak position in the loops.
For the hot leg break, there can be no bypass of the injected ECC water unless that water has already passed through the core.
Furthermore, the elevation head driving reflood can rise to as high as the spillover elevation in the steam generator tubes.
For the pump suction break, the core flooding also improves because the break is moved to the other side of the pump.
The effect,
- 6.11 -
L 1
l however, is not as strong as for a hot leg break.
Therefore, in conjunction with the better blowdown cooling, the suction break produces a Jover peak cladding temperature than does the pump 1
discharge break.
While the lower peak temperature alone precludes the suction break from being limiting, there is further reason to I
concentrate analyses upon breaks in the pump discharge piping.
Because of the volume of pumped ECCS flow spilled from the system for the pump discharge break (roughly 25 percent), that condition represents a more limiting ECCS flow than does the pump suction I
break.
For these reasons, and based upon the discussions of the foregoing paragraphs, the spectrum and LOCA limits cases performed for Catawba and McGuire are for large breaks in the reactor coolant pump discharge piping, q
O
- 6.12 -
.l
g i
i
- \\,
7i FIGURE 61 CORE MASS FLUX j
a, Es_,/.
EM SPECTRUM STUDY VERSUS MoGUIRE/ CATAWBA BASE MODEL
]
3,o i
1 120 -
0 1
30 -
2NG AT PD WITH Cd = 1.0 40 -
,' \\
1 p
,e w
i 1
ge-- -
e
-, s,
'..,,,,,,,,., ~ ' '..,'
g e
E
-s 80 -
7 120 -
McGUIRE/ CATAWBA BASE MODEL EM SPECTRUM STUDY 160 i
0 4
8 12 16 20 24 28 TIME, S t,
A FIGURE 6 2' TACO 3 FUEL TEMPERATURE AS A FUNCTION OF BURNUP i
l:
HIGH LOCAL POWER l'
W l
1 E
lD AVERAGE LOCAL POWER i.
i r
0 20000 40000 60000
(
ROD AVERAGE BURNUP (mwd /mTU)
- 6.13 -
l 7.
Plant-sDecific Studies and Snectrum Analysis p)
(v Although a
considerable portion of the analysis inputs and assumptions can be set by the evaluation model and its sensitivity studies, some parameters are dependent on plant-specific inputs and can only be established by individual plant studies.
These studies and the spectrum analyses are performed to identify a worst case break to use in calculating the LOCA Limits.
Figure.7-1 shows the l'
order in which the plant-specific sensitivity studies and ths spectrum analysis cases were performed.
The calculations were all done with the changes to the BEACH spacer grid and rupture models (the first step in the modifications) described in Chapter 5.
This chapter presents the results of the studies and describes the applicability of the results to analyses done with the final i
evaluation model, which incorporates all of the modifications.
L.1 Base Case The first step in performing a series of sensitivity studies is to establish a base case.
For the studies presented in this chapter, the base case is a double-ended guillotine cold leg break, with a discharge coefficient of 1.0, located between the reactor coolant pump and the reactor vessel.
Figures 7-2 through 7-9 present key l
parameters for this case.
The results compare well to previous cases analyzed with the original evaluation model, done for a standard 3411 Mwt plant with dry containment.
The major differences in the results are in the length of blowdown, which is approximately four seconds longer, owing to the lower pressure of l
the ice condenser containment in this base case.
Core flow is l
affected by the lower initial fluid temperatures in the upper head l
region, as shown by negative flows greater than those calculated for the design assumed in the evaluation model case.
Core reflooding is slower because of the additional steam binding l
encountered for the lower containment backpressure.
This lower l
' O V
- 7.1 -
l l
i
i r
L i
flooding rata. is somewhat alleviated by the venting paths afforded by the upper head spray nozzles, also not a feature of the configuration assumed for the original EM studies.
4 L
M Accumulator,gonfiauration l
As was discussed in Chapter 4, one difference between the McGuire and the Catawba plants is the design of the accumulators and the r
L
' accumulator surge lines.
The McGuire accumulator liquid volumes L.
- are lower than those of Catawba, and the pressurizing gas volumes i
are higher.
- Also, the surge line resistances for McGuire are higher than for Catawba.
The table below contrasts parameters for the plants.
[
1 l
Parameter McGuire Catawba 3
l Accumulator Liquid Volume, ft 950 1050 3
l Accumulator Gas Volume, ft 450 350 s
Accumulator Tank Pressure, psia 615 615 l
Supply Line Flow Area, ft 0.42 0.42 2
Supply L3ne Resistance Factor 12.4 5.7 l
Since the downcomer is full, or nearly so, before the accumulators empty, the different liquid volumes are unlikely to affect the LOCA results.
The higher gas volume, which will maintain a higher tank pressure during discharge, and the higher line resistances for
- McGuire, would tend to be offsetting.
Therefore, it is not surprising that the study showed essentially no difference in peak cladding temperatura.
Selected parameters for these runs are shown in Figures 7-10 and 7-11.
The analysis included the BEACH grid and rupture blockage modifications but only altered input parameters for RELAPS/ MOD 2-B&W and REFLOD3B to form the comparison.
Neither of these codes is affected by the evaluation model changes, and the results are consistent over the length of the pin (see Figure 7-11).
'I hus,
the study remains valid for use with the final evaluation model.
7.2 -
t 1
.fQ The results do not show a significant difference in effects between the two accumulator configurations.
The Catawba configuration, J
because it is 4 F highor in peak cladding temperature, is used for the remaining studies in this chapter and for the LOCA limits studies.
- L_3, Break Snectrum Analysis The break. spectrum analysis is performed to determine the worst case break size, the worst case break configuration, and the. worst case break location.
The break location study was discussed in Section 6.3 with the conclusion that the hot leg and reactor coolant pump suction breaks are assured to be substantially less limiting than the pump discharge breaks because of the lack of ECCS spillage for these events.
The differences between the split and l-the guillotine break and the range of break sizes cannot be j
generalized,
- however, and those studies were run for the l
McGuire/ Catawba evaluation.
The break size study was performed first and followed by the break type study.
1 For the BWFC large break evaluation model, the break size study is interchangeable with a discharge coefficient study since the break l
flow is directly proportional to the product of the break area and j
the discharge coefficient.
For the McGuire/ Catawba model, the discharge coefficient study was conducted for a guillotine break of twice the area of the cold leg piping located at the pump discharge with discharge coefficients of 1.0, 0.8, and 0.6.
Key parameters for these cases are shown grouped by case in Figures 7-12 through 7-35.
Table 7-1 presents a comparison of the timing of events for the three cases.
1 L
There are no major differences among the sequences of events for the three cases that make up the discharge coefficient study; the peak cladding temperatures differ by only 110 F.
As expected, q
- 7.3 -
a
.,a,
,n-
.n..
i blowdown is extended as the break flow is decreased.
The 0.8 and 0.6 discharge coefficient cases remove slightly more heat from the fuel during
- blowdown, as indicated by the centerline fuel temperatures at the end of blowdown: 1067 F, 1013 F, and 1019 F for the cd =
- 1. 0, 0.8, and n.6 cases, respectively.
All three cases leave slightly over 50 cubic feet of liquid in the reactor vessel at the end of blowdown, making the adiabatic heatup periods during lower head refilling nearly the same.
The reflooding transients are almost identical for the three cases, except that the 0.6 case floods slightly faster from 50 to 90 seconds, and thereafter, floods at the same rate as the other two cases, having built up a higher core liquid level.
The difference in peak cladding temperatures between the cd = 1.0
= 0.8 cases is due primarily to the better blowdown and ~ the Cd cooling in the latter case.
The difference in centerline fuel temperatures at the end of blowdown, 54 F, is preserved through the remainder of the calculations, creating a 50 F difference in cladding temperature at the time of the peak.
The Cd = 0.6 case experiences the better blowdown cooling, being 48 F cooler than the Cd = 1.0 case at the end of blowdown, but it also experiences extra
]
flooding between 50 and 90 seconds.
As can be seen in the cladding
. temperature plots, all three of the cases experience an increase in
- cooling, caused by the occurrence of rupture in the adjacent upstream node, at about 80 or 90 seconds.
Because the Cd = 1.0 and 0.8 cases have very similar flooding histories, this the cd
=
effect is nearly the same for both.
For the Cd = 0.6 case, the higher flooding rates near the rupture time cause the effect to be more pronounced.
The resultant temperature difference, 50 to 60 F, remains throughout the rest of the evaluation, placing the Cd = 0.6 0.8 case at the time of peak case 60 F cooler than the cd
=
cladding temperature and 110 F cooler than the cd = 1.0 case.
The peak cladding temperatures were all for nodes that were adjacent to the rupture location.
These evaluations were conducted
- 7.4
6 with only the BEACH spacer grid and rupture modifications described
()
'in Chapter 5; thus, the final cladding temperature results were calculated in FRAP-T6.
As such, there was no axial conduction allowed along the fuel pin.
With axial conduction, the peak cladding temperature location may move to the grid span above the location of peak power.
To be sure that the conclusions of the study. remain valid for the complete set of evaluation model changes, a comparison of the cladding temperatures in the grid span j
above - the ruptured span is required.
The table below shows the.
peak cladding temperatures for the cladding in the non-grid nodes one span above the rupture location.
NODE 14 NODE 15 CASE PCT, F / Time, s PCT, F / Time, s Cd = 1.0 1683 / 321 1770 / 321 l
cd = 0.8 1668 / 281 1761 / 378 Cd = 0.6 1676 / 352 1752 / 378 O}
\\'
As can be seen, the differences in peak temperatures are compressed at the higher elevation and there is little difference in the results across the spectrum.
From this study, it would appear to matter little which case was used for the remaining evaluations.
Still, the cd = 1.0 is slightly worse than the others and it is more than only marginally worse at the lower elevations.
Therefore, the cd =1.0 was selected for the remaining sensitivity studies and as the worst case for the LOCA limits calculations with the final evaluation model.
LA Break Tvoe Following the selection of the discharge coefficient, the type of break, split or guillotine, was considered.
The guillotine break is modeled as a complete severance of the pipe, allowing separate discharges through the full area of the cold leg piping from both
,F'N V
- 7.5 -
i the reactor vessel and pump sides of the break location.
No mixing of the flows from the two sides of the break is allowed.
The split
~
~ break assumes discharge from the cold leg piping through an area twice the1 size <tf the cold leg piping cross section.
Although the flows from the two sides, RC pump and reactor vessel, must still pass through limiting pipe areas, they are allowed to mix at the i
break location.
The blowdown rates and system flow splits are i
somewhat different for the two types of breaks, and that can lead to some differences in cooling response.
Figures 7-36. through 7-43 present the results of the split type break case.
These should be compared to Figures 7-12 through 7-19 which are for the reference guillotine case.
Both are double-area r
breaks with Cd = 1.0 and located at the pump discharge.
Key data from the split case are included with the spectrum studies in Table 7 - 1. - As can be seen, the period of blowdown for the split break is shorter by two seconds.
Although the blowdown flows differ somewhat, the degree of cooling is essentially the same between the split and the guillotine, as shown by the centerline fuel temperature of the hot pin at the end of blowdown, 1067 F for the
. guillotine and 1061 F for the split.
Similar to the Cd 0.6
=
guillotine break from the break size study, the flooding rate for the split is higher than for the guillotine between 70 and 100 seconds.
Therefore, when rupture occurs, the interphase heat transfer increases at the ruptured node are more effective with the result that the cladding temperature increase is slowed substantially more for the cplit than for the guillotine.
The temperature difference thus set up is preserved for the remainder of the run, resulting in the 59 F lower peak cladding temperature for the split case.
An examination of the results for the grid span just above the ruptured location verifies that the split should remain the less limiting case when the complete set of evaluation model techniques is applied.
n
)
'l f:
i
~
") :'
NODE 14 NODE 15 x
phEE PCT. F / Time, s PCT, F / Time, s i
f Guillotine 1683 / 321 1770 / 321 Split 1665 / 334 1750 / 365 l
l As with the discharge coefficient study, the differeraces in results are almost marginal.
Nonetheless, the double-ended break produces the higher peak temperatures, and on that basis, was selected as
)
the worst configuration for the remaining sensitivity studies and L
for the LOCA limits calculations using the complete set of EM changes.
)
M Maximum ECCS Analysis The final sensitivity study, prior to the LOCA limits studies, is to determine which condition for the ECCS is more severe, with or without a single failure.
Under a single failure assumption, only g,
(
one train of pumped ECCS injection is available.
With no failure, two full trains are available.
Because the sizing of each individual train must be sufficient to mitigate an accident, the second train is redundant relative to providing adequate water for core cooling.
In an analysis, nearly all the extra injection capacity will spill from the primary coolant system to the containment where it may interact with the atmosphere to reduce the containment pressure.
As the lowering of the containment pressure causes a reduction in core reflooding rate, the evaluation of a fully functional ECCS may actually show a higher peak cladding temperature than would be predicted using the single failure assumption.
The calculations presented thus far have all been performed assuming a single active failure.
The assumed failure, that of the diesel generator emergency power supply, results in a loss of
.g
(
- 7.7 -
b~
.h &
L i
L electrical power to half of the pumped ECC systems.
This is generally referred to as a " minimum ECCS" case.
To evaluate the c
assumption of no failure, a calculation was made with all of the SCCS functional, the " maximum ECCS" case.
Figures 7-44 through 7-59 present relevant parameters for the maximum ECCS case.
These should be compared to the guillotine Cd = 1.0 from the discharge
]
coefficient study, Figures 7-12 through 7-19.
1 i
The extra ECC available in the maximum ECCS case has two effects f
during reflood:
First, the ECCS water that is injected into the intact loops condenses more of the steam flowing through the loops and lowers the RCS pressure.
- Second, the ECCS water that is spilled into the containment mixes with the containment atmosphere and reduceis its pressure.
The reduced containment pressure in turn reduces the RCS pressure.
The effect of reducing the RCS pressure is to increase the specific volume of the steam created during core reflooding.
As a result, the steam is more difficult to vent through the system, and the core flooding rate decreases.
A comparison of Figures 7-50 and 7-14 shows the flooding rate reduction for the McGuire/ Catawba model.
As there are no blowdown cooling benefits to compensate for the lower flooding rate, the cladding temperature increases by 82 F as shown in Figures 7-54 and 7-16.
An examination of the cladding temperatures one grid span above the rupture location shows that the cladding temperature is increased at.that location as well.
Therefore, the conclusion that the peak cladding temperature will be higher for the maximum ECCS case will also apply for the analyses done with the final evaluation model.
- 7.8 -
u 1
)
O. '
l 3
ar x Table 7-1 Spectrum'and Break Type Comparison t
i o
Guillotines Split Item or Parameter Cd =>
1.0 0.8 0.6 1.0 End of Blowdown, a 22.3 23.6 29.0 20.3 Liquid in: Reactor Vessel 3
at'EOB, ft 57.4 72.3 69.1 64.8-
]
Bottom-of-Core Recovery, s 35.2 36.2 42.6 32.7 j
Time of. Rupture, s 74.4 78.6 90.4 73.5
]
Ruptured Node
- 11 11 11 11 g
PCT at Rupture Node, F 1613 1597 1597 1608 Node Adjacent to Rupture
- 12 12 12 12 PCT of Adjacent Node, F 1862 1812 1750 1803 j
Node in Adjacent Grid Span 15 15 15 15 PCT of Adjacent Grid Span, F 1770 1761 1752 1750
[
~ Pin PCT Node
- 12 12 15 12
)
- Refer to Figure 4-4 for noding arrangement s
f'\\_/
1
'l L
o l
l:
l:
L 1,
l
- 7.9 -
..,,.. - -,..... - ~
.,......... - - - - -... -. - - -. - ~. -. -.. ~ ~.. - -. - -, - - - - - -.. - -. -
FIGURE' 7-1 PLANT SPECIFIC STUDIES ANALYSIS -DIAGRAM i
i ACCUMLAATOR I
CONFIGURATION SPECTRUM STUDY ECC STUDY STUDY DISCHARGE BREAK COEFFICIENT TYPE l
i McGUIRE SPUT PCT = 1858 F PCT = 1803 F i,
?
BASE MODEL Cd = 1.0 MINIMUM l
i PCT = 1858 F PCT = 1862 F PCT = 1862 F l
CATAWBA Cd = 0.8 GUILLOTINE PCT = 1862 F PCT = 1812 F PCT = 1862 F Cd = 0.6 MAXIMUM LOCA PCT =2 1752 F PCT = 1944 F UMITS i
BASE MODEL: 2NG @ PD Cd = 1.0 9
O O
~
s.
. e Ei-
'! t,,
i f'(fy FIGURE 7 2 SENSITIVITY STUDY - BASE MODEL
{
SYSTEM PRESSURE DURING BLOWDOWN j
2400 2000 -
p i
1000 -
2NG AT PD WITH Cd = 1.0 i
t n!
1200 -
[
i 800 -
~
400 -
I 0
0 4
8 12 16 20 24 28 TIME, S
~
Qf FIGURE 7 3 SENSITIVITY STUDY - BASE MODEL i
MASS FLUX DURING BLOWDOWN AT PEAK POWER LOCATION-160 120 -
i 80 -
2NG AT PD WITH Cd = 1.0 40 -
0 l
40 -
80 -
120 -
I 160 0
4 8
12 19 20 24 28 TIME, S
- 7.11 -
p w
~
y.
, - -.,. -,-.,-i,,,---
aw e-w,
..,,,-_w a w
., +.,
v.
V I
FIGURE 7-4 SENSITIVITY STUDY - BASE MODEL REFLOODING RATE 10 d
j
.a
~
l 8-h.
0~
g j
2NG AT PD WITH Cd = 1.0
]
j 4-2-
\\.
0 4
i 0
200 400 600 t
TIME, S
. 9:
y FIGURE 7-5 SENSITIVITY STUDY - BASE MODEL HEAT TRANSFER COEFFICIENT AT PEAK POWER' LOCATION 1000 -
500f
(
t I
100 - l l2 2NG AT PD WITH Cd = 1.0 j.
'g
- go _~
NODE 11 l
o l'
~
l r
~
10
=;.:
'lly % f.'- g 5 - 0 200 400 600 TIME, S \\ - 7.12 - [ t.
i ' t 4 [;.,... L.. ;', m. FIGURE 7-6 SENSITIVITY STUDY - BASE MODEL i
- A._)
PEAK CLADDING TEMPERATURE p i* 2000 -^ r 1000 - E' l i l,,., 1200 - 5 2NG AT PD WITH Cd = 1.0 PCT LOCATION-NODE 12 l l p-l 400 0 i 0 200 400 000 TIME, S .f Lt i \\._, FIGURE 7 7 SENSITIVITY STUDY - BASE MODEL CLADDING TEMPERATURE AT RUPTURE LOCATION I 5 2000 - 4 1900 - w l. 1200 - l-1 s j 800 - 2NG AT PD WITH Cd = 1.0 RUPTURE LOCATION-NODE 11 [. 400.. in r 0 't g 4 0 200 400 600 TIME, S - 7.13 - D
pm j e t t 6 "<x ? FIGURE 7 8 SENSITMTY STUDY BASE MODEL CLADDING TEMPERATURE IN ADJACENT GRID. SPAN I i i f 2000 - 1000 - I ~ 1200 - j 2NG AT PD WITH Cd = 1.0 NODE 15 e 400 - f 0 0 200 400 000 TIME, S O FIGURE 7-9a SENSITMTY STUDY BASE MODEL FLUID TEMPERATURE AT PCT LOCATION 2000 - t 1800 - . u. 2NG AT PD WITH Cd = 1.0 1200 - ~ PCT LOCATION-NODE 12 ~ [ .M 400 - 0 4 8 12 16 20 24 28 TIME, S - 7.14 - ~
i i i r' FIGURE 7 9b SENSITIVITY STUDY CASE MODEL i ~ ()s FLUID TEMPERATURE AT PCT LOC ATION I 2000 - t 2A'O tT PD WITH Cd = 1.0 PCT LOCATION-NODE 12 i 1000 - ( l k i 1200 -
- ; W
, % %,3 4 400 - ~ i 0 0 200 400 000 t TIME, 8 FIGURE 710 ACCUMULATOR STUDY McGUIRE vs CATAWBA ACCUMULATOR FLOW RATES 2A/G AT PD WITH Cd = 1.0 [ McGUIRE CONFIGURATION i 6000 - ,/ CATAWBA CONFIGURATION 's, l l 4000 - 7 b i i t ( 0 20 40 00 80 100 120 TIME, S - 7.15 - I
i i FIGURE 711 ACCUMUL.ATOR STUDY MoGUIRE vs CATAWBA g PEAK CLADDING TEMPERATURES W { 3000 - [ [ S ~ f I1200-2NG AT PD WITH Cd = 1.0 i McGUIRE CONFIGURATION-NODE 12 [ CATAWBA CONFIGURATION-NODE 12 000 - ~*"*"" i [ 400 - i i o l 0 200 400 e00 [ Time, s O FIGURE 712 DISCHARGE COEFFICIENT STUDY Cd = 1.0 i SYSTEM PRESSURE DURING BLOWDOWN 2400 l f 2000 -- L [ 1000 - i 2AfG AT PD WITH Cd = 1.0 l im - r .x - r 400 - [ ~ g o s. 20 2. 2. TIME, S - 7.16 - r
yQ i i.; N FIGURE 713 DISCHARGE COEFFICIENT STUDY Cd = 1.0 (), MASS FLUX DURING BLOWDOWN AT PEAK POWER LOCATION ~-. i 120 - i 80 - ING AT PD WITH Cd = 1.0 g 0 [ r a. 120 - 100 0 4 8 12 16 20 24 28 TIME, S FIGURE 714 DISCHARGE COEFFICIENT STUDY Cd = 1.0 REFLOODING RATE 10 8-g 2A/O AT PD WITH Cd = 1.0 p 1! 2~ l gt l C 1 l 0 ( 0 200 400 600 l TIME, S b
[ l1 FIGURE 715 DISCHARGE COEFFlCIENT STUDY Cd = 1.0 HEAT TRANSFER COEFFICIENT AT PEAK POWER LOCATION 1000 - I-i 100 -- SNQ AT PD WITH Cd = 1.0 go. NODE 11 1 10 - }% n % v-m W Wc _An_ 8~ 0 = = a Tiut, s O FIGURE 716 DISCHARGE COEFFICIENT STUDY Cd = 1.0 PEAK CLADDING TEMPERATURE 2000 - 1000 - 1200 - 2A/G AT PD WITH Cd = 1.0 PCT LOCATION-NODE 12 TIME, S - 7.18 - l
i I i FIGURE 717 DISCHARGE COEFFICIENT STUDY Cd = 1.0 ( s J Cl. ADDING TEMPERATURE AT RUPTURE LOCATION l i 2000 - + j f 1000 - L I i ~ h 1200 - i i W I g 800 - 2NG AT PD WITH Cd = 1.0 o RUFTURE LOCATION-NOD $ 11 l 400 - l i t' 0 { O m a e TIME, S (/ FIGURE 718 DISCHARGE COEFFICIENT STUDY Cd = 1.0 l CLADDING TEMPERATURE IN ADJACENT GRID SPAN i i 2000 - 1 r t 1000 - l l 1200 - h 1 2NG AT PD WITH Cd = 1.0 NODE 15 g 800 - u l 400 - l o. l ( 0 200 400 600 tME,S - 7.19 - D
o FIGURE 719a DISCHARGE COEFFICIENT STUDY Cd = 1.0 l FLUID TEMPERATURE AT PCT LOCATION sooo - 1e00 - ~ 2NQ AT PD WITH Cd = 1.0 [ 1200 - PCT LOCATION-NODE 12 i j 800 - i .d 400 - 0 0 4 8 12 16 20 24 28 TIME, D O FIGURE 7-19b DISCHARGE COEFFICIENT STUDY Cd = 1.0 FLUID TEMPERATURE AT PCT LOCATION 2000 - ~ 2NG AT PD WITH Cd ' 1.0 PCT LOCATION-NODE 12 1600 - a 1200 - i kh l ll 1 l 400 - I h 0 200 400 soo TIME, S - 7.20 -
7 3 FIGURE 7-20 DISCHARGE COEFFICIENT STUDY Cd = 0.8 ! () SYSTEM PRESSURE DURING BLOWDOWN 2400 2000 - I# ~ 2NQ AT PD WITH Ca = 0.8 1200 - 800 - 400 - I o J i o 4 6 12 16 20 24 28 TIME, S (- . 6, - (_ FIGURE 7 21 DISCHARGE COEFFICIENT STUDY Cd = 0.8 MASS FLUX DURING BLOWDOWN AT PEAK POWER LOCATION 100 ( ~ 120 - 2A/G AT PD WITH Cd = 0.8 40 - o-- i 40 - 120 - 1 / o 4 8 12 16 20 24 28 i. TIME. S l - 7.21 -
t r e FIGURE 7 22 DISCHARGE COEFFICIENT STUDY Cd = 0.8 g REFLOODING RATE w e 10 i p. y 0- [ l i L I t c.. 2A/G AT PO WITH Cd = 0.8 . e i 0 0 200 400 600 TIME, S 9: FIGURE 7 23 DISCHARGE COEFFICIENT STUDY Cd = 0.8 HEAT TRANSFER COEFFICIENT AT PEAK POWER LOCATION i 1000 - 100 ; 2NG AT PD WITH Cd = 0.8 So _~. NODE 11 I, 10 5-TIME, S j - 7.22 -
s l i 'r FIGURE 7 24 DISCHARGE COEFFICIENT STUDY Cd = 0.8 ) x j x _,, PEAK CLADDING TEMPERATURE i 2000 - r'.. - I N i 1600 - (f I1200-2A/O AT PD WITH Cd = 0.8 [ PCT LOCATION-NODE 12 i 400 - 0 0 200 400 000 l TIME, S b ts FIGURE 7 25 DISCHARGE COEFFICIENT STUDY Cd = 0.8 i CLADDING TEMPERATURE AT RUPTURE LOCATION 2000 - 1600 - w }f 120C, - i W 800 - 2NO AT PD WITH Cd = 0.8 RVPTURE LOCATION-NODE 11 c 400 - t i r. 0 ( 0 200 400 600 TIME, S - 7.23 -
i i I B FIGURE 7 26 DISCHARGE COEFFICIENT STUDY C. ~. 0.8 CLADDING TEMPERATURE IN ADJACENT GRID SPAN i I 2000 - ,000 - u. i ,200 - ) I 2A/O AT PD WITH Cd = 0.8 NODE,5 4oo. i i e i 0 l 0 200 400 000 i TIME, S O! FIGURE 7 27a DISCHARGE COEFFICIENT STUDY Cd = 0.B l FLUID TEMPERATURE AT PCT LOCATION l 2000 - ~ t f ~ WG AT PD WITH Cd = 0.8 g,_ Pc1 _ _ ~..,2 1 ~ G TIME, S - 7.24 - l
3 i e i l FIGURE 7 27b DISCHARGE COEFFICIENT STUDY Cd = 0.8 i 1 I () FLUID TEMPERATURE AT PCT LOCATION 2000 - I 2AtG AT PD WITH Cd = 0.8 PCT LOCATION-NODE 12 gg _ f Sam - n usaggg % l-400 - 0 O m a m l Tue, s FIGURE 7 28 DISCHARGE COEFFICIENT STUDY Cd = 0.S SYSTEM PRESSURE DURING BLOWDOWN i 2400 t 2000 - l ~ l sex - c f 2AtG AT PD WITH Cd = 0.6 g ,2w.- .w - 400 - TIME. S - 7.25 -
\\ a i FIGURE 7 29 DISCHARGE COEFFICIENT STUDY Cd = 0.6 [ MASS FLUX DURING BLOWDOWN AT PEAK POWER LOCATION }. 100 120 - i f 80 - 2NG AT PD WITH Cd a 0.6 0-l i l to - L 1 120 - 100 0 4 6 12 16 20 24 20 TIME, S O FIGURE 7 30 DISCHARGE COEFFICIENT STUDY Cd = 0.6 REFLOODING RATE 10 8-6- g 2NG AT PD WITH Cd = 0.6 g 4 2- 'Y ~ TIME, S - 7.26 -
t FIGURE 7 31 DISCHARGE COEFFICIENT STUDY Cd = 0.6 c 3 O HEAT TRANSFER COEFFICIENT AT PEAK POWER LOCATION { 1000 t S00 2 i ~ t i l l 500, l 2NG AT PD WITH Cd = 0.6 i go. NODE 11 l I- ~ I h gNNdyww%M 50 si i 0 E e TIME, S I FIGURE 7 32 DISCHARGE COEFFICIENT STUDY Cd = 0.6 j PEAK CLADDING TEMPERATURE i 2000 - 1200 2A/G AT PD WITH Cd = 0.6 NODE 12 t i 400 - I t 0 200 400 600 TIME, S l - 7.27 -
i i I >i F!GURE 7 33 DISCHARGE COEFFICIENT STUDY Cd = 0.6 I CLADDING TEMPERATURE AT RUPTURE LOCATION l 1 2000 - l 1000 - w i 1200 - i f l i i g 800 - ING AT PD WITH Cd = 0.6 O RUPTURE LOCATION-NODE 11 f 400 -- o t o 200 400 000 TIME, S 9: FIGURE 7-34 DISCHARGE COEFFICIENT STUDY Cd = 0.6 CLADDING TEMPERATURE AT ADJACENT GRID SPAN f 2000 - 1000 - E ~ i 1200 - l' 2A/G AT PD WITH Cd = 0.6 [ PCT LOCAT'ON-NODE 15 l 400 - i 0 200 400 600 l TIME, S - 7.28 -
l i I ( 3 FIGURE 7 35a DISCHARGE COEFFICIENT STUDY Cd = 0.6 j i j
- _)
FLUID TEMPERATURE AT PCT LOCATION a [ f-2000 - ~ 1000 - j w ~ l 2A/G AT PD WITH Cd = 0.6 1200'- PCT LOCATION-NODE 12 4,,.. O i [ 0 4 8 12 16 20 24 28 i TIME, S FIGURE 7 35b DISCHARGE COEFFICIENT STUDY Cd = 0.6 FLUID TEMPERATURE AT PCT LOCATION l 1 2000 - l ~ 2NG AT PD WITH Cd = 0.6 PCT LOCATION-NODE 12 $g _ l,,00 $j % h ) $ % g i I-400 - t 0 200 400 600 TIME, S -7.29 - ,,w--
- f
,tf ~ ~ ~ ^ ~ ^ l U l l FIGURE 7 36 BREAK TYPE STUDY SPLIT, Cd = 1.0 SYSTEM PRESSURE DURING BLOWX)WN i 2400 i i j 2000 - } ~ l 1803 - I 2A/S AT PD WITH Od = 1.0 E 1200 - ~ l 800 - I t 400 - I E r 0 0 4 8 12 16 20 24 28 i TIME, S O-FIGURE 7 37 BREAK TYPE STUDY SPLIT, Cd = 1.0 MASS FLUX DURING BLOWDOWN AT PEAK POWER LOCATION j 160 l 120 -- 60 - 2NS AT PD WITH Cd = 1.0 40 - d 0 v 2 i f 30 - 120 - -100 0 4 8 12 16 20 24 28 TIME. S - 7. 30 -
F pc FIGURE 7 38 BREAK TYPE STUDY SPLIT, Cd = 1.0 I_)'- REFLOODING RATE 10 J e-I 6-N 2A/S AT ?O WITH Cd = 1.o ) E d 1 l 2-k ) ~ i o o 200 400 soo TME, S (-. ( FIGURE 7 39 BREAK TYPE STUDY SPLIT, Cd = 1.0 HEAT TRANSFER COEFFICIENT AT PEAK POWER LOCATION 1000 _ l500.. l too _ l 2A/S AT PD WITH Cd = 1.0 $o NODE 11 I. I 10 4 %%WA 2^A L 5-l t o 200 400 soo l TIME, S j- - 7.31 -
i I FIGURE 7 40 BREAK TYPE STUDY - SPUT, Cd = 1.0 l PEAK CLADDING TEMPERATURE i 2000 - l 1000 - a _ I ( 1200 - l 2NS AT PD WITH Cd = 1.0 l t PCT LOCATION-NODE 12 ~ f 400 - i 0 0 200 400 000 f TIME, S O; FIGURE 7-41 BREAK TYPE STUDY - SPLIT, Cd = 1.0 CLADDING TEMPERATURE AT RUPTURE LOCATION i 20m - i I 1000 - I Y ~ 1200 - l 800 - 2A/S AT P3 WITH CC = 1.0 j RUPTURE LC#ATION-NODE 11 [ 400 - l l TME, S l - 7.32 - 1
r __ _ l 4 l FIGURE 7-42 BREAK TYPE STUDY SPUT. Cd = 1.0 1 (,) CLADDING TEMPERATURE IN ADJACENT GRID SPAN 2000 - l i t 1 1000 - u. i F 1200 - 1 f-2A/S AT PD WITH Cd = 1.0 i NODE 15 t - 400 - I 0 o a e m l t TIME, s ,.{v FIGURE 7-43a BREAK TYoE STUDY - SPLIT, Cd = 1.0 FLUID TEMPE9AT1.lRE AT PCT LOCATION j i 2000 - 1600 - ) g f i g 2NS AT PD WITH Cd = 1.0 I 1200 - PCT LOCATION-NODE 12 h l s 800 - [ \\: 400 - 1 i. f' O 0 4 8 12 16 20 24 28 ( TIME, S l. l 1 - 7.33 - ? +
i FIGURE 7-43b BREAK TYPE STUDY SPLIT, Cd = 1.0 FLUID TEMPERATURE AT PCT LOCATION i 2000 - l l SArs AT PD WITH Cd = 1.0 PCT LOCATION-NODE 12 1000 - ~ l NYNVjh h 4gg '~ 400 - i o 0 200 400 600 TIME S FIGURE 7-44 MAXIMUM ECCS STUDY. DECLB, Cd = 1.0 3 MINIMUM ECC PUMPED INJECTION CONTAINMENT PRESSURE to 8- \\ l e-f, ~ D-TIME, S - 7.34 -
i l l l FIGURE 7-45 MAXlMUM ECCS STUDY DECLB, Cd = 1.0 f, ,i / SYSTEM PRESSURE DURING BLOWDOWN 2400 l 2000 - ) \\ 1900 - g ~ 2NQ AT PD WITH Cd = 1.0 1 i 1200 - i 1 800 - t ) 400 - l l 1 0 i 0 4 8 12 16 20 24 28 TIME, S !7_s) i V i FIGURE 7-46 MAXIMUM ECCS STUDY DECLB, Cd = 1.0 MASS FLUX DURING BLOWDOWN AT PEAK POWER LOCATION i i 100 120 - ? 30 I 2NO AT PD WITH Cd = 1.0 40 - l 0- + g ~ f i 120 - ( m 160 0 4 8 12 14 20 24 20 TIME. S - 7.35 -
i l t-1 FIGURE 7-47 MAXIMUM ECCS STUDY DECLB, Cd = 1.0 i [ MAXIMUM ECC PUMPED INJECTION CONTAINMENT PRESSURE ] 10 i- ) s-s i j 6-1 i l 4- ) l \\ 2-i ~ o ' o 200 400 soo l 9 TIME, 5 FIGURE 7-48 MAXIMUM ECCS STUDY DECLB, Cd = 1.0 l PUMPED ECC INJECTION FLOW RATE 1000 l .00 : E l l /:,........................................................................................ 400 - h l MAX! MUM ECC PUMPED INJECTION [: MINIMUM ECC PUMPED INJECTION 200 - l ? / I '.*IME, S - 7.36 -
t i ^ FIGURE 7-49 MAXIMUM ECCS STUDY DECLB, Cd = 1.0 V DOWNCOMER WATER LEVEL so j f 18 - 1e - t ] 14 - 12 - h l 10 - 4 ? MAXIMUM ECC PUMPED INJECTION 6-MINIMUM ECC PUMPED INJECTION 4-i 2-0 ( o m a m t TIME, S FIGURE 7 50 MAXIMUM ECCS STUDY DECLB, Cd = 1.0 i REFLOODING RATE j 10 I 8-I 6-g 2NG AT PD WITH Cd = 1.0 j g 4 g = l 0 200 400 soo l TIME S l I l - 7.37 - l
3r l l FIGURE 7 51 MAXIMUM ECCS STUDY DECLB, Cd = 1.0 f HEAT TRANSFER COEFFICIENT AT PCT LOCATION l 1000 -- i s00 i i t I100. 2NG AT PD WITH Cd = 1.0 'h! { gZ NODE 12 i l. j 10 -
- =..t e l
1 5-O N TIME, S FIGURE 7 52 MAXIMUM ECCS STUDY - DECLB, Cd = 1.0 HEAT TRANSFER COEFFICIENT AT RUPTURE LOCATION 1000 _ I g-- 1 I t I N l Soo -.i 2NG AT PD WITH Cd = 1.0 gi NODE 11 I. 10 g ( pW '%y4yv..... ep.,. . g,wp L G, 0 200 400 W TIME, S - 7.38 - .i.
p 1 1 1 s / FIGURE 7 53 MAXIMUM ECCS STUDY DECLB, Cd = 1.0 ()3 HEAT TRANSFER COEFFICIENT IN ADJACENT GRID SPAN i 1000 m i n ~ l i 100 - 2A/G AT PD WITH Cd = 1.0 so NODE 15 ) l MvmgW:.... ^t..t.vid;. .. :# avg l J s- ~ o a e e TIME, S U(N i FIGURE 7 54 MAXIMUM ECCS STUDY - DECLB, Cd = 1.0 PEAK CLADDING TEMPERATURE 2000 - --g._ t 1.= - u. ~ 1200 - 2A/G AT PD WIT 11 Cd = 1.0 f PCT LOCATION-NODE 12 i t
- ~
Q 0 ex a a TIME, S - 7.39 -
l FIGURE 7 55 MAXIMUM ECCS STUDY DECLB, Cd = 1.0 [ CLADDING TEMPERATURE AT RUPTURE LOCATION i ~ o, 2000 - l ~ 1900 -< i l i ~ Y 1200 - j 800 - 2NG AT PD WITH Cd = 1.0 RUPTURE LOCATION-NODE 11 l 400 - t 0 0 200 400 e00 TlWE. S 'O FIGURE 7 56 MAXIMUM ECCS STUDY - DECLB, Cd = 1.0 CLADDING TEMPERATURE IN ADJACENT GRID SPAN l i 2G00 - j 1000 - w 1200 - / l r ANG AT PD WITH Cd== 1.0 1 NODE 15 { 400 - ) 1 h' 0 200 400 000 TIME S - 7.40 - )
t FIGURE 7 57a MAXIMUM ECCS STUDY DECLB, Cd = 1.0 3
- (.)
FLUID TEMPERATURE AT PCT LOCATION m0-I 1600 - g ~ 2NG AT PD WITH Cd = 1.0 i 1200 - PCT LOCATION-NODE 12 400 - 0 i. 0 4 8 12 16 20 24 18 TIME, S FIGURE 7 57b MAXIMUM ECCS STUDY DECLB, Cd = 1.0 FLUID TEMPERATURE AT PCT LOCATION 2000 - 2NG AT PD WITH Cd = 1.0 PCT LOCATION-NODE 12 1600 - a \\ g l 1j*($WAdaggg 400 -- l { l 0 200 400 600 j TIME, S j - 7.41 -
L i I FIGURE 7 58a MAXIMUM ECCS STUDY DECLB, Cd = 1.0 l FLUID TEMPERATURE AT RUPTURE LOCATION ao00 - 1000 - g 2A/G AT PD WITH Cd = 1.0 l 100 - NOOe 11 ~ l 800 - -sO 400 - t 0 O 4 8 12 16 20 24 26 l Twe,s FIGURE 7 58b MAXIMUM ECCS STUDY DECLB, Cd = 1.0 FLUID TEMPERATURE AT RUPTURE LOCATION I 2000 - ~ 2A/G AT PD WITH Cd = 1.0 NODE 11 goon _ ~ 1200 - i Y ~ l V' I \\ { l d %mMMIIf h 11111Il11ll 91y U l o u ,h l 3 400 - ~ g; 0 2. TIME, S I 7.42 - t .w.-, --n, ,a-. -_,-,.,-,,.-,--n a
y, i l o l ' 'x FIGURE 7 59a-MAXIMUM ECCS STUDY DECLB, Cd = 1.0 3 . (.I FLUID TEMPERATURE IN ADJACENT GRID SPAN ' = l: 2000 - 1600 -
- u. -
~ ' 2NG AT PD WITH Cd = 1.0 1200 - NODE 15 g soo. dw3 q O i i i 0 4 8 12 16 20 24 28 WE, S 9G FIGURE 7 59b MAXIMUM ECCS STUDY - DECLB, Cd = 1.0 FLUID TEMPERATURE IN ADJACENT GRID SPAN 2000 - ~ 2NG AT PD WITH Cd = 1.0 NODE 15 1600 - g l '": [694fth(hhMkh 800 - 400 - l3 l
- l.
O c ~ ~ ~ TIME, S - 7.43 - i m
L s P [ D 8. IDCA Limits The LOCA evaluation is completed with a set of analyses done to show compliance with 10 CFR 50.46 for the core power and peaking that will be taken as the limiting LOCA conditions for core operation, that is, the LOCA limits. The term limit is applied because these cases are run at the limit of allowable local power operation. ,Actually, these LOCA evaluations serve as the bases for the allowable local power. As
- such, the LOCA limits calculations comprise the cases that are used to demonstrate compliance of the reload fuel cycles and peaking limits to the criteria of 10 CFR 50.46.
Five runs are made at differing axial elevations such that a curve of allowable peak linear heat rates as a function of elevation in the core can be constructed or, in this case, confirmed. This curve becomes a part of the plant technical specifications, and plant operation is controlled such that the local peaking and power do not exceed the allowable values. nd' l' M LOCA Limits Conditions i The absolute LOCA limits to power and peaking for each elevation in the core can be determined through repeated calculations at i each elevation, with successively higher local power levels, L l until the analysis shows one or more of the applicable acceptance l criteria to be exceeded. The highest linear heat rate for which l the criteria are not exceeded is the absolute LOCA limit for a l particular elevation. The more practical
- approach, the one adopted for this report, assumes a set of peaking limits at a given power level that have been determined to be acceptable for fuel cycle design and plant operations purposes.
The LOCA limits analyses are then done to confirm that the assumed limits will meet the applicable criteria. { - 8.1 - s
r h ? 5 Figure 8-1 shows the axial power and peaking selected and g confirmed as applicable to the McGuire and Catawba plants for T l operation with Mark-BW fuel. With the axial power and peaking dependency established, LOCA calculations are performed with the ( core power level and total peaking initialized at different positions on the curve to demonstrate that these peaking i limitations assure compliance with 10 CFR 50.46. Should the results not comply, the allowed peaking is reduced, and. the i analysis is repeated until acceptable results can be. obtained. Likewise, if the results show large margins of compliance, the peaking may be increased to provide additional operational flexibility. For those analyses, neither of these steps was taken. although-the results do show considerable margins at certain elevations. 'An additional condition assumed in these analyses is that the allowable peaking will be dependent on fuel assembly burnup in accordance with Figure 8-2. This limitation is made necessary because, at burnups approaching 50000 mwd /MTu, the initial fuel enthalpy and internal pressure can become a more severe combination than the beginning-of-life values. By assuring that the local heating rates will be limited to those shown in Figure 8-2, the reduction in power compensates for the increases in fuel temperature and pin pressure such that the beginning-of-life conditions remain the most severe. (This is discussed in greater detail in the time-in-life sensitivity studies, Section 6.2.) Because. Figure 8-2 is a limit of operation which justifies the basis upon which the LoCA evaluation is constructed, it is part of the plant technical specifications. 3l LOCA Limits Results To validate Figure 8-1, five separate LOCA calculations were performed. Power peaks were run centered at the middle of the i second through the sixth grid spans. Figure 8-3 shows the axial - 8.2 - l l
r ~ I power shapes evaluated. For all cases, the radial power peaking (j ' was 1.55. The combination of the axial peaking of Figure 8-3 and a 1.55 radial yields the total peaking at the corresponding elevation shown in Figure 8-1. The results of the calculations are tabulated in Table 8-1 and shown in Figures 8-4 through 8-23. The figures comprise five sets with four figures in each set. The four figures of each set i show (1) the mass flux at the elevation of peak power, (2) the cladding temperature for three different locations on the pin, { (3) the heat transfer coefficient at the location of highest cladding temperature, and (4) the distribution of cladding oxidation along the pin. Only one mass flux plot is provided for each case because the axial variations in mass flux are not strong. This can be observed by comparing the five mass flux curves for the different peaking cases. .To demonstrate the cladding temperature results, three curves are presented for each case. Temperature histories are shown for the rupture location, for the node adjacent to the rupture, and for the high temperature node in an adjacent grid span. For p6 wor distributions peaked toward the middle of the core, rupture location is almost certain to correspond to the location of peak power. Near the time of
- rupture, the portion of the pin immediately above the rupture site will be at nearly the same
' temperature. Following rupture, the burst location cools quickly as the cladding pulls away from the fuel, and the area for heat transfer is increased. Due to axial heat conduction in the cladding and the effect of tne rupture on flow conditions, the cooling in the node just above the rupture is substantially improved. This means that although one of the nodes in the adjacent grid span is at a lower power, it can develop as the location of the highest cladding temperature. G - 8.3 -
The heat transfer coefficient (HTC) is shown for the peak cladding temperature location. HTC variations with elevation are as expected (see Figuren 7-51 through 7-53), such that the HTC from one elevation reasonably characterizes the other elevations. The last figure in each set shows the local oxide thickness as a i function of elevation for the fuel pin. Each figure show's total ' oxidation including that assumed prior to the start of the accident. Additionally, oxidation is included up to the time the cladding temperature falls below 1500 F or the elevation has been covered by mixture, as measured by the REFLOD3B core water level. The large variations of the resultant curve reflect the relatively lower cladding oxidation in the vicinity of the grid and rupture locations. 2.9-ft Peak Power Case In this case, the axial power shape is peaked well below the core
- midplane, and the cladding temperature responses differ accordingly from those calculated in the 4
,6 , and 8-ft cases. The peak power locations on the rod are cooled rapidly during reflood and have not reached temperatures sufficient to cause a rupture by the time of temperature turnaround. Therefore, the rupture occurs in node 8, the center node of the grid span above the location of peak power. This region of the core is also cooled rapidly, and the peak cladding temperature occurs in the grid span above the ruptured location. Although the power at the midplane is about 80 percent of that at the peak power location, the central node in the mid-core grid span produces the highest cladding temperature, 1816 F. The highest local oxidation, 3.4
- percent, occurs at the ruptured location.
The whole core oxidation calculated for this LOCA is 0.25 percent. 4.6-ft Peak Power Case With the power peaked at 4.6
- feet, the cladding temperature
- 8.4 -
3 i i i.' ( i responses resemble closely those obtained for the other two teid-j core peaks. The rupture occurs at the location of peak power. Q i' - L The node above the rupture experiences increased cooling post j
- rupture, and the peak. cladding temperature occurs in the i
downstream grid span (node 11). The temperature at this location 7 is about 100 F above the temperatures near the rupture location. The highest local oxidation, 5.2 percent, also occurs at the mid-core elevation. The whole core oxidation during this LOCA is 0.41 percent, the highest obtained in the set of LOCA limits c analyses. 1 p.3-ft Peak Power Case For a peak power situated at the core midplane, the cladding temperature response corresponds to that described in the l previous paragraph. The rupture is at the location of peak power. For this case, however, the post rupture cooling axial j conduction does not outweigh the effect of the relatively lower [D power at the next span, and the peak cladding temperature occurs ] '~) \\ in the node just above the rupture location. As shown in Figure 1 8-13, the peak temperature, 1873 F, is only slightly higher, by J L about 70 F, than that predicted for the next higher grid span. The highest local oxidation in this case, 4.8 percent, occurs for the peak cladding temperature node. The whole core oxidation is 1 0.40 percent. j. I-8.0-ft Peak Power Case ) Again, the temperature responses follow the pattern described for the previous two cases. Here, with the power peaked toward the
- outlet, the grid span that will produce high cladding temperatures lies below the location of peak power.
The rupture 1 occurs at the location of peak power, and the peak cladding l temperature, 1930 F, is predicted to occur in the grid span below the peak location. The markedly higher flow velocities at the 1 ( - 8.5 - 1 l 1' l
M j 1 ? i J higher elevations, in conjunction with rupture cooling effects and the drop-off of
- power, combine to produce a
cladding temperature in the node above the rupture location that is nearly 200 F'below the peak cladding temperature (node 12). The highest local oxidation is 4.7 percent, and the whole core oxidation is 0.32 percent. 9.7-ft Peak Power Case In accordance with the. axial dependency of power peaking shown in Figure 8-1, this case is run at a slightly lower total peaking than the other four cases. The location of peak power is in node l 17, which experiences some cooling due to grid effects. With the reduction in peaking-and the severe outlet shape, the power in node 15 is close to that in node 17. Because the lower location, node 15, is at the end of the grid span, there is little, if any, grid effect. Thus, node 15 is the first location on the fuel pin to reach the rupture. temperature. Since the rupture occurs at a node adjacent to the grid span, the rupture and spacer grid-effects combine to provide better cooling ' in a higher powered grid span. The peak temperature, 1823 F, occurs just below the rupture location. The peak local oxidation is 3.7 percent and the whole core oxidation is 0.29 percent. M Comoliance to 10 CFR 50.46 The LOCA limits calculations directly demonstrate compliance to two of the criteria of 10 CFR 50.4 6 and serve as the basis for demonstrating compliance with two others. As seen in the figures and in Table 8-1, the highest peak cladding temperature, 1963 F, and the highest local oxidation, 5.2 percent, are well below the ' 2200 F and 17 percent criteria. Chapter 9 documents compliance with the whole core oxidation limit based on the local oxidations calculated for these evaluations, and Chapter 10 documents the core geometry based on the deformations predicted for the LOCA - 8.6 - ,n -n. .m., + ,,.n..
N Table 8-1 LOCA Limits Results ?'%) Elevatien of Peak Power, Feet Item or Parameter 2.9 4.6 6.3 8.0 9.7 End of Blowdown, s 21.0 21.3 21.2 20.7 20.8 Liquid in Reactor Vessel 3 at EOB, ft 93.1 70.2 71.8 83.9 79.0 Bottom-of-Core Recovery, s 33.0 33.7 33.5 32.8 32.9 Time of. Rupture, s 81.8 74.4 67.6 73.8 84.4 Ruptured Node
- 8 8
11 14 15 PCT at Rupture Node, F 1611 1669 1666 1655 1602 Oxide at Rupture Node, % 3.4 3.5 4.8 1.5 0.8 Node Adjacent to Rupture
- 9 9
12 15 14 PCT of Adjacent Node, F 1804 1839 1873 1753 1823 Oxide at Adjacent Node, % 2.9 3.1 4.4 3.0 3.2 Node in Adjacent Grid Span 11 11 14 12 12 ,s / T J. PCT of Adjacent Grid s Span, F 1816 1963 1805 1930 1718 Oxide at Adjacent ' Grid Span, % 3.0 5.2 3.4 4.7 2.0 1 Pin PCT Node
- 11 11 12 12 14 I
Peak Local Oxidation, % 3.4 5.2 4.8 4.7 3.7 Whole Core Oxidation, % 0.25 0.41 0.40 0.32 0.29 L
- Refer to Figure 4-4 for noding arrangement I
l l 1 - 8.7 -
4 .c h FIGURE 81 AXIAL DEPENDENCE OF ALLOWED TOTAL PEAKING FACTOR I 1.ARGE BREAK LOCA MARK BW i 2.00 t \\ 2.40 n f ~ ~ 2.00 1.00 1.20 K 0.80 0.40 t 0 0 2 4 6 8 10 12 CORE ELEVATION, FEET O FIGURE 8-2 NORMALIZED LOCAL POWER-BURNUP DEPENDENCY FACTOR 1.20 s 0.80 O.00 0.40 i, Y s 0.20 0 2mm 400m ,000, j BURNUP, mwd /MTu L - 8.8 - l-
r %,, .: i .3, ', ' i .I 1 r ~ FIGURE 8 3 LOCA LIMIT STUDY AXIAL POWER SHAPES v i g(e] 2.0 .q 1.8 - )
- i 1.6 -
....,,........,.7,..,'>c,;;,. -~ .'N 1.4 - ,. '; -N--- sN s 4 .,e s N s 3-\\ + / 1.2 - / s s N \\ l s r 's, ,8 j s ',*y 'h. N \\ 1.0 - l ,r 3 ,s,,- N, K 8 l / , f fe' N o,e. ,/ s, N\\ ,.,,j .., /,:.<, - s,%.,,'s x y s.,- s f' p ., s g, f; P, 2.9 FOOT CASE '..3, t A.. 4.6 FOOT CASE O.4 - 6.3 FOOT CASE 8.0 FOOT CASE 9.7 FOOT CASE 0.2 - l l 0 i 0 20 40 00 80 100 120 140 CORE ELEVATION, INCHES O'; \\ t LJ FIGURE 8-4 LOCA' LIMITS STUDY 2.9 FOOT CASE L' MASS FLUX DURING BLOWDOWN AT PEAK POWER LOCATION 160 120 - 80 - 2NG AT PD WITH Cd = 1.0 i M- ~ g 0 -j-d
- )
40 - 30 - 1 120 - .p. 160 - -, f 0 4 8 12 16 20 24 20 ( TIME, S 8.9 -
n-4 l FIGURE 8 5 LOCA LIMITS STUDY 2.9 FOOT CASE
- /
CLADDING TEMPERATURES N' h 2000 - s# ~~"****' Q --- 1,00 j h 1200 - l ([ i / 2A/O AT PD WITH Cd = 1.0 soO - RUPTURE LOCATION-NODE 8 i 400 - ADJACENT TO RUPTURE LOCATION-NODE 9 PCT LOCATION-NODE 11 0 0 200 400 W TIME, S i-l FIGURE 8-6 LOCA LIMITS STUDY - 2.9 FOOT CASE HEAT TRANSFER' COEFFICIENT AT PCT LOCATION 1000 7 500 - I 100 ; 2NG AT PD WITH Cd = 1.0 g _ NODE 11 - { \\ l10 - i. p h, I 5- ~ 0 200 400 600 TIME, S i - 8.10 -
w r,,, b, x; ,- ( 'Sp' FIGURE 8-7 LOCA LIMITS' STUDY - 2.9 FOOT CASE O LOCAL OXIDATION ,. c. 10 ? i 8-t i. i s 6-l E 2NG AT PD WITH Cd = 1.0 i 1 I 2-0 0 2 '4 6 8 10 12 l CORE ELEVATION, FEET \\. A FIGURE 8 8 LOCA LIMITS STUDY 4.6 FOOT CASE L MASS FLUX DURING BLOWDOWN AT PEAK POWER-LOCATION 100 i 120 - 80 - 2NG AT PD WITH Cd = 1.0 40 - f V 40 - 80 - 120 - 160 0 4 8 12 16 20 24 28 TIME, S - 8.11 -
e 4 i i L 1 FIGURE 8 9 LOCA LIMITS STUDY - 4.6 FOOT CASE j CLADDING TEMPERATURES j 2*0 - p,.*........... ~~.,, j gs" 1000 - ,'%~._,. - l o' f n i 1200 - /.' 2A/G AT PD WITH Cd = 1.0 N g g om - i RUPTURE LOCATION-NODE 8 400 - - ADJACENT TO RUPTURE LOCATION-NODE 9 PCT LOCATION-NODE 11 0 0-200 400 600 TIME, S O, FIGURE 810 LOCA LIMITS STUDY - 4.6 FOOT CASE HEAT TRANSFER' COEFFICIENT AT PCT LOCATION ~ 1000. 1 J $00 - j 100 -. 2NG AT PD WITH Cd = 1.0 ) 50 NODE 11 ~ 10 ~ =. :
- x. -
n f
- p:.
- . =
i 5 l i 0 200 400 000 (. TIME, S - 8.12 -
te .t'. i FIGURE 811' LOCA LIMITS STUDY 4.6 FOOT CASE ~' (j ' ' L)t LOCAL OXIDATION { 10 I it ra g_ ~ f g + 6-2NG AT PD WITH Cd = 1.0 ~ I 2-0 0 2 4 6 8 10 12 l CORE ELEVATION, FEET ) f,') 1 Q. ' FIGURE 8-12 LOCA LIMITS STUDY 6.3 FOOT CASE i MASS FLUX DURING BLOWDOWN AT PEAK POWER LOCATION 160 L l 120? i h y 80 - 2NG AT PD WITH Cd = 1.0 l' 40 - 1 0-120 - C 160 ( 0 4 8 12 16 20 24 28 TIME, S 8.13 - w w, -m o r e -- m -,--wn e e- -w---
___..--s-,
F
- n 1
p-(! FIGURE 813 LOCA LIMITS STUDY 6.3 FOOT CASE l CLADDING TEMPERATURES i 2000 - ,e, .s ~"}[... A. '- ~B"~' " ~-~--*e.-..~......., I ___,,a 1000 - i e L 1200 - 2NG AT PD WITH Cd = 1.0 1 1 1 RUPTURE LOCATION-NODE 11 400 - PCT LOCATION-NODE 12 ADJACENT GRID SPAN-NODE 14 'O O E O' TIME. S O' FIGURE 6-14 LOCA LIMITS STUDY 6.3 FOOT CASE HEAT TRANSFER COEFFICIENT AT PCT LOCATION 1000 _ 1 t l h 1 , l 2A/G AT PD WITH Cd = 1.0 g; NODE 12 ' i q ---:-^^ 10 g vvv.* -e ; -- -~ 62 j 3 TIME, S 8.14 -
t ' fi ?
- w fh FIGURE 815 LOCA LIMITS STUDY - 6.3 FOOT CASE Cl ~
LOCAL OXIDATION 3 10 8-6- l 2NO AT PD WITH Cd = 1.0 ,1 4 '- t 2-i 0 'O 2 4 6 8 10 12 CORE ELEVATION, FEET FIGURE 816 LOCA LIMITS STUDY 8.0 FOOT CASE MASS FLUX DURING BLOWDOWN AT PEAK POWER LOCATION 160. ~ ,( 120 - 80 - 2NG AT PD WITH Cd = 1.0 40 0 L ? _ -120.- e 160 (s) 0 4 8 12 16 20 24 28 TIME S - 8.15 -
L s t 1 i 1 FIGURE 817 LOCA LIMITS STUDY 8.0 FOOT CASE CLADDING TEMPERATURES j 20C0 - ,?',,,,...-v~-----~~~~^-:==~~..,.g..___ e g, s_- k 1200 - 2A/G AT PD WITH Cd = 1.0 W RUPTURE LOCATION-NODE 14 400 - ADJACENT TO RUPTURE LOCATION-NODE 15 PCT LOCATION-NODE 12 ~ 0 0 200 400 600 TIME, S O-FIGURE 818 LOCA LIMITS STUDY 8.0 FOOT CASE HEAT TRANSFER COEFFICIENT AT PCT LOCATION 1000 ; 500 - 100 t 2A/G AT PD WITH Cd = 1.0 50 - NODE 12 5 ~ ~ R to -
- rq g
1 5 - 0 2b0 400 600 TIME, S - 8.16 - ..v .n,. -~
I i f f'~ 'y FIGURE. 819 LOCA LIMITS STUDY 8.0 FOOT CASE A_).. LOCAL OXIDATION 10 L. L 8-
- 1 1 1 i
6-2A/G AT PO WITH Cd = 1.0 4-i a i e 2-I b 0 0 2 4 6 8 10 12 CORE ELEVATION, FEET (~T j N. s ' FIGURE 8 20 LOCA UMITS STUDY 9.7 FOOT CASE MASS FLUX DURING BLOWDOWN AT PEAK POWER LOCATION 160-120 - 80 - 2NG AT PD WITH Cd = 1.0 40.- l g 80 - l. p 120 - l. ,1 1e0 j Q 0 4 8 12 16 20 24 28 TIME, S 1 l - 8.17 - 1
. +. i b l. r i FIGURE 8 21 LOCA LIMITS STUDY - 9.7 FOOT CASE CLADDING TEMPERATURES 2000 - l v rv.W^ I d b.3. ~ ~ ~... -A------ MNvNN %g 2"'; 1 1000 - l ' ~~ ~~- 1200 - ,j 2NG AT PD WITH Cd = 1.0 800 - w RUPTURE LOCATION-NODE 15 PCT LOCATION-NODE 14 400 - ADJACENT GRID SPAN-NODE 12 PEAK POWER LOCATION-NODE 17 0 i 0 200 400 600 TIME, S O FIGURE'8-22 LOCA LIMITS STUDY - 9.7 FOOT CASE. .r HEAT TRANSFER COEFFICIENT AT PCT LOCATION 1000 - ~ g 1 100 2NG AT PD WITH Cd = 1.0 50 NODE 14 I. 10 - fN \\D\\ I l .1 I1 j k 0 200 400 600 TIME, S - 8,18 - ~.
t.; .y-p (j FIGURE 8 23 LOCA LIMITS STUDY - 9.7 FOOT CASE c LOCAL OXIDATION t,..L/ 10 - - m 8-l- g ~ 6-2NG AT PD WITH Cd = 1.0 1 2.. h k g 0 2 4 6 8 10 12 CORE ELEVATION, FEET V I h - 8.19 -
I l 9. Whole-Core Oxidation and Hydrocen Generation . /_.h (f The third criterion of 10 CFR 50,46 states that the calculated 1 total amount of hydrogen generated from the chemical reaction of 'the cladding with water or steam shall not exceed 0.01 times the hypothetical amount that would be generated if all of the metal in the cladding cylinders surrounding the fuel, excluding the cladding surrounding the plenum volume, were to react. The I method - provided in the BWFC evaluation model, Reference 1, has been applied to determine corewide oxidation for each of the LOCA limits cases. In the calculations, local cladding oxidation was computed as long as the cladding temperature remained above 1500 F, and the REFLOD3B analysis did not show that the cladding wac within the core flooded region. The flooded region of the core was conservatively taken to be twenty percent above the core collapsed liquid level. These local oxidations are summed over the core to give the core-wide oxidation. The figures in Chapter 8 give the local oxidation for the hot pin, including the initial [ oxide layer. The only difference between these distributions and the ones used for the whole core calculation is that the initial oxide layer is subtracted before the integration in order to provide a measure of the hydrogen produced during the LOCA. The results of these calculations for each of the power distributions of the LOCA limits cases are: Case Whole Core Oxidation. % 2.9-ft Peak O.25 4.6-ft Peak O.41 6.3-ft Peak O.40 8.0-ft Peak O.32 9.7-ft Peak O.24 - 9.1 -
1 s( l L i Because these cases represent a range of the possible power distributions that can occur in.the plant, the maximum possible f oxidation:that can occur during a LOCA at the McGuire or Catawba l plants is calculated to be less than 0.41 percent. Thus, the I third' criterion of 10 CFR 50.46, which limits the reaction to 1 ' percent or less, is met with considerable margin. i I 1 1 J e i e' i i 6 Y,: 6 - 9.2 - o . + - = - -, ___m,- ..m., e--% .e,- - - - - -, - -, - - - -, - ~,. e
m 10. Core Geometry j s L) The fourth acceptance criterion of 10 CFR 50.46 states that I calculated changes in core geometry shall be such that the core remains amenable to cooling. The calculations in Chapter 8 l directly assess the alterations in core geometry, which result from the 14CA, at the most severe location in the core. These l calculations demonstrate that the fuel pin cooled successfully. As discussed in Section 7 of the original BWFC evaluation model report (Reference 1), clad swelling and flow blockage due to rupture can be estimated based on NUREG-0630. For the McGuire and Catawba plants, the hot assembly flow area reduction at rupture is less than 60 percent for all LOCA limits cases. Furthermc,re, the upper limit of possiblo channel blockage, based on NUREG-0630, is less than 90 percent. Neither 90 parcent blockage nor 60 percent blockage constitutes total subenannel cbstruction. As the position of rupture in a fuel assembly is distributed within the upper part of a grid span, subchannel blockage will not become coplanar across the assembly. Therefore, the assembly retains its pin - coolant channel - pin - coolant channel arrangement and is capable of passing coolant along the pin to provide cooling for all regions of the assembly. The effects of fuel rod bowing on whole core blockage are i considered in the BWFC fuel assembly and fuel rod designs, which minimize the potential for rod bowing. The minor adjustments of fuel pin pitch due to rod bowing do not alter the fuel assembly flow area substantially and the average subchannel flow area is preserved. Therefore, due to the axial distribution of blockage caused by rupture, no coplanar blockage of the fuel assembly will i l occur and the core will remain amenable to cooling. Deformation 1 of the fuel pin lattice at the core periphery can occur from the i i combined mechanical loadings of the LOCA and a seismic event. l These loads have been analyzed separately in the original plant structural designs to ensure that they have no adverse effect on - 10.1 -
the core cooling processes. The loadings and effect on the Mark-BW assembly are presented in the Mark-BW fuel mechanical design report (BAW-10172, Reference 7). Although deformations can occur, they are limited to the outer two or three points on the i lattice structure of the core and do not cause a subchannel flow area reduction larger than 35 percent. The fuel pins at these lattice points do not operate at power levels sufficient to produce a cladding rupture during LOCA. Therefore, the only reduction in channel flow area is from the mechanical effect, and the assemblies retain a coolable configuration. The consequences of both thermal and mechanical deformation of the fuel assemblies in the core have been assessed and the resultant deformations have been shown to maintain coolable core configurations. Therefore, the coolable geometry requirements of 10 CFR 50.46 have been met and the core has been shown to remain amenable to core cooling. O' o - 10.2 -
y,- l L d. /N 11. Lona-Term Coolina The fifth acceptance criterion of 10 CFR 50.46 states that the calculated core temperature shall be maintained at an acceptably low value, and decay heat shall be removed for the extended period of time required by the long-lived radioactivity remaining in the core. Successful initial operation of the ECCS is shown i by demonstrating that the core is quenched and the cladding temperature is returned to near saturation temperature. i Thereafter, long-term cooling is achieved by the pumped injection systems. These systems are redundant and able to provide a continuous flow of cooling water to the core fuel assemblies so long as the coolant channels in the core remain open. For a cold leg break, the concentration of boric acid within the core might induce a crystaline precipitation which could prevent the coolant flow from reaching certain portions of the core. This chapter presents the evaluation of the final stages of the initial operation of the ECCS, a discussion of the long-term supply of (J' water to the core, and a discussion of the procedures to prevent the build-up of boric acid in the core. 11.1 Initial Claddina Cooldqya The heat transfer models used to determine the peak cladding temperature are most conservative following the initiation of cladding cooldown and cannot be used directly to predict cladding quench. Rather, the occurrence of core quenching is determined from the core liquid inventory, as predicted by the REFLOD3B code, with a conservatively assumed 20 percent mixtilre swell. After quenching, core heat transfer is by pool nucleate boiling or by forced convection to liquid, depending on the location of the break in the reactor coolant system (cold leg breaks are in pool nucleate and hot leg breaks in forced convection). Either mechanism is fully capable of maintaining the core within a few [ ( - 11.1 -
n l.\\. degrees of. the saturation temperature of the coolant.
- Thus, within ten to fifteen minutes of the LOCA, the core has been returned to an a.ceptably low temperature level.
11.2 Extended Coolant Suoolv once the core has been cooled to low temperatures, maintaining i that condition relies upon the systems available to provide a continuous supply of coolant to the core. Detailed descriptions of the plant systems and functions are provided in the safety analysis reports for the McGuire and Catawba units (References 8 and 9). Provision for long-term core cooling with the ECCS as demonstrated in the FSARs is independent of the fuel design. Thus, the licensing basis for previous operation remains valid .for Mark-BW reload fuel. 11.3 Boric Acid Conceritration As discussed, the long-term cooling mechanism for a hot leg break is by forced convection to liquid; once established, the coolability of the core is assured and need not be further considered. For the cold leg break, however, there is no forced 1 flow through the core. The liquid head balances between the core j and the downcomer prevent ECCS water from entering the core at a rate faster than the rate of core boiling. Extra injection simply flows out of,the break and spills to the containment. With no throughput, the boiling in the core region acts to concentrate boric acid. To limit the concentration of boric . acid, the operator is required to establish a hot leg recirculation mode of operation within 15 hours of the initiation of the accident. In this mode, the piping is aligned so that injection takes place in both the hot and cold legs. By doing so, the amount of j. l -injection to the hot leg becomes a through-flow that can control - 11.2 - i i
c f the concentration of boric acid. The timing and effectiveness of the hot leg injection is established by demonstrating that the in-vessel concentrations are well below the soluability limits for the disolved solids. Therefore, there is no dependency on the fuel element design as the concentrations depend only on tho injection rate, the Reactor Coolant System geometry, and the core power level. Since none of these factors have been altered by the fuel change, the evaluation in the referenced FSARs remains valid for the plant. The operator actions and procedures to establish the operation are also described in the FSARs. 11.4 Adherence to Lona-Term Coolina Criterion ' Compliance to this criterion is demonstrated for the systems and components specific to the Catawba and McGuire units in the l. referenced FSARs and is not related to the fuel design. The initial phase of core cooling has been shown to result in low cladding and fuel temperatures. A pumped injection system V capable of recirculation is available and operated by the plant to provide extended coolant injection. The concentration of l dissolved solids has been shown to be limited to acceptable levels through the timely implementation of hot leg recirculation. Therefore, the capability of long-term cooling l I has been established and compliance to 10 CFR 50.46, demonstrated. l l 1 1' l l l l l l - 11.3 l l l .)
l
- 12. Small Break LOCA
. Q )' The. current licensing bases for the McGuire and Catawba plants f comprise - a spectrum of large and small break loss-of-coolant accidents (LOCAs) analyzed by Westinghouse and documented in j Chapter 16 of each of the plant final safety analysis reports t (FSARs). For operation of Catawba and McGuire with BWFC-supplied
- fuel, BWFC has reanalyzed the large break LOCA transient as presented in the foregoing chapters.
Reanalysis was considered necessary for two reasons: (1) for Catawba and McGuire, the large break is clearly the limiting loss-of-coolant event, and (2) the large break results can be sensitive to changes in fuel design. By these same considerations, reanalysis of the small break LOCA f for operation with Mark-BW reload fuel is not required; the 1 current SBLOCA results are not the most limiting or severe LOCA cases,. and SBLOCA evaluations are unaffected by the design differences between the Mark-BW and resident Westingbouse fuel assemblies. Thus, the referenced FSAR analyses, performed by ~ Westinghouse, remain the bases for plant licensing even after the cores are loaded with BWFC-supplied fuel assemblies. 12.1 SBLOCA Transient l SBLOCA transients can be generally characterized as developing in l I five distinct phases: (1) subcooled depressurization, (2) 1 j pump / loop flow coastdown and natural circulation, (3) loop
- draining, (4) vessel / core boil-off, and (5) long-term cooling.
These phases are examined in the following paragraphs as a lead-in to a discussion in the next section of the effects, if any, of i fuel design differences--between the resident Westinghouse fuel and the Mark-BW reload assemblies--upon the sequence of events and consequences of the small break LOCA transient for the l Catawba and McGuire units. l L l' - 12.1 - l l c
m I i l The limiting SBloCA events begin with a subcooled reactor coolant system (RCS) depressurization until the primary system pressure f rea:hes the initial hot leg temperature saturation pressure. f During this depressurization phase, the low pressure reactor trip, ECCS injection, and reactor coolant pump trip signals are j generated. Tripping of the pumps begins the pump and loop flow l coastdown period. i { To11owing reactor trip, the core power drops sharply. The initial forced flow and subsequent coastdown f 3 ow provide I continuous heat removal via the steam generators.
- Thus, the initial stored energy and the core power and decay heat during this phase are trandferred directly to the steam generators.
The i pump coastdown and natural circulation flows during this period i are sufficient to prevent critical heat flux (CHF) from occurring in the core. As a result, the fuel pins are cooled toward the j quasi-steady temperature distribution required to simply conutet i and corsvect the decay heat energy out of the pins. These pin f temperatures approach the RCS saturation temperature. Loss of j continuous loop flow marks the end of this period. The third phase in the transient is characterized as a period of loop draining. During this
- period, the system reaches a
quiescent state in which the core decay heat, leak flows, pumped f ECCS injection, and steam generator heat transfer combine to control the development of steam-water mixture levels within the RCS. The system inventory distribution is a strong function of j the system geometry and break location. RCS liquid inventory will continue to decrease until component mixture levels provide a continuous vent path for core steam production. Relief of core steam production allows the RCS to further depressurize end enter I the boil-off mode. The, develcpment and timing of events that mark the end of the ) loop draining and onset of core boil-off are governed by the - 12.2 - l ~ \\
I i l ] break location. For hot leg breaks, the continuous core staan j Q venting path is readily established. A significant system j inventory loss is required to establish the vent path for steam i g generator downstream piping breaks. The most severe of all f SBloCAs occurs in the cold leg pump discharge breaks. In these breaks, liquid inventory is lost until the primary levels descend to the spill-ender elevation at the low point in the cold leg pump suction piping. This liquid trap or loop seal must be t cleared of liquid to establish the steam venting path to the leak. Since the loop seal elevation is located slightly above l the middle of the core, the core collapsed liquid level will be depressed by the manometric pressure balance imposed by the RCS geometry. Once the loop seals clear, the steam venting path is I established and the residual liquid inventory in the pump discharge and downcomer regions drains into the core region. i i The onset of the boil-off period typically coir.cides with the beginning of a final saturated depressurization. Voiding at the break increases the leak volumetric flow rate which ultimately depressurizes the system until the accumulator fill pressure is reached or the pumped ECCS injection matches core steaming. j During this period, the reactor vessel mixture levels may drop f into the core heated region. Pin temperature excursions calculated for the upper elevations are maximized by the assumption of a bounding, core outlet-skewed peaking. During { these heatups, the cladding may swell and even rupture if the temperature approaches 1500 F. However, as long as the cladding temperature remains below 1800 F, the occurrence of rupture is treated conservatively by not including the effects of rupture in the evaluation. t 4 Swelling and rupture produce three primary effects on the temperature calculation. First, the cladding expansion increases the fuel pin gap allowing a momentary cooling of the clad. This condition is temporary;
- however, it delays the temperature
- 12.3 -
excursion, resulting in a lower peak cladding temperature because I the decay heat level has decreased slightly.
- Secondly, the rupture may divert flow out of the channel.
The evaluation model l uses only average channel flows to cool the hot channel. This l l flow is significantly lower than that expected in the hot channel. Thus, the effects of rupture-induced flow diversion are l already conservatively bounded by the modeling.
- Finally, the f
l ' inside of the cladding is exposed to metal-water reaction which l creates a new heat source. The metal-water reaction is expcnential with the cladding temperature, becoming significant l relative to decay heat levels as the temperature approaches 1800 i F. Below 1800 F the metal-water heating is a small fraction of the decay heat. The extra heat from the inside of the cladding t need not be considered to conservatively evaluate the cladding temperature so long as that temperature remains below 1800 F. l l The temperature exc'trsions are arrested as the combined ECCS flows exceed the core decay heat level and final core refill begins. The suppression of core steam production further depressurizes the RCS, and thus increases the ECCS injection flow I and hastens core refill. Eventually the RCS system will be depressurized to the containment pressure and the core will be l refilled. At this point, the start of a long-term cooling configuration has been established and the transient is mitigated. ) 12.2 Fuel Desian Effects fib 14CA transients are affected primarily by system design and core decay heat levels. Fuel assembly design influences the i calculated sequence of events only to the extent that it affects overall system behavior. In that regard, differences between the !! ark-BW reload fuel assemblies and the resident Westinghouse OFA assemblies should not materially affect the bounding SBLOCA sequences of the reference FSARs. The BWFC and Westinghouse - 12.4 - \\
l t g3 assemblies differ in the following areas: unrecoverable pressure () drops across the assemblies, initial fuel temperatures, initial l pin internal gas pressure, and the axial power profile. The j impact of each of these items, with respect to the controlling j aspects of the sBIOCA transient, will be evaluated in the i following paragraphs. l l Mark-BW fuel assemblies have unrecoverable pressure drops that l are approximately 1 psi lower than those of the Westinghouse OFA assemblies. The associated effect in overall loop pressure drop l would translate to less than 1 percent difference in the initial forced flow. At the same steady-state core power and effectively identical loop flows, the controlling hot leg initial temperature is also essentially unaffected. The maximum hot leg temperature l variation will be less than 1 F. Thus, the initial subcooled depressurization phase of the SBLOCA will be unaltered. The reactor trip signal and pump trips will occur at the same time in l the transient as in the reference FSAR calculations. O The impact of the fuel bundle resistance will be even less during the pump coastdown and natural circulation phase because the flows during this phase are much reduced. Significant margins exist such that CHF will not be exceeded. All of the initial stored energy in the fuel will still be transferred to and removed by the steam generators. Therefore, core resistance j variations will not change the fuel thermal transient or impact i the existing evaluations. Changes in the initial fuel temperature add or subtract overall l energy from the RCS. The initial fuel energy is removed from the fuel pin during the reactor coolant pump coastdown p h a s o, a r.d i l rejected from the system via the steam generators. Therefore, the initial fuel enthalpy of operatior has virtually no impact i f( ~ 12.5 - 1 , - +,. +e--.- ..-w-.
- __m_,
m
beyond the loop coastdown period. The core energy content during i the loop draining and boil-off mode will be identical to the current licensing base. The fuel pin internal gas fill pressures are similar to the Westinghouse values, out may differ slightly. The internal gas pressure could affect the fuel / cladding gap dimensions and rupture time. However, the fuel temperatures approach the system saturation temperature within a fraction of a minute following reactor trip. The impact of gap differences at low temperatures i is negligible. Since the SBLOCA temperatures peak at approximately 1500 F, the impact of a rupture or diffarence in rupture timing would be negligible. As a final point, the technical specifications for allowable local power levels, core peaking, for core elevations at or above 8 feet will not be increased due to the use of BWFC-supplied fuel. Thus, the axial power profile used by Westinghouse in the SBLOCA analyses remains bounding. This assures that the thermal load imposed on the fuel during a temperature excursion remains conservatively modeled. The thermal
- results, cladding temperatures, for the present FSAR evaluations are, therefore, conservative for Mark-BW fuel.
In summary, the core resistance variations will not affect the loop flows such that the controlling hot leg temperature or CHF points are altered. The steam generator heat removal rate during the flow coastdown period will compensate for any initial fuel stored energy fluctuations. All controlling parameters in the phases following the pump coastdown and natural circulation phase will be unchanged. Therefore, since the overall RCS geometry, initial operating conditions, licensed
- power, and governing phenomena are effectively unchanged, the existing FSAR calculations should remain bounding for operation of the Catawba and McGuire units with BWFC-supplied fuel.
- 12.6 -
d <~~ 12.3 Current FSAR Results The Westinghouse calculations of SBICCA accidents for the McGuire and Catawba units are not the limiting LOCAs as predicted by the NOTRUMP and JACTA-IV computer codes. The calculated results documented in the current McGuire and Catawba FSARs predict peak SBLOCA cladding temperatures less than 1500 F. All parameters i t are well within the acceptance criteria limits of 10 CFR 50.46. I Even wide variations in SBLOCA results would not cause the SBLOCA to be limiting.
- Thus, considerable margins exist such that variations in the SBLOCA results would not alter either the plant technical specifications or operating procedures, i
12.4 Como11ance with Accantance Criteria The existing SBLOCA calculations contained in the McGuire and Catawba FSARs are valid and bounding for the BWFC Mark-BW fuel. ( The reactor coolant system, decay heat levels, and other system controlling parameters remain unchanged by the reload fuel. A significant safety margin exists between the calculated results and 10 CFR 50.46 limits. The fuel design differences between the Westinghouse OFA and the BWFC Mark-BW do not substantially alter the results of SBloCA ovaluations. Adequate core cooling has already been demonstrated and does not need to be repeated because of the change in fuel design. The present SBLOCA evaluation calculations remain valid for the McGuire and Catawba fuel reloads supplied by BWFC. These analyses remain the small break evaluhtions of record for demonstrating compliance with the criteria of 10 CFR 50.46. I l ( O 12., - I w e- ,,. - -.,. +. m
I I I l i
- 13. References
-y i ) U i 1. RSG IDCA-B&W LOCA Evaluation Model for Recirculatina i Steam Generator Plants, BAW-10168, B&W Fuel Company, j July, 1988. f I 2. RELAPS/ MOD 2-B&W. An Advanced Proaram for Licht Water [ Reactor IDCA and Non-IDCA Transient Analysis, BAW-f 10164P, B&W Fuel Company, October, 1988. 3.
- SAVER, Dicital Computer Code to Determine Pressure
{ Drops, BAW-10072-A, Babcock & Wilcox, July, 1973. I { 4. CR AF"?2-FORTRAN Procram for Diaital Simulation of a Multinode Reactor Plant Durina Loss-of-Coolant, BAW-i i 10092, April, 1975, 5. V. H. Ransom et al., RELAPS/ MOD 2 CODE MANUAL, Volumes 1 and 2, NUREG/CR-4312, EGG-2396, August 1985. ) i 6. BEACH. Bent Estimate Analysis Core Heat Transfer, BAW-10166P, B&W Fuel Company, April, 1988. 7. Mark-BW Mechanical Desian Reoort, BAW-10172P, B&W Fuel t Company, July, 1988. l l 8. McGuire Nuclear Plant Upgraded Safety Analysis Report, j Duke Power Company, 1986 Update. 9. Catawba Nuclear Plant Upgraded Safety Analysis Report, l Duke Power Company, 1988 Update, i I i l - 13.1 - .4
l Annandir A. Evaluation of Transition Cores Q \\ During the period of transition from a full Westinghouse OFA core to a full Mark-BW core, the two types of fuel assemblies will reside next to each other in various mixes for several cycles of operation. This appendix addresses the LOCA and ECCS aspects of both types of fuel assemblies in the mixed core configuration. As will be shown, the Mark-BW assembly will experience slightly l better blowdown cooling in a mixed core and a slightly slower reflooding rate than it does in a full Mark-BW core. Likewise the OFA assembly will experience slightly higher blowdown temperatures in a mixed core and slightly increased core flooding rates than it would in a full OFA core. In both cases, the mixed I core impacts balance each other so that there is little change in expected peak cladding temperature for either fuel. Thus, the mixing of the fuel assemblies will not alter the LDCA evaluations of either fuel substantially and the evaluations of the fuels, performed as full
- cores, remain valid for the mixed core condition.
L.1 OFA and Mark-BW Desian Differences t t Table A-1 presents a comparison of the main design differences between the OFA and the Mark-BW fuel assemblies. Differences between the assemblies occur in the pin dimensions, the guide l tube dimensions, and the irrecoverable pressure drop across the assembly. Guide Tube and Instrument Tube Differences These differences do not affect a large portion of the fuel assembly and do not substantially affect the LOCA evaluation. Depending on core elevation, the flow area of the guide tube and instrument tube channels is larger for the Mark-BW than for the OFA. The shift in flow area from inside the guide and instrument O - A.1 - I g l l \\
i tubes to the heated channels is small and inconsequential. The f effect of the change has been included in the computer runs in Section A.2. Fuel and claddina Dimension Differences t The flow area of the heated channels is lower in the Mark-BW, the i heat transfer area of the fuel pin is increased, the clad thickness is increased, the gap between fuel and clad is slight 3y increased, and the radius of the fuel pellet is increased. The length of the pellet is decreased, but this dimension is not modeled by present techniques which treat the fuel as axially i uniform. The effect of these dimensional changes is not consequential. A smaller core flow area can have a beneficial effect on flooding rate if the coolant supply from the downcomer is limited. For McGoire and Catawba, however, the downcomer is nearly always full and ECCS water is spilling from the system. With the excess ECCS water available, the decrease in core flow area going from OFA to Mark-BW provides no benefit. The difference in cladding surface area (310.2 ft' per fuel 2 assembly for the MK-BW to 298.6 ft for the OFA) will affect the assemblies in their individual evaluations an hot assemblies but will not affect the average core calculations. The changes in heat transfer area are somewhat offset by the changes in material thicknesses. Overall, these have no substantial effect on other than the hot pin temperature calculations. The dimensional changes combine to make the Mark-BW materials more massive than the OFA materials. At the beginning of reflooding, the Mark-BW assemblies contain about five percent more stored energy than do the OFAs. Although this would tend to reduce the flooding rate, it is too small a difference to have a significant impact. The flooding rate comparison discussed later includes this effect and shows very low impact. - A.2 -
t i i ) (O i ) Fuel Assembiv Pressure Dron f The only difference between the assemblies that can produce a { meaningful and discernable change in Inca results is the decreased assembly pressure drop of the Mark-BW. Most of the l decrease in pressure drop occurs at the fuel assembly inlet nozzle where the OFA design offsets the pin ends from the top surface of the end nozzle. With no offset in the Mark-BW design, there is less turbulence and a lower pressure loss. The effect of the different pressure drop during blowdown is to divert some 6 flow away from the OFA assemblies and towards the Mark-BW. I During reflood the impact is on the whole core pressure drop which will gradually increase the flooding rate as the core makes t the transition from OFA to Mark-BW assemblies. i t M Assessment of Imoact on Peak Claddino Temoeratures The previous section assessed the differences between the two l fuel assembly designs and the immediate impact those differences would have on system evaluations. The differences in the guide tube /inetrument tube and the pin dimensions, will not significantly change the LOCA responses of the plants and need not be further evaluated. The third
- change, the assembly pressure
- drop, although only slightly different, should be considered further to assure that the resultant deviation is not substantial.
To determine the impact on cladding temperature predictions from a LOCA evaluation, changes in results caused by the mixing of the assemblies must be assessed relative to a base for which results are known. Two such bases exist. The McGuire and Catawba FSARs contain evaluations of the plants when fully loaded with the r Westinghouse OFA assembly; this report contains an evaluation when the plants are fully loaded with the Mark-BW assembly. The - A.3
r-t l following subsections assess the impact of the mixed core relative to these bases. The evaluation for the Mark-BW is made by determining the impact, of exchanging Mark-BW assemblies in i the average core for OFA assemblies, on the results presented in I the main body of this report. This procedure keeps a Mark-BW as the hot assembly. Likewise, the impact on the OFA assembly is assessed by exchanging OFA assemblies in the average core for j Mark-BW assemblies in the FSAR analyses. This keeps an OFA as the hot assembly. Mixed Core Imoact on the Mark-BW Evaluations t The higher pressure drop of the OFA assemblies present in the everage core will slightly increase the average core pressure drop during blowdown. This will divert some small percentage of t flow out of the average channel and into the hot assembly. Studies done during the evaluation model development show that, i for a change of this magnitude, the difference in cladding and fuel temperature at the end-of-blowdown will be 3d to 50 F. Thus, blowdown temperatures can be expected to decrease slightly. During the reflood period, the OFA assemblies will cause an increase in the core flow resistance that will make it harder for i water to flow into the core. This will produce a slight decrease in the flooding rate of a mixed core, when compared to a full i Mark-BW core, and a resultant increase in cladding temperature. A direct assessment was made using REFLOD3B to model a complete OFA core. The modeling included the appropriate guide tube, instrument tube, pin dimensional, and pressure drop changes. The decrease in flooding rate for a complete change of core was less than 2 percent. As expected, the core inlet mass flow increased and the core inventory increased by 2 to 5 percent. The decreased flooding rate will tend to cause an increase in i - A.4
t-cladding temperature on the same order of magnitude as the g ) decrease in temperature caused by the flow diversion, that is, an increase of about 30 F. ( 9 There is virtually no difference in cladding temperature during a LOCA for a Mark-BW assembly in a mixed core when compared to a Mark-BW in a pure core. The impact is certainly within the 50 F l reportability measure used in the most recent revision of 10 CFR f 50.46. Therefore, the calculations performed in this report, which have 200 to 300 F margins with respect to the limits of 10 CFR 50.46, can be applied to the licensing of the Mark-BW during mixed core operation. Mixed Core Impact on the OFA Evaluations To assess the impact of a mixed core on the OFA assembly, the FSAR results are used as a baseline. During blowdown, the introduction of Mark-BW assemblies into the average core with their lower pressure drop will cause some small amount of flow diversion out of the hot assembly and into the average core. From the BWFC studies this can be expected to cause a 30 to 57 F increase in the hot channel fuel and cladding temperatures at end-of-blowdown. t During reflooding, the replacement of OFA assemblies with the Mark-BW assemblies will decrease the core pressure drop and increase the flooding rate by less than 2 percent. The baseline here is a calculation of the flooding rate performed with a full OFA core. This increase in flooding rate will allow a small decrease in peak cladding temperature, estimated to be 30 F. Thus, the difference in cladding temperature during a LOCA for an OFA assembly in a mixed core when compared to an OFA in a full core is also marginal. The impact is, again, well within the current reporting criteria of 10 CFR 50.46. Therefore, the - A.5 - l l
calculations performed in the FSAR, which have 200 to 300 F margins to the limits of 10 CFR 50.46, can be applied to the licensing of the OFA during mixed core operation. W conclusions An assessment of the design differences between the OFA and the Hark-BW assemblies has concluded that the LOCA cladding temperatures for mixed core operation will not vary substantially from those calculated for the two designs in pure core operation. Furthermore, the calculations for each of the designs show considerable margin to 10 CFR 50.46 criteria. The peak cladding temperature reported for the Mark-BW in Chapter 8 of this report is 1963 F, the peak reported in the McGuire FSAR is 1841 F, and the peak in the Catawba FSAR is 1704 F. Therefore, the full core evaluations of the respective fuel assemblies can be applied for the licensing during mixed core operation. The analysis contained in the current FSAR will justify the uce of the OFA assemblies, and the technical specifications applied to those assemblies will be based on those analyses. The analysis presented in the main body of this report will be applied to the licensing of the Mark-BW during the mixed core period, and the technical specifications applied to the Mark-BW assemblies can be based on the assumptions of this analysis. Operational limits or technical specifications required by either analysis that are not directly applied to the fuel assemblies will be based on the analysis which generates the most stringent limit. - A.6 -
? 1 Table A-1 OFA/ Mark-BW Design Differences s MK-BW OFA Guide Thimbless p i' Upper section OD/r (in) 0.482/0.016 0.474/0.016 Lower Section OD/r (in) 0.429/0.016 0.429/0.016 Instrument Tube OD/r ' (in) 0.482/0.016 0.474/0.016 Puel Pin Pin OD (in) 0.374 0.360 Clad Thickness (in) 0.024 0.0225 Pellet OD (in) 0.3195 0.3088 Pellet Length (in) 0.400 0.507 Diametral Gap (in) 0.0065 0.0062 Pressure Drop Across Core (psi) 22.7 23.7 (at full flow) rs l l 9 ) - A.7 - (
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