ML20101N744

Nonproprietary TR, Mark-BW Fuel Assembly Application for Sequoyah Nuclear Units 1 & 2

ML20101N744
Person / Time
Site:	Sequoyah
Issue date:	03/31/1996
From:	FRAMATOME COGEMA FUELS (FORMERLY B&W FUEL CO.)
To:
Shared Package
ML20013A240	List: ML20101N698 ML20101N707 ML20101N731 ML20101N744
References
BAW-10220NP, BAW-10220NP-R, BAW-10220NP-R00, NUDOCS 9604090100
Download: ML20101N744 (286)

Text

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Non Proprietary BAW-10220NP Rev. O Topical Report March 1996 MARK-BW FUEL ASSEMBLY APPLICATION FOR SEQUOYAH NUCLEAR UNITS 1 AND 2 9

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! FRAMATOME COGEMA FUELS l P. O. BOX 10935

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j- LYNCHBURG, VIRGINIA 24506-0935 9604090100 960404 PDR ADOCK 05000327 P PDR l.,_ __ _-,,_ __.. - _ - . _ _ , _ _ , .

Non Proprietary FRAMATOME COGEMA FUELS P. O. BOX 10935 LYNCHBURG, VIRGINIA 24506-0935 Topical Report BAW-10220NP Rev. O March 1996 Key Words- LOCA. non-LOCA. Safety. Thermal. Hydraulic.

Mechanical. Transient. Water Reactor Abstract The Framatome Cogema Fuels will be delivering reload fuel to the Tennessee Valley Authority's Sequoyah Nuclear Plant Units 1 and 2 beginning in early 1997. This report presents a complete LOCA, non-LOCA, thermal-hydraulic, mechanical and containment evaluation for operation of the sequoyah Nuclear Units with Mark-BW reload fuel. Other NRC-approved, Framatome Cogema Fuels topical reports describe in detail the Mark-BW fuel assembly design; the LOCA, non-LOCA, mechanical, nuclear, and thermal-hydraulic methods and codes supporting the design and are l appropriately referenced herein. The analysis and evaluations presented in this report serve in conjunction with these other topical reports as reference for the future reload r,afety i evaluations applicable to cores with Framatome Cogema Fuels  ;

supplied fuel assemblies. The scope of events / evaluations is l consistent with that addressed in the final safety analysis  ;

report for Sequoyah and/or NUREG requirements. l l

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Non Proprietary l Acknowledoments

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v' Framatome Cogema Fuels wishes to acknowledge M. L. Miller and M.

V. Parece of Framatome Technologies, Inc; L. W. Newman, J. L.

Griffith, S. J. Shah, and J. R. Rodes of Framatome Cogema Fuels, and D. M. Lafever, W. M. Justice, J. F. Burrow, and A. J. Sanislo of the Tennessee Valley Authority for their efforts in performing the analyses that demonstrate the acceptability of Mark-BW fuel for use in Sequoyah and for their support in the preparation of i

this topical report.

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Non Proprietary Topical Report Revision Record Documentation Revision Description 1

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l l Noli Proprietary l

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Pace 1.0 Introduction . . . . . . . . . . . . . . . . . . . . . . 1-1 2.0 Summary and Conclusions . . . . . . . . . . . . . . . . 2-1 2.1 References . . . . . . . . . . . . . . . . . . . . 2-4 30 Resident Fuel Assembly Compatibility . . . . . . . . . . 3-1 3.1 Fuel Assembly Description and Component Measurements . . . . . . . . . . . . . . 3-1 3.2 Fuel Assembly Structural Testing . . . . . . . . . 3-3 3.2.1 Fuel Assembly Lateral Stiffness and Natural Frequency Tests . . . . . . . . 3-3 3.2.2 VANTAGE SH Fuel Assembly Holddown Spring Test . . . . . . . . . . . . 3-4 3.3 Hydraulic Flow Testing . . . . . . . . . . . . . . 3-5 3.4 Mark-BW Nuclear Design Evaluation . . . . . . . . . 3-6 3.5 References . . . . . . . . . . . . . . . . . . . . 3-6 4.0 Plant Description . . . . . . . . . . . . . . . . . . . 4-1 i

4.1 Physical Description . . . . . . . . . . . . . . 4-1 Reactor Vessel . . . . . . . . . . . . . . . . . . 4-1 Reactor Core and Fuel Assembly . . . . . . . . . . 4-1 l Reactor Coolant Loops . . . . . . . . . . . . . . . 4-2 Steam Generators . . . . . . . . . . . . . . . . . 4-2

,-~s 4.2 Description of Emergency Core Cooling System . . . 4-2

[ T 4.3 Plant Parameters . . . . . . . . . . . . . . . . . 4-4

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('~ 5.0 LOCA Analysis . . . . . . . . . . . . . . . . . . . . . 5-1 5.1 Computer Codes and Methods . . . . . . . . . . . . 5-2 5.2 Inputs and Assumptions . . . . . . . . . . . . . . 5-4 5.2.1 RELAPS Modeling . . . . . . . . . . . . . . 5-4 Core . . . . . . . . . . . . . . . . . . . . 5-5 Reactor Vessel . . . . . . . . . . . . . . . 5-5 Reactor Coolant Loops . . . . . . . . . . . 5-6 Reactor Coolant Pumps . . . . . . . . . . . 5-6 Pressurizer . . . . . . . . . . . . . . . . 5-6 Recirculating Steam Generator . . . . . . . 5-7 Break Characteristics . . . . . . . . . . . 5-7 Primary Metal Heat Model . . . . . . . . . . 5-7 Emergency Core Coolant System . . . . . . . 5-7 5.2.2 REFLOD3B Modeling . . . . . . . . . . . . . 5-8 5.2.3 BEACH Modeling . . . . . . . . . . . . . . . 5-9 5.3 Sensitivity Studies . . . . . . . . . . . . . . . 5-15 5.3.1 Evaluation Model Generic Studies . . . . . 5-15 RELAP5/ MOD 2-B&W Time Step Study . . . . . 5-15 RELAP5/ MOD 2-B&W Loop Noding Study . . . . 5-15 REFLOD3B Primary Coolant Pump Rotor Resistance Study . . . . . . . . . . 5-16 RELAP5/ MOD 2-B&W Break Noding Study . . . . 5-16 RELAP5/ MOD 2-B&W Pressurizer Location Study . . . . . . . . . . . . . . 5-16 t

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Non Proprietary Iable of Contents (cont'd)

Pace RELAP5/ MOD 2-B&W Core Crossflow Study . . . 5-16 RELAP5/ MOD 2-B&W Core Noding Study . . . . 5-17 5.3.2 Confirmable Sensitivity Studies . . . . . 5-17 RELAP5/ MOD 2-B&W Pump Degradation Study . . 5-17 REFLOD3B Loop Noding Study . ....... 5-18 Time-in-Life Study . . . . . ....... 5-18 5.3.3 Break Location . . . . . . . ....... 5-19 5.4 Plant-Specific Studies and Spectrum Analysis . . 5-22 5.4.1 Base Case . . . . . . . . . ....... 5-22 5.4.2 Minimum Emergency Core Cooling System Analysis . . . . . . ....... 5-22 5.4.3 Discharge Coefficient Study ....... 5-23 5.4.4 Break Type Study . . . . . . ....... 5-24 5.4.5 Containment Pressure Study . ....... 5-25 5.4.6 Gadolinia Fuel Pin Study . . ....... 5-26 5.5 LOCA Limits . . . . . . . . . . . . . . ....... 5-51 5.5.1 LOCA Limits Dependencies . . ....... 5-51 5.5.2 LOCA Limits Resul's c . . . . ....... 5-52 2.9-ft Peak Power Case . . . ....... 5-53 4.6-ft Peak Power Case . . . ....... 5-53 6.3-ft Peak Power Case . . . ....... 5-54 8.0-ft Peak Power Case . . . ....... 5-54 9.7-ft Peak Power Case . . . ....... 5-54 5.5.3 Compliance to 10CFR50.46 . . ....... 5-54 5.6 Whole-Core Oxidation and Hydrogen Generation . . 5-73 5.7 Core Geometry . . . . . . . . . . . ....... 5-73 5.8 Long-Term Cooling . . . . . . . . . ....... 5-74 5.8.1 Initial Cladding Cooldown . ....... 5-75 5.8.2 Extended Coolant Supply . . ....... 5-75 5.8.3 Boric Acid Concentration . . ....... 5-75 5.8.4 Adherence to Long-Term Cooling Criterion . 5-76 5.9 Small Break LOCA . . . . . . . . . ....... 5-76 5.9.1 Small Break LOCA Transient Description . . 5-76 5.9.2 Small Break LOCA Evaluation Model . . . . 5-78 5.9.3 Inputs and Assumptions . . . ....... 5-79 5.9.4 Analysis Results . . . . . . ....... 5-80 5.9.5 Compliance to 10CFR50.46 . . ....... 5-81 5.10 Evaluation of Transition Cores . . ....... 5-99 5.11 Summary and Conclusion . . . . . . . . . . . . 5-102 5.12 References . . . . . . . . . . . . . . . . . . . 5-105 6.0 Non-LOCA Transient Analysis And Evaluation . . . . . . 6-1 6.1 Plant Modeling and Input Assumptions . . . . . . . 6-2 6.1.1 RELAPS Model Description . . . . . . . . . . 6-2 Fuel Pin and Nucleonics . . . . . . . . . . 6-3 Reactor Core . . . . . . . . . . . . . . . . 6-3 Reactor Vessel . . . . . . . . . . . . . . . 6-4 Reactor Coolant Loops . . . . . . . . . . . 6-4 ,

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Non Proprietary i

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_ Table of contents (cont'd)

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Reactor Coolant Pumps . . . . . . . . . . . 6-4 Pressurizer . . . . . . . . . . . . . . . 6-4 Recirculating Steam Generator . . . . . . . 6-4 -

j Feedwater System . . . . . . . . . . . . . . 6-5

Steam Line . . . . . . . . . . . . . . . . . 6-5 ,

i Emergency Core Cooling System . . . . . . . 6-5 6.1.2 Core and Plant Parameters . . . . . . . . . 6-6 6.1.3 Reactor Protection System Setpoints . . . . 6-6 i 6.1.4 Power Distribution . . . . . . . . . . . . . 6-6 l 6.1.5 Reactivity Coefficients . . . . . . . . . . 6-7 6.1.6 Rod Cluster Control Assembly Insertion Characteristics . . . . . . . . . . . . . . 6-7 4 j 6.2 Condition II - Faults of Moderate Frequency . . . 6-15  !

4 6.2.1 Uncontrolled Rod Cluster Control  !

Assembly Bank Withdrawal From a Subcritical Condition . . . . . . . . . . 6-15

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5 6.2.2 Uncontrolled Rod Cluster Control Assembly Bank Withdrawal At Power . . . . 6-16 6.2.2.1 Analysis of Effects and l Consequences . . . . . . . . . . 6-17 l j 6.2.2.2 Environmental Consequences . . . 6-18  ;

i s 6.2.2.3 Conclusions . . . . . . . . . . 6-19 I

' [f-w ) 6.2.3 Rod Cluster Control Assembly Misalignment 6-19

\s_,/ 6.2.3.1 Analysis of Effects and Consequences . . . . . . . . . . 6-21 6.2.3.2 Conclusions . . . . . . . . . . 6-23 6.2.4 Uncontrolled Boron Dilution- . . . . . . . 6-24 6.2.5 Partial Loss of Reactor Coolant Flow . . . 6-25 6.2.6 Startup Of An Inactive Reactor Coolant Loop . . . . . . . . . . . . . . . 6-26 6.2.7 Loss of External Electrical Load and/or Turbine Trip . . . . . . . . . . . 6-26 6.2.7.1 Analysis of Effects and Consequences . . . . . . . . . . 6-27 6.2.7.2 Conclusions . . . . . . . . . . 6-29 6.2.8 Loss of Normal Feedwater . . . . . . . . . 6-29 6.2.9 Loss of Non-Emergency AC Power to the Station Auxiliaries . . . . . . . . . 6-30 6.2.10 Excessive Heat Removal Due to Feedwater System Malfunctions . . . . . . 6-30 6.2.11 Excessive Increase in Steam Flow . . . . . 6-31 6.2.12 Accidental Depressurization of the Reactor Coolant System . . . . . . . . 6-32 6.2.13 Accidental Depressurization of the Main Steam System . . . . . . . . . . . . 6-33 6.2.14 Spurious Operation of the Safety Injection System at Power . . . . . . . . 6-34 D /

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Non Proprietary Table of Contents (cont'd)

Pace 6.3 Condition III - Infrequent Faults . ....... 6-42 6.3.1 Inadvertent Loading of a Fuel Assembly into an Improper Position . ....... 6-42 6.3.2 Complete Loss of Forced Reactor Coolant Flow . . . . . . . . ....... 6-44 6.3.2.1 Analysis of Effects and Consequences . ....... 6-45 6.3.2.2 Conclusions . . . . ....... 6-46 6.3.3 Single Rod Cluster Control Assembly Withdrawal at Full Power . . ....... 6-46 6.4 Condition IV - Limiting Faults . . ....... 6-50 6.4.1 Rupture of a Main Steam Line ....... 6-50 6.4.1.1 Analysis of Effects and Consequences . . . ....... 6-52 6.4.1.2 Environmental Consequences . . . 6-55 6.4.2 Major Rupture of a Main Feedwater Pipe . . 6-56 6.4.3 Steam Generator Tube Rupture ....... 6-56 6.4.4 Single Reactor Coolant Pump Locked Rotor . 6-57 6.4.4.1 Analysis of Effects and Consequences . . . ....... 6-57 6.4.4.2 Conclusions . . . ....... 6-58 6.4.5 Rupture of a Control Rod Drive Mechanism Housing (Rod Cluster Control Assembly Ejection) . . . . . . . . . ....... 6-58 6.4.5.1 Analysis and Evaluation . . . . 6-59 6.4.5.2 Results . . . . . ....... 6-59 6.4.6 Steam Line Break at Power With Coincident Rod Withdrawal . ....... 6-59 6.4.6.1 Analysis of Effects and Consequences . . . ....... 6-60 6.4.6.2 Conclusions . . . ....... 6-61 6.5 Conclusions . . . . . . . . . . . . ....... 6-80 6.6 References . . . . . . . . . . . . ....... 6-80 7.0 Thermal-Hydraulic Evaluation . . . . . . . . . . . . . 7-1 7.1 Thermal-hydraulic Core Models . . . . . . . . . . . 7-1 7.1.1 LYNXT Modeling . . . . . . . . . . . . . . . 7-1 7.1.2 Form Loss Coefficients . . . . . . . . . . . 7-2 7.1.3 Core Power Distribution . . . . . . . . . . 7-2 7.1.4 Core Conditions . . . . . . . . . . . . . . 7-3 7.1.5 Engineering Hot Channel Factors . . . . . . 7-3 7.1.5.1 Local Heat Flux Engineering Hot Channel Factor . . . . . . . . 7-3 7.1.5.2 Average Pin Power Engineering Hot Channel Factor . . . . . . . . . . 7-3 7.1.6 Fuel Rod Bowing . . . . . . . . . . . . . . 7-4 7.1.7 Active Fuel Stack Height . . . . . . . . . . 7-4 7.1.8 Reactor Coolant Flow Rate . . . . . . . . . 7-4 viii

Non Proprietary s Table of Contents (cont'd)

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\ Page 7.2 Application of Statistical Core Design . . . . . 7-10 7.2.1 Statistical Design Limit -- SDL . . . . . 7-10 7.2.2 Thermal Design Limit -- TDL ....... 7-11 -

7.3 Calculation of Safety Limits . . . ....... 7-13 7.3.1 Hot Leg Boiling' Limits . . . ....... 7-13 i 7.3.2 DNB Limits . . . . . . . . . ....... 7-14 '

7.3.3 Conversion to AT Versus Tavs Coordinates . 7-14 7.4 Transient Analysis Methods . . . . ....... 7-17 7.4.1 SCD Transients . . . . . . . ....... 7-17 7.4.2 Non-SCD Transients . . . . . ....... 7-17 7.5 Maximum Allowable Peaking Limits . ....... 7-18 7.5.1 Safety Limit Maximum Allowable Peaking Limits . . . . . . . ....... 7-19 7.5.1.1 Safety Limit MAPS Variation With Power . . . . ....... 7-19 7.5.2 Operating Limit Maximum Allowable Peaking Limits . . . . . . . ....... 7-20 7.5.2.1 Operating Limit MAPS Variation with Power . . . . ....... 7-20 7.5.3 Core Safety'and Operating Limits . . . . . 7-20 7.5.3.1 Physics Check Cases . . . . . . 7-21 7.6 Mixed Core Analysis . . . . . . . . ....... 7-24

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\ 7.6.1 Methods and Models for Resident Fuel . . . 7-24 N ,/ 7.6.2 Mixed Core DNBR Analysis . . ....... 7-25 7.6.3 Mixed Core Pressure Drop, Lift and Crossflow . . . . . ....... 7-27 7.7 Fuel Thermal Performance Analysis . ....... 7-30 7.8 Thermal-hydraulic Conclusions . . . ....... 7-30 7.9 Thermal-hydraulic References . . . ....... 7-31 8.0 Mechanical Analysis and Evaluations . . . . . . . . . 8-1 8.1 Normal Operations . . . . . . . . . . . . . . . . . 8-1 l 8.1.1 Growth Allowance Evaluations . . . . . . . . 8-1 8.1.2 Fuel Assembly Component Stress Analysis . . 8-2 8.1.2.1 Fuel Assembly Holddown Springs . . 8-2 8.1.2.2 Guide Thimble . . . . . . . . . . . 8-3 8.1.2.3 Spacer Grids . . . . . . . . . . . 8-4 8.1.2.4 Top and Bottom Nozzles . . . . . . 8-4 l 8.1.3 Mechanical Compatibility . . . . . . . . . . 8-5 i 8.2 Faulted Condition Loads . . . . . . . . . . . . . . 8-5 8.2.1 Horizontal Seismic and LOCA . . . . . . . . 8-6 8.2.2 Vertical LOCA Analysis . . . . . . . . . . . 8-7 8.2.3- Fuel Assembly Component Stresses Under Faulted Condition Loads . . . . . . . 8-7 8.3 Fuel Assembly Shipping and Handling . . . . . . . . 8-8 8.4 Mechanical Conclusions . . . . . . . . . . . . . . 8-9

! 8.5 References . . . . . . . . . . . . . . . . . . . . 8-9 4

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Non Proprietary Table of contents (cont'd)_

Pace 9.0 Long Term Containment Integrity Evaluation . . . . . . . 9-1 9.1 Evaluation Results . . . . . . . . . . . . . . . 9-1 9.2 References . . . . . . . . . . . . . . . . . . . . 9-2 Appendix A: Computer Codes . . . . . . . . . . . . . . . . . A-1 A.1 LOCA and Non-LOCA Codes . . . . . . . . . . . . . . A-1 RELAP5/ MOD 2-B&W (BAW-10164) . . . . . . . . . . . . A-1 REFLOD3B (BAW-10171) . . . . . . . . . . . . . . . A-1 BEACH (BAW-10166) . . . . . . . . . . . . . . . . . A-1 A.2 Thermal-hydraulic Analysis Codes . . . . . . . . . A-1 LYNXT (BAW-10156A, Rev. 01) . . . . . . . . . . . . A-1 TACO 3 (BAW-10162P-A) . . . . . . . . . . . . . . . A-2 GDTACO (BAW-10184P-A) . . . . . . . . . . . . . . . A-3 A.3 Mechanical Codes . . . . . . . . . . . . . . . . . A-3 ANSYS (BWNT-TM-83) . . . . . . . . . . . . . . . . A-3 STARS . . . . . . . . . . . . . . . . . . . . . . . A-4 A.4 Physics Codes . . . . . . . . . . . . . . . . . . . A-5 NEMO (BAW-10180-A, Rev. 1) and CASMO-3 . . . . . . A-5 A.5 References . . . . . . . . . . . . . . . . . . . . A-5 O

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Non Proprietary l

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bl Table Pace 3-1 Comparison of the Mark-BW and VANTAGE SH Design . . . . 3-7 3-2 Mark-BW Prototype Test Program . . . . . . . . . . . . . 3-8 3-3 Summary of Fuel Assembly Test Results . . . . . . . . . 3-9 4-1 Plant Parameters and Operating Conditions . . . . . . . 4-4 5.4-1 Min vs. Max ECC Flow Comparison . . . . . . . . . . 5-28 5.4-2 Spectrum and Break Type Comparison . . . . . . . . . 5-28 5.4-3 Containment Backpressure Study Comparison . . . . . 5-28 5.5-1 LOCA Limits Results . . . . . . . . . . . . . . . . 5-55 5.9-1 Small Break LOCA Time Sequence Of Events . . . . . . 5-83 5.9-2 Small Break LOCA Results . . . . . . . . . . . . . . 5-83 5.10-1 VANTAGE SH/ Mark-BW Design Differences . . . . . . . 5-101 5.11-1 Summary of Large Break LOCA Limits Analyses . . . . 5-104 5.11-2 Summary of Small Dreak LOCA Analyses . . . . . . . . 5-104 6.1-1 Summary of Non-LOCA Assessment for Reload With Mark-BW Fuel . . . . . . . . . . . . . . . . . . 6-8 6.1-2 Typical Input Parameters and Initial Conditions for Transients . . . . . . . . . . . . . . . . . . . . 6-9 6.1-3 Trip Points and Time Delays to Trip . . . . . . . . 6-10 6.2-1 Sequence of Events for an Uncontrolled RCCA Bank Withdrawal at Power . . . . . . . . . . . . . . . . 6-35 6.2-2 Sequence of Events for Loss of Electrical ,

j -~g Load Without Automatic Pressure Control . . . . . . 6-35  !

) 6.3-1 Sequence of Events for Complete Loss of

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( ~,/ Reactor Coolant Flow . . . . . . . . . . . . . . . . 6-46 1 6.4-1 Sequence of Events for Main Steam Line 1 Break Upstream of the Steam Flow Measurement Device With Offsite Power Available . . . . . . . . 6-61 6.4-2 Limiting Core Parameters Used in Steam Line Break DNB Analysis . . . . . . . . . . . . . . . . . 6-62 6.4-3 Sequence of Events for Reactor Coolant Pump l Shaft Seizure (Locked Rotor) . . . . . . . . . . . . 6-63 I 6.4-4 Parameters Used in the Evaluation of the  !

Rod Cluster Control Assembly Ejection Accident . . . 6-63 l 6.4-5 Sequence of Events for a 0.583 Ft2 Steam Line I Break With Coincident Control Rod Withdrawal . . . . 6-63 7.1-1 Thermal-Hydraulic Analysis Design Parameters . . . . . 7-5 7.1-2 Core Thermal-Hydraulic Conditions . . . . . . . . . . 7-7 7.2-1 Statistical Core Design Application Summary . . . . 7-12 7.6-1 Comparison of Mark-BW, W VANTAGE SH and W Standard Fuel Nominal Designs . . . . . . . . . . 7-29 8-1 Limiting Load Conditions for Fuel Assembly Components 8-11 8-2 Summary of Reactor Coolant System Design Transients . 8-13 l 8-3 Comparison of Mark-BW and VANTAGE SH Critical l

Interface Dimensions . . . . . . . . . . . . . . . . . 8-14 8-4 Component Vertical LOCA Forces . . . . . . . . . . . . 8-15 8-5 Mark-BW Fuel Assembly Stress Analysis Results for SSE and Combined SSE plus LOCA Conditions . . . . . . 8-16

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Non Proprietary List of Tables Table Ea9.2 9-1 Comparison of Key Inputs Affected by the Fuel Change Important to the Containment Integrity Evaluation for Sequoyah . . . . . . . . . . 9-3 l

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Non Proprietary List of Figures i

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Figure Page 3-1 Mark-BW Fuel Assembly . . . . . . . . . ....... 3-10 3-2 Mark-BW Versus VANTAGE SH Unrecoverable Pressure Drop Comparison . . . . . . . . ....... 3-11 5.1-1 RSG Large Break LOCA EM Computer Code Interface . . 5-3 5.2-1 RELAP5 Large Break LOCA Sequoyah Noding

for Reactor Vessel and Core . . . . . ....... 5-11 l 5.2-2 RELAPS Large Break LOCA Sequoyah Noding for Primary Loops with Model 51 SGs . ....... 5-12 5.2-3 REFLOD3B Large Break LOCA Model . . . ....... 5-13 5.2-4 RELAPS/ BEACH Large Break LOCA Model . ....... 5-14 5.3-1 Time-In-Life Study Burnup Limit . . . ....... 5-21 i 5.3-2 Time-In-Life Study Max Average Fuel Temperature . . 5-21

! 5.3-3 Time-In-Life Study Hot Pin Internal Pressure ... 5-21 l 5.4-1 Plant Specific Studies Analysis Diagram ...... 5-29 5.4-2 Sensitivity Study - Base Model System Pressure During Blowdown . . . ....... 5-30

5.4-3 Sensitivity Study - Base Model Mass Flux

! During Blowdown at Peak Power Location ...... 5-30 5.4-4 Sensitivity Study - Base Model l Minimum Containment Back Pressure . . ....... 5-31

' 5.4-5 Sensitivity Study - Base Model Reflooding Rate . . 5-31 l 5.4-6 Sensitivity Study - Base Model Hot Channel l Quench Front and Collapsed Liquid Level ...... 5-32

( 5.4-7 Sensitivity Study - Base Model

! Hot Pin Cladding Temperature at PCT Segment 9 ... 5-32 5.4-8 Sensitivity Study - Base Model Hot Pin Cladding Temperature at Rupture Segment 11 5-33 5.4-9 Sensitivity Study - Base Model Hot Pin Cladding Temperature at Segment 12 .... 5-33 5.4-10 Pumped ECC Injection Flow Rate . . . ....... 5-34 l 5.4-11 ECC Study Downcomer Water Level . . . ....... 5-34 5.4-12 ECC Study Reflooding Rate . . . . . . ....... 5-35 5.4-13 ECC Study Hot Channel Quench Front I and Collapsed Liquid Level . . . . . ....... 5-35 5.4-14 Hot Pin Cladding Temperature at Segment 9 ..... 5-36 !

l 5.4-15 Hot Pin Cladding Temperature at Rupture Segment 11 . . . . . . . . . . . . . ....... 5-36 5.4-16 Hot Pin Cladding Temperature at PCT Segment 12 . . 5-37 5.4-17 Discharge Coefficient Study - Cd = 0.8 System Pressure During Blowdown . . . ....... 5-37 5.4-1B Discharge Coefficient Study - Cd = 0.8 l Mass Flux During Blowdown at Peak Power Location . 5-38 1

5.4-19 Discharge Coefficient Study - cd = 0.8 Reflooding Rate . . . . . . . . . . . ....... 5-38 5.4-20 Discharge Coefficient Study - Cd = 0.8

! Hot Channel Quench Front and l

Collapsed Liquid Level . . . . . . . ....... 5-39

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Non Proprietary List of Figures (cont'd)

Figure Pace 5.4-21 Discharge Coefficient Study - Cd = 0.8 Hot Pin Cladding Temperature at Segment 9 . . . . . 5-39 5.4-22 Discharge Coefficient Study - Cd = 0.8 Hot Pin Cladding Temperature at Rupture Segment 11 5-40 5.4-23 Discharge Coefficient Study - Cd = 0.8 Hot Pin Cladding Temperatures at PCT Segment 12 . . 5-40 5.4-24 Discharge Coefficient Study - Cd = 0.6 System Pressure During Blowdown . . . ....... 5-41 5.4-25 Discharge Coefficient Study - Cd = 0.6 Mass Flux During Blowdown at Peak Power Location . 5-41 5.4-26 Discharge Coefficient Study - Cd = 0.6 Reflooding Rate . . . . . . . . . . ....... 5-42 5.4-27 Discharge Coefficient Study - Cd = 0.6 Hot Channel Quench Front and Collapsed Liquid Level . . . . . . . . . . . . ....... 5-42 5.4-28 Discharge Coefficient Study - Cd = 0.6 Hot Pin Cladding Temperature at Segment 9 . . . . . 5-43 5.4-29 Discharge Coefficient Study - Cd = 0.6 Hot Pin Cladding Temperature at Rupture Segment 11 5-43 5.4-30 Discharge Coefficient Study - Cd = 0.6 Hot Pin Cladding Temperature at PCT Segment 12 . . 5-44 5.4-31 Break Type Study - Split System Pressure During Blowdown . . . ....... 5-44 5.4-32 Break Type Study - Split Mass Flux During Blowdown at Peak Power Location . 5-45 5.4-33 Break Type Study - Split Reflooding Rate . . . . . . . . . . . ....... 5-45 5.4-34 Break Type Study - Split Hot Channel Quench Front and Collapsed Liquid Level . . . . . . . . . . . . ....... 5-46 5.4-35 Break Type Study - Split Hot Pin Cladding Temperature at Segment 9 . . . . . 5-46 5.4-36 Break Type Study - Split Hot Pin Cladding Temperature at Rupture Segment 11 5-47 5.4-37 Break Type Study - Split Hot Pin Cladding Temperature at PCT Segment 12 . . 5-47 5.4-38 Containment Pressure Study Minimum Containment Backpressure . . ....... 5-48 5.4-39 Containment Pressure Study Reflooding Rate . . . . . . . . . . . ....... 5-48 5.4-40 Containment Pressure Study Hot Channel Quench Front and Collapsed Liquid Level . . . . . . . . . . . . ....... 5-49 l 5.4-41 Containment Pressure Study l

Hot Pin Cladding Temperature at Segment 9 . . . . . 5-49 xiv l

t

Non Proprietary List of Figures (cont'd)

Figure Page 5.4-42 Containment Pressure Study [

Hot Pin Cladding Temperature at Rupture Segment 11 5-50 5.4-43 Containment Pressure Study Hot Pin Cladding Temperature at PCT Segment 12 .. 5-50 5.5-1 LOCA Limits Study Peaking Limit . . . . . . . . . . . . ....... 5-56 5.5-2 LOCA Limits Study Burnup Limit . . . . . . . . . . . . ....... 5-56 5.5-3 LOCA Limits Study ,

Axial Power Profiles . . . . . . . . ....... 5-57 5.5-4 LOCA Limits Study Containment Backpressure . . . . . . ....... 5-57 5.5-5 LOCA Limits Study - 2.9 Foot Case Pressurizer Pressure . . . . . . . . ....... 5-58 5.5-6 LOCA Limits Study - 2.9 Foot Case Core Mass Flux . . . . . . . . . . . ....... 5-58 5.5-7 LOCA Limits Study - 2.9 Foot Case l Reflooding Rate . . . . . . . . . . . ....... 5-59 5.5-8 LOCA Limits Study - 2.9 Foot Case Hot Channel Quench Front and Collapsed Liquid Level . . . . . . . . . . . . ....... 5-59 5.5-9 LOCA Limits Study - 2.9 Foot Case

\_,/ Cladding Temperature . . . . . . . . ....... 5-60 5.5-10 LOCA Limits Study - 2.9 Foot Case Hot Channel Local Oxidation . . . ....... 5-60 5.5-11 LOCA Limits Study - 4.6 Foot tl9 a Pressurizer Pressure . . . , . . ....... 5-61 5.5-12 LOCA Limits Study - 4.6 Foot age Core Mass Flux . . . . . . . . . . . ....... 5-61 5.5-13 LOCA Limits Study - 4.6 Foot Case Reflooding Rate . . . . . . . . . . . ....... 5-62 i 5.5-14 LOCA Limits Study - 4.6 Foot Case l

Hot Channel Quench Front and Collapsed Liquid Level 5-62 5.5-15 LOCA Limits Study - 4.6 Foot Case

! Cladding Temperature . . . . . . . . ....... 5-63 5.5-16 LOCA Limits Study - 4.6 Foot Case l Hot Channel Local Oxidation . . . . . ....... 5-63 l 5.5-17 LOCA Limits Study - 6.3 Foot Case Pressurizer Pressure . . . . . . . . ....... 5-64 l 5.5-18 LOCA Limits Study - 6.3 Foot Case l

Core Mass Flux . . . . . . . . . . . ....... 5-64 5.5-19 LOCA Limits Study - 6.3 Foot Case Reflooding Rate . . . . . . . . . . . ....... 5-65 5.5-20 LOCA Limits Study - 6.3 Foot Case Hot Channel Quench Front and Collapsed l Liquid Level . . . . . . . . . . . . ....... 5-65

! (> .

Non Proprietary List of Figures (cont'd)  !

Fiaure Pace 5.5-21 LOCA Limits Study - 6.3 Foot Case Cladding Temperature . . . . . . . . ....... 5-66 5.5-22 LOCA Limits Study - 6.3 Foot Case Hot Channel Local Oxidation . . . . . ...... 5-66 5.5-23 LOCA Limits Study - 8.0 Foot Case Pressurizer Pressure . . . . . . . . ....... 5-67 5.5-24 LOCA Limits Study - 8.0 Foot Case Core Mass Flux . . . . . . . . . . . ....... 5-67 5.5-25 LOCA Limits Study - 8.0 Foot Case Reflooding Rate . . . . . . . . . . . ....... 5-68 5.5-26 LOCA Limits Study - 8.0 Foot Case Hot Channel Quench Front and Collapsed Liquid Level . . . . . . . . . . . . ....... 5-68 5.5-27 LOCA Limits Study - 8.0 Foot Case Cladding Temperature . . . . . . . . ....... 5-69 5.5-28 LOCA Linats Study - 8.0 Foot Case Hot Channel Local Oxidation . . . . . ....... 5-69 5.5-29 LOCA Limits Study - 9.7 Foot Case Pressurizer Pressure . . . . . . . . ....... 5-70 5.5-30 LOCA Limits Study - 9.7 Foot Case Core Mass Flux . . . . . . . . . . . ....... 5-70 5.5-31 LOCA Limits Study - 9.7 Foot Case Reflooding Rate . . . . . . . . . . . ....... 5-71 5.5-32 LOCA Limits Study - 9.7 Foot Case Hot Channel Quench Front and Collapsed Liquid Level . . . . . . . . . . . . ....... 5-71 5.5-33 LOCA Limits Study - 9.7 Foot Case Cladding Temperature . . . . . . . . ....... 5-72 5.5-34 LOCA Limits Study - 9.7 Foot Case Hot Channel Local Oxidation . . . . . ....... 5-72 5.9-1 RELAP5 Small Break LOCA Model Sequoyah Noding for Reactor Vessel and Core . . . . . ....... 5-84 5.9-2 RELAP5 Small Break LOCA Model Sequoyah Noding for Primary Loops with Model 51 SGs . ....... 5-85 5.9-3 Small Break LOCA Study Hot Channel Power Profile . . . . . . ....... 5-86 5.9-4 2-inch Pump Discharge Break Primary System Pressure . . . . . . . ....... 5-86 5.9-5 2-inch Pump Discharge Break Leak Flow Rate . . . . . . . . . . . ....... 5-87 5.9-6 2-inch Pump Discharge Break Pump Suction Loop Seals Levels . . . ....... 5-87 5.9-7 2-inch Pump Discharge Break Core Collapsed Level . . . . . . . . ....... 5-88 5.9-8 2-inch Pump Discharge Break Hot Rod Clad Temperature . . . . . . ....... 5-88 xvi

l Non Proprietary List of Figures (cont'd)

V Figure Page 5.9-9 3-inch Pump Discharge Break Primary System Pressure . . . . . . . ....... 5-89 5.9-10 3-inch Pump Discharge Break Leak Flow Rate . . . . . . . . . . . ....... 5-89 5.9-11 3-inch Pump Discharge Break Pump Suction Loop Seal Levels . . . . ....... 5-90 5.9-12 3-inch Pump Discharge Break Core Collapsed Level . . . . . . . . ....... 5-90 5.9-13 3-inch Pump Discharge Break Hot Rod Clad Temperature . . . . . . ....... 5-91 5.9-14 3-inch Pump Discharge Break Primary System Pressure . . . . . . . ....... 5-91 5.9-15 4-inch Pump Discharge Break Leak Flow Rate . . . . . . . . . . . ....... 5-92 5.9-16 4-inch Pump Discharge Break Pump Suction Loop Seal Levels . . . . ....... 5-92 5.9-17 4-inch Pump Discharge Break Core Collapsed Level . . . . . . . . ....... 5-93 5.9-18 4-inch Pump Discharge Break Hot Rod Clad Temperature . . . . . . ....... 5-93 5.9-19 6-inch Pump Discharge Break

, Primary System Pressure . . . . . . . ....... 5-94 5.9-20 6-inch Pump Discharge Break l s Leak Flow Rate . . . . . . . . . . . ....... 5-94

'~

5.9-21 6-inch Pump Discharge Break l Pump Suction Loop Seal Levels . . . . ....... 5-95 I i 5.9-22 6-inch Pump Discharge Break l Core Collapsed Level . . . . . . . . ...... 5-95  !

5.9-23 6-inch Pump Discharge Break Hot Rod Clad Temperature . . . . . . ....... 5-96 5.9-24 CCI.Line Break Primary System Pressure . . . . . . . ....... 5-96 5.9-25 CCI Line Break Leak Flow Rate . . . . . . . . . . . ....... 5-97 5.9-26 CCI Line Break Pump Suction Loop Seal L . . . . . . ....... 5-97 5.9-27 CCI Line Break Core Collapsed Level . . . . . . . . ....... 5-98 4 5.9-28 CCI Line Break Hot Rod Clad Temperature . . . . . . ....... 5-98 6.1-3 RELAPS/ MOD 2-B&W Model Used for Non-LOCA Analysis of the SQN Plant . . . . . . ....... 6-11 6.1-2 Doppler Power Coefficient . . . . . . ....... 6-13 6.1-3 RCCA Position Versus Time . . . . . . ....... 6-13 6.1-4 Normalized Rod Worth Versus Rod Fraction Inserted . 6-14 6.1-5 Tripped Rod Worth Versus Time . . . . ....... 6-14

{

^

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s/ xvii

Non Proprietary List of Figures (cont'd)

Figure 6.2-1 75 PCM/s Rod Withdrawal - Neutron Power ...... 6-36 6.2-2 75 PCM/s Rod Withdrawal - Thermal Power ...... 6-36 6.2-3 75 PCM/s Rod Withdrawal - Pressurizer Pressure .. 6-37 6.2-4 75 PCM/s Rod Withdrawal - Pressurizer Water Volume 6-37 6.2-5 75 PCM/s Rod Withdrawal - Core Average Fluid Temperature . . . . . . . . . . . . . ....... 6-38 6.2-6 75 PCM/s Rod Withdrawal - Calculated DNBR ,.... 6-38 6.2-7 Loss of Electric Load - Neutron Power ....... 6-39 6.2-8 Loss of Electric Load - Reactor Vessel Lower Plenum Pressure . . . . . . . . . . . ....... 6-39 6.2-9 Loss of Electric Load - Steam Generator Downcomer Pressure . . . . . . . . . . . . . . ....... 6-40 6.2-10 Loss of Electric Load - Average Core Fluid Temperature . . . . . . . . . . . . . ....... 6-40 1 6.2-11 Loss of Electric Load - Pressurizer Water Volume . 6-41 6.2-12 Loss of Electric Load - DNBR . . . . ....... 6-41 6.3-1 Four Pump Coastdown - Neutron Power . ....... 6-47 6.3-2 Four Pump Coastdown - Thermal Power . ....... 6-47 6.3-3 Four Pump Coastdown - Core Flow Fraction ..... 6-48 )

6.3-4 Four Pump Coastdown - Pressurizer Pressure .... 6-48 6.3-5 Four Pump Coastdown - DNBR . . . . . ....... 6-49 6.4-1 Main Steam Line Break With Offsite Power Available -

K Effective Versus Average Core Fluid Temperature . 6-64 6.4-2 Main Steam Line Break With Offsite Power Available -

Doppler Power Feedback . . . . . . . ....... 6-64 6.4-3 Main Steam Line Break With Offsite Power Available -

Single Train High Pressure Injection Flow ..... 6-65 6.4-4 Main Steam Line Break With Offsite Power Available -

Neutron Power . . . . . . . . . . . . ....... 6-65 6.4-5 Main Steam Line Break With Offsite Power Available -

Pressurizer Pressure . . . . . . . . ....... 6-66 6.4-6 Main Steam Line Break With Offsite Power Available -

Pressurizer Water Level . . . . . . . ....... 6-66 6.4-7 Main Steam Line Bredk With Offsite Power Available -

Single Loop Reactor Vessel Inlet Temperature ... 6-67 6.4-8 Main Steam Line Break With Offsite Power Available -

Triple Loop Reactor Vessel Inlet Temperature ... 6-67 6.4-9 Main Steam Line Break With Offsite Power Available -

Core Average Temperature . . . . . . ....... 6-68 6.4-10 Main Steam Line Break With Offsite Power Available -

Core Reactivity . . . . . . . . . . . ....... 6-68 6.4-11 Main Steam Line Break With Offsite Power Available -

Core Boron Concentration . . . . . . ....... 6-69 6.4-12 Main Steam Line Break With Offsite Power Available -

Single RSG Feedwater Flow . . . . . . ....... 6-69 6.4-13 Main Steam Line Break With Offsite Power Available -

Triple RSG Feedwater Flow . . . . . . ....... 6-70 xviii

Non Proprietary

) List of Figures (cont'd) v Figure Pace 6.4-14 Main Steam Line Break With Offsite Power Available -

Single RSG Steam Flow . . . . . . . . ....... 6-70 6.4-15 Main Steam Line Break With Offsite Power Available -  ;

Triple RSG Steam Flow . . . . . . . . ....... 6-71 6.4-16 Main Steam Line Break With Offsite Power Available - )

Single RSG Steam Pressure . . . . . . ....... 6-71 6.4-17 Main Steam Line Break With Offsite Power Available -

Triple RSG Steam Pressure . . . . . . ....... 6-72 6.4-18 Main Steam Line Break With Offsite Power Available - l Reactor Coolant Flow . . . . . . . . ....... 6-72 1 6.4-19 Locked Rotor - Neutron Power . . . . ....... 6-73 6.4-20 Locked Rotor - Thermal Power . . . . ....... 6-73 6.4-21 Locked Rotor - Core Flow Fraction . . ....... 6-74 6.4-22 Locked Rotor - Pressurizer Pressure . ....... 6-74 I 6.4-23 Locked Rotor - Clad Surface Temperature . . . . . . 6-75 6.4-24 Locked Rotor - DNBR . . . . . . . . . ....... 6-75 6.4-25 Rod Ejection - Neutron Power Fraction ....... 6-76 6.4-26 Rod Ejection - Average Rod Thermal Response . . . . 6-76 6.4-27 SLB with RWAP - Neutron Power . . . . ....... 6-77 6.4-28 SLB with RWAP - Thermal Power . . . . ....... 6-77

, - 6.4-29 SLB with RWAP - Pressurizer Pressure ....... 6-78 j 6.4-30 SLB with RWAP - T average . . . . . . ....... 6-78 j i [j

( 6.4-31 SLB with RWAP - Steam Pressures . . . .......

6.4-32 SLB with RWAP - DNBR . . . . . . . . .......

6-79 6-79 7.1-1 LYNXT 31-Channel Bundle-By-Bundle Model . . . . . . . 7-8 7.1-2 LYNXT 12-Channel Model . . . . . . . . . . . . . . 7-9 7.3-1 Typical Reactor Core Safety Limits . . . . . . . . 7-15 7.3-2 Typical Overtemperature Limit Lines . . . . . . . . 7-16 7.5-1 Typical 118% Power Safety Limit MAP Curves . . . . 7-22 7.5-2 Typical 100% Power Operating Limit MAP Curves . . . 7-23 1 8-1 Horizontal Core S;ismic and LOCA Model . . . . . . 8-17 8-2 Fuel Assembly Vertical Model . . . . . . . . . . . 8-18 l

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l Non Proprietary I

List of Acronyms l

AC - Alternating Current AFD - Axial Flux Difference ANS - American Nuclear Society ARC - Alliance Research Center ASME - Ame: 1can Society of Mechanical Engineers B&W - Babcock and Wilcox Company BOL - Beginning of Life BWNT RSG LOCA EM - B&W Nuclear Technologies Recirculating Steam Generator Loss of Coolant Accident Evaluation Model CC - Centrifugal Charging Cd - Discharge Coefficient CHF - Critical Heat Flux CLB - Cold Leg Break CNFP - Commercial Nuclear Fuel Plant CRDL - Control Rod Drive Line DECLG - Double Ended Cold Leg Guillotine DNB - Departure from Nucleate Boiling DNBR - Departure from Nucleate Boiling Ratio ECC - Emergency Core Cooling ECCS - Emergency Core Cooling System EOB - End of Blowdown EOE - End of Event EOL - End of Life FCF - Framatome Cogema Fuels FSAR - Final Safety Analysis Report HCF - Hot Channel Factor HLB - Hot Leg Break HPSI - High Pressure Safety Injection ID - Inside Diameter LOCA - Loss of Coolant Accident M/W - Metal / Water MAP - Maximum Allowable Peaking MDFR - Mechanical Design Flow Rate NRC - Nuclear Regulatory Commission OD - Outside Diameter OD/t - Outside Diameter / thickness OFA - Optimized Fuel Assembly (Westinghouse design)

OPAT - Overpower Delta Temperature OTAT - Overtemperature Delta Temperature PCT - Peak Cladding Temperature PIE - Post Irradiation Examination PORV - Power-Operated Relief Valve PWR - Pressurized Water Reactor RCCA - Rod Cluster Control Assembly RCS - Reactor Coolant System l RHR - Residual Heat Removal l RSG - Recirculating Steam Generator RTP - Reactor Thermal Power xx

Non Proprietary ws List of Acronyms (cont'd)

RV - Reactor Vessel RWST - Refueling Water Storage Tank SCD - Statistical Core Analysis i SDL - Statistical Design Limit SG - Steam Generator SI - Safety Injection SQN - Sequoyah Nuclear Plant  ;

SRSS - Square Root of Sum of Squares  ;

SSE - Safe Shutdown Earthquake '

TDL - Thermal Design Limit TFTR - Transportable Flow Test Rig j TVA - Tennessee Valley Authority I

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Non Proprietary f-^s 1.0 Introduction

/ \

b Framatome Cogema Fuels (FCF) will be delivering reload fuel to the Tennessee valley Authority (TVA) for use in their Sequoyah Nuclear Plant (SON) Units 1 and 2 beginning in aarly 1997. The FCF-supplied Mark-BW fuel assembly for SQN is similar in design and has compatible performance characteristics as assemblies previously licensed and operated in SON. This Mark-BW fuel design has been licensed for use in Duke Power Company's Catawba and McGuire units and Portland General Electric's Trojan unit.

This same design is now being applied to TVA's SON units.

In justifying the application of the Mark-BW fuel for SON, FCF conducted a series of mechanical and hydraulic tests and measurements on the resident fuel at SQN to verify compatibility.

These tests and measurements were followed by a series of analyses and/or evaluations that encompass the regulatory requirements for each of several disciplines. These analyses and evaluations were performed to establish Mark-BW application for thermal-hydraulic, mechanical, and neutronic characteristics, and for LOCA, non-LOCA, and containment building response. The summary and conclusions of this work is provided in Section 2 of this topical report.

__ After the Summary and Conclusions, the remainder of this report is organized to provide the details of the work mentioned above.

s Each section is devoted to a category or discipline such that each section is complete within itself, including references.

Where NRC-approved topical report references are appropriate, a brief discussion is presented herein with reference to the approved topical report for details. The appendix contains brief descriptions of the computer codes with reference to NRC-approved topical reports that provide detailed descriptions.

Section 3 contains the evaluation of the resident Westinghouse VANTAGE SH fuel assemblies and shows that they are compatible with the Mark-BW. Section 4 presents a brief description of the SQN plant systems. Section 5 provides the large- and small-break LOCA analyses. Section 6 addresses the non-LOCA events from Chapter 15 of the SON FSAR. Sections 7 and 8 contain the thermal-hydraulic and mechanical efforts, respectively. Section 9 presents the long term containment integrity analysis evaluation.

(O V) 1-1

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l Non Proprietary l l

2.0 Summarv and Conclusions l

l

'\_s/

i Framatome Cogema Fuels (FCF) will be delivering reload fuel to l

the Tennessee Valley Authority (TVA) for use in their Sequoyah Nuclear Plant (SQN) Units 1 and 2 beginning in early 1997.

The SQN plant uses a nuclear steam supply system design by l Westinghouse that is representative of the standard Westinghouse  ;

l four-loop, 193 fuel assembly, 3411 MWt design.

The assembly design being supplied is the Mark-BW design described in BAW-10172P (Reference 2-1). The Mark-BW 17x17 fuel assembly includes these key features for compatibility with Westinghouse-design fuel asremblies:

- Leaf-type holddown springs

- Dashpot region in the guide thimbles for control rod deceleration  !

- Mixing vanes on selected intermediate spacer grids l

- Floating spacer grid restraint system This design was previously approved for use in Duke Power Company's Catawba and McGuire units and Portland General

-~s Electric's Trojan unit (References 2-2 through 2-8). In these

[(

applications, the Mark-BW assembly was shown to be compatible with the Westinghouse standard and OFA fuel assembly designs.

For application to SQN, mechanical and hydraulic tests and measurements, and neutronic evaluations were conducted to establish that the Mark-BW fuel assembly is structurally, l hydraulically, and neutronically compatible with the Westinghouse-supplied VANTAGE SH fuel assembly currently in use at SQN. Using two VANTAGE SH fuel assemblies supplied by TVA, data were collected from the interface dimensions and key performance features. Damping and harmonic tests were performed on these assemblies to verify similarity with the Mark-BW and to obtain results for future use in mixed core mechanical analyses.

The hydraulic tests were performed in FCF's cold-water loop, which has been shown to be acceptable through previous hot-to-cold loop results comparisons. These flow tests provide a distribution of pressure drops through the length of the assembly. Resultant assembly and component pressure drop values were used to determine form loss coefficients for the VANTAGE SH assembly and were applied in the thermal-hydraulic evaluation of a mixed core of VANTAGE SH and Mark-BW assemblies.

From a neutronic standpoint, the VANTAGE SH and the Mark-BW fuel assemblies are almost identical. The structural materials within

/N

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2-1

Ncn1 Prorncietary the active fuel region are similar in composition and weight.

The slight differences in uranium loading will be modeled such that isotopic composition and burnup differences are properly calculated. Thus, the use of the Mark-BW assembly in conjunction with the Westinghouse VANTAGE SH assembly in the core does not adversely affect plant operation or neutronic parameters.

In accordance with requirements of 10CFR50.46, an evaluation of the emergency core cool'ag system (ECCS) performance was performed for FCF Mark-BW reload fuel in SON utilizing the guidelines of 10CFR50, Appendix K. At the time of initial operation, SQN was fueled by Westinghouse with standard fuel.

Later the Westinghouse VANTAGE 5H design was implemented at SON.

Compliance was demonstrated by Westinghouse for both fuel types.

FCF calculations and evaluations documented herein demonstrate that SON continues to meet these criteria when operated with Mark-BW fuel assemblies. Large break LOCA calculations performed in concurrence with an approved evaluation model (BAW-10168 and revisions - Reference 2-9) demonstrate compliance up to and including the double ended severance of the largest primary coolant pipe. The small break LOCA calculations also demonstrate that the plant meets 10CFR50.46 criteria for small breaks when loaded with Mark-BW fuel. The coexistence of Mark-BW fuel with resident Westinghouse fuel is shown to be inconsequential and does not cause the calculated temperatures for the different assembly types to approach the limits of 10CFR50.46.

Specifically, for LOCA the results demonstrate that when SON is operated with Mark-BW fuel:

1. The calculated peak clad temperatures for the limiting cases are less than 22001

2. The maximum calculated local cladding oxidation is less than 17.0%.

3. The maximum amount of core-wide oxidation does not exceed 1.0% of the fuel cladding.

4. The cladding remains amenable to cooling. l l

S. Lona-term coolina is established and maintained after the I LOCA.

All FSAR Chapter 15 non-LOCA transient events were evaluated using approved methodology (BAW-10169P - Reference 2-10). For those transients potentially affected by operation of Mark-BW l fuel, the bounding cases were reanalyzed. For other transients, i the relevant core-related parameters, pertinent to future j reloads, were evaluated for their effect on transient events. l The events specifically analyzed were the following: i 2-2

Non Proprietary

,m I l - Control rod bank withdrawal at power

\s_,) - Loss of forced reactor coolant flow

- Locked reactor coolant pump rotor

- Loss of electrical load

- Main steam line break

- Steam line break with coincident rod withdrawal The non-LOCA analyses and evaluations confirme'd that operation of SQN units for reload cycles with Mark-BW fuel continues to be bounded by the previously reviewed and licensed safety limits.

Reanalyses of the transients affected by the fuel reloads demonstrate that tne acceptance and design criteria specified in Regulatory Guide 1.70 continue to be met.

Thermal-hydraulic analyses were performed to support the Mark-BW fuel application and future Mark-BW reload fuel use at SQN. The result of the thermal-hydraulic analyses is a set of operating limits that insure fuel and clad integrity are maintained during normal operation and transients of moderate frequency. The design criteria that were established and met to achieve this goal were as follows:

1. During Condition I and II events, there must be at least a 95 % probability with a 95% confidence level that the hot

.-s

, pin will not experience a departure from nucleate boiling

[ T (DNB); or a 99.9% probability that DNB will not occur core (s_,/ Wide.

2. During Condition I and II events, there must be at least a 95 % probability with a 95% confidence level that no fuel will experience centerline melting.

The structural design requirements for the Mark-BW fuel assembly were derived in large part from FCF experience, both in design and in-core operation of similar designs. For application to SON, plant specific design requirements and parameters augmented the established Mark-BW design criteria. These requirements in total are consictent with the acceptance criteria of the Standard Review Plan (NUREG-0800), Section 4.2, and follow the guidelines established by Section III of the ASME code. Code Level A criteria are used for normal operational and Code level D criteria are used for LOCA/ seismic. The design bases analyses performed to verify the adequacy of the Mark-BW fuel assembly in SON were as follows:

o Normal operations

- Growth allowance l - Fuel assembly holddown

- Guide thimble buckling

- Spacer grid loads

/N s_/ 2-3 l

Non Proprietary I

l - Interface with adjacent assembly

- Lateral seismic and LOCA loading l - Fuel assembly vertical LOCA loading

! - Fuel assembly component stress

- Shipping and handling loads These analyses and evaluations / review of the content of the NRC-approved topical report BAW-10172P, " Mark-BW Mechanical Design Report," confirm that the Mark-BW fuel assembly maintains its mechanical integrity when operated in SON either as a full complement of Mark-BW assemblies or in conjunction with the resident fuel assemblies.

An evaluation of the long term containment integrity as detailed in Chapter 6 of SON's FSAR shows that when the SQN core contains Mark-BW fuel the existing analysis results remain bounding. The important aspects of the fuel change that potentially impact the )

analysis are changes in the flow characteristics past the fuel, the reactor coolant system average operating temperature, the core stored energy and fuel heat capacity, and the decay heat.

Each aspect was evaluated considering detailed assembly testing and measurement results and/or comparisons with existing I

analyses. S9. sed on these evaluations, there is no consequence to the containment systems when the SON units are fueled with FCF-supplied Mark-BW assemblies.

All analyses and evaluations presented herein confirm and justify the operation of TVA's SQN units with Mark-BW fuel reloads in combination with resident fuel design or a complete core of Mark-BW assemblies.

2.1 References 2-1 BAW-10172P, Mark-BW Mechanical Design Report, July 1988.

2-2 BAW-2119, Reload Report - Catawba Unit 1 Cycle 6, October 1990.

2-3 BAW-10173P Rev. 2, Mark-BW Reload Safety Analysis for Catawba and McGuire, November 1990.

2-4 BAW-10174, Mark-BW Reload LOCA Analysis for Catawba and McGuire, September 1989.

2-5 BAW-2136 Rev. 2, Reload Safety Evaluation Trojan Cycle 14, November 1991.

2-6 BAW-10176P, Mark-BW Reload Safety Analysis for Trojan Nuclear Plant, March 1989.

(

! 2-4

. .. - . - . . _ - ~ . . . . . . . _ . - - . .__.-.. . . _ _ _ - - - - . _ - . - - = . . . - - .

5 Non Proprietary  !

2-7 BAW-10177P, Mark-BW Reload LOCA Analysis for Trojan Nuclear Plant, January 1990.

k.

2-8 BAW-10178P, Mark-BW Thermal-Hydraulics. Applications for Trojan, March 1990.

l 2-9 BAW-10168P Revision 02, BWNT Loss-of-Coolant Accident Evaluation Model for Recirculating Steam Generator Plants, October 1992.

BAW-10168P Revision 03, BWNT Loss-of-Coolant Accident Evaluation Model for Recirculating Steam Generator j Plants, November 1993.

2-10 BAW-10169P-A, RSG Plant Safety Analysis, October 1989.

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Non Proprietary

/N 3.0 Resident Fuel Assembly Compatibility

\ s_s 3.1 Fuel Assembly Description and Component Measurements The Mark-BW fuel assembly is a 17x17, standard lattice fuel assembly designed for use in Westinghouse-designed reactors.

The Mark-BW fuel assembly, shown in Figure 3-1, consists of 264 fuel rods, 24 guide thimbles, and one instrument sheath in a 17x17 square array. The guide thimbles are annealed Zircaloy-4 and provide guidance for control rod insertion. The fuel assemblf contains 8 spacer grid assemblies, 6 Zircaloy-4 intermediate spacer grids, and 2 Inconel-718 end grids. The bottom nozzle is a proven debris resistant design. A detailed description of the Mark-BW fuel assembly may be seen in topical report BAW-10172P (Reference 3-1).

l The Mark-BW fuel assembly design for the Sequoyah Nuclear Plant l (SON) reload incorporates three advanced design features which I differ from those outlined in BAW-10172P. Each of these improvements is discussed below along with an explanation for its need.

1

[ ]

n kv /

(g i

t 3-1

1 than Propurietary The basic design parameters of the Mark-BW are comparable to those of the Westinghouse Standard, OFA, and VANTAGE SH 17x17 fuel assemblies. The Mark-BW fuel assembly incorporates proven design features while naintaining compatibility with the Westinghouse reactor internals and resident fuel assemblies.

[ ]

Mechanical compatibility with SQN core internals and all resident fuel assemblies was confirmed by direct comparisons with measurements taken from two (2) Westinghouse VANTAGE SH fuel assemblies. All accessible components of the fuel assembly were measured and recorded. The comparison between the two assemblies indicated that the VANTAGE SH and the Mark-BW are very similar dimensionally with only a few exceptions. The values noted below are primarily due to differences in design philosophy and the methodology used in component connection. A comparison of the major interface dimensions may be seen in both Tables 3-1 and 8-3. The following is a summary of the most significant design differences:

I l O

1 3-2 0

Non Proprietary l

f 3.2 Fuel Assembly Structural Testing A

s/ A comprehensive test program, described in detail in BAW-10172P (Reference 3-1) was conducted to characterize and verify the performance of the Mark-BW fuel assembly design. A complete list of the tests conducted is provided in Table 3-2. In addition, mechanical and pressure drop tests as described in the following sections were performed on Westinghouse VANTAGE SH fuel i assemblies. In the mechanical tests, static load deflection and pluck tests were performed to identify lateral stiffness, natural frequency and damping as a function of mid-span deflection. The results of these tests are used to benchmark fuel assembly analytical models.

As a part of the mechanical test program, tests were also conducted to determine load versus deflection behavior of the VANTAGE SH fuel assembly holddown springs at room temperature.

The results of these tests are used to compare the holddown capabilities of the Westinghouse VANTAGE SH fuel assembly and Mark-BW fuel assembly and serve as a basis to characterize the loads exerted by the fuel assembly on reactor internals.

3.2.1 Fuel Assembly Lateral Stiffness and Natural Frequency Tests A fuel assembly natural frequency test was performed on a

[-~sj Westinghouse VANTAGE SH fuel assembly in air at room temperature

( _,/ to obtain fundamental natural frequency and damping values at l various amplitudes. The test is conducted by measuring and recording the displacements of selected spacer grids as the fuel I assembly is deflected at the midplane and quickly released. l The fuel assembly tests were performed with initial deflections of [ ] inch. A minimum of three l

tests were performed for each test setup. Natural frequencies l

and logarithmic decrement damping were calculated from the '

amplitude decay plots. The natural frequency was computed relative to the amplitude as follows:

Fn =

1/ [Tn.1-Tn ]

l Where: Fn = frequency of nth cycle, Hz Tn = time of the nth peak, sec I T n.1 = time of the first peak of the same sign following the nth peak, sec The fuel assembly damping is calculated assuming a single degree of freedom system with viscous damping based on the log decrement method. The percent critical damping is calculated by:

\

i D! 3-3

Non Proprietary Dn = 100/(2*H)

• In [An/ ( A .1) n }

Where: Dn = % critical damping of the nth cycle An = amplitude of nth cycle A n.1 = amplitude of the same sign of the first cycle following the nth peak cycle Force versus deflection tests were conducted to determine the lateral stiffness of the fuel assembly. Fuel assembly lateral stiffness was tested by imposing midspan deflections of

[ ] inch. At least two tests were run for each test setup. The lateral stiffness was computed by taking the average load at each of the peak deflections and dividing by the total deflection.

The results of these tests are used in comparison with similar results obtained from the Mark-BW test program. Table 3-3 summarizes the test results.

The mechanical behavior of the Westinghouse fuel assemblies is consistent with the behavior exhibited by Mark-BW fuel assemblies. [ ]

3.2.2 VANTAGE SH Fuel Assembly Holddown Sprina Test The set of four holddown springs on the fuel assembly was loaded to a maximum load of [ ]lbs. and then unloaded. During the loading and unloading cycles, load steps of [ 11bs. were used for the first [ ]lbs. of load and load steps of [] lbs. were used thereafter. At each load step, holddown spring deflections were measured at all four sides of the test fixture. Each set of test data was then reduced to obtain an average load versus deflection data. This data was plotted to obtain the load / deflection curve for the spring system.

The VANTAGE SH fuel assembly holddown spring is preloaded to about [ ]lbs. After this load is applied, the spring shows a spring rate of approximately [ ] lbs./in. Holddown spring tests performed previously on Standard Westinghouse 17x17 fuel assemblies show [ ]when compared to the VANTAGE SH fuel assembly. The Mark-BW holddown spring has a spring rate of approximately [ ] lbs/in. Thus for the same holddown spring deflection, the Mark-BW fuel assembly design will have a [ ]

holddown force than that of the Westinghouse VANTAGE SH and Standard fuel assembly designs.

3-4

T Non Proprietary l

l /"'N The VANTAGE SH fuel assembly pressure drop results show that the

( ) total unrecoverable pressure drop across the VANTAGE SH fuel

\m/ assembly is [ ] than that of the Mark-BW fuel assembly. )

Hence, the VANTAGE 5H fuel assembly pressure drop and holddown i spring stiffness tests confirm the conservative design of the l Mark-BW fuel assembly holddown system.

3.3 Hydraulic Flow Testing A series of flow tests was performed, using c transportable flow test rig (TFTR), to verify the compatibility Letween the Mark-BW fuel design and the Westinghouse VANTAGE SH resident design. The TFTR is a " cold flow" loop, with typical test conditions encompassing water flow rates of [ ] gpm at pressures from [ ] psig and temperatures from [ ).

Pressure drop testing was performed on individual fuel assemblies, providing pressure drop characteristics for the VANTAGE SH fuel design, as well as for the Mark-BW.

The TFTR was validated as a viable source for determining fuel assembly hydraulic characteristics in 1988, when, as part of the effort to license Mark-BW fuel for Duke Power Company's McGuire and Catawba Plants and for Portland General Electric Company's Trojan Plant, a preliminary set of flow tests was performed to l

,-~s qualify the TFTR and establish a benchmark for comparison to l l \ later tests. These tests were performed with the same Mark-BW l

( _,

) prototype assembly that had been previously tested at the B&W Alliance Research Center, Alliance, Ohio, in the Control Rod l

Drive Line (CRDL) flow loop. The CRDL test spanned a wide range of pressure, temperature, and flow conditions encompassing those that occur during reactor operation. Core thermal-hydraulic analysis models for the Mark-BW fuel assembly design, including form loss coefficients for individual components, were developed from the CRDL test results. The measurements obtained with the TFTR demonstrated that the TFTR reproduces the CRDL results at comparable Reynolds numbers.

For the current test program, the goal was to first verify loop performance and then to hydraulicly characterize the VANTAGE SH fuel through a series of lift and pressure drop tests. [ ]

[

p

(

3-5

Non Proprietary

[ ]

Using the measured pressure drops, form loss coefficients for the fuel assembly subcomponents were determined. These form loss coefficients were then incorporated into a LYNXT hydraulic model which showed that the total pressure drop of the Westinghouse VANTAGE SH design is approximately [ ] than that of the Mark-BW. A plot of unrecoverable pressure drop versus axial elevation for the two designs at typical in-reactor conditions is provided on Figure 3-2.

For SON, TVA plans to use some Westinghouse Standard fuel assemblies as reinserts. As re, ported in BAW-10178P (Reference 3-2), a series of TFTR tests were also run to characterize the i Westinghouse Standard fuel design. The results from those tests I indicated that the Westinghouse Standard fuel assembly is approximately [ lin pressure drop than the current Mark-BW.

3.4 Mark-BW Nuclear Design Evaluation l

The Mark-BW is neutronically similar in design to both the Westinghouse Standard 17x17 and Westinghouse VANTAGE 5H fuel assemblies. Both the Mark-BW and the VANTAGE SH use Zircaloy spacer grids in the active fuel region, while the Standard uses Inconel grids. The fuel pin pitch and fuel pin outside diameter for all three designs are identical, and uranium loadings of all three designs are very similar. Therefore, reactivity depletion characteristics and moderator and Doppler coefficients are essentially the same for all three designs. From a physics standpoint (Reference 3-3), the Mark-BW assembly design represents only a small change from previous Westinghouse assembly designs already licensed and operated in SQN and will not adversely affect plant operation or neutronic parameters, l either alone or in combination with the Westinghouse designs.

3.5 References 3-1 BAW-10172P, Mark-BW Mechanical Design Report, July 1988.

3-2 BAW-10178P, Mark-BW Thermal-Hydraulics Applications for Trojan, March 1990.

3-3 BAW-10163P-A, Core Operating Limit Methodology for Westinghouse-Designed PNRs, June 1990.

3-6

,_._____.m_________.._.__ - - - - - - - - - - - - -- - - -

I Non Proprietary Table 3-1 1 1

I Comparison of the Mark-BW and VANTAGE SH-Design Mark-BW VANTAGE SH Fuel Assembly Fuel Assembly 1 Parameter Design Design

[ ]-

)

1 l

)

1 4

! 3-7

Non Proprietary Table 3-2 Mark-BW Prototype Test Program Hydraulic Flow Testing Fuel Assembly Lift Pressure Drop with Thimble Plug First 500 Hour Life and Wear Pressure Drop with thimble Plug Second 500 Hour Life and Wear Spacer Grid LDV Critical Heat Flux RCCA Drag

! Pressure Drop with RCCA RCCA Trip Fuel Assembly Structural Frequency and Damping in Air Lead Assembly Frequency and Damping in Air Axial, Lateral, Torsional, and Cantilever Stiffness Fuel Assembly Vertical Drop Component Structural Spacer Grid Impact Spacer Grid Crush Spacer Grid Slip Holddown Spring Force / Deflection Guide Thimble Buckling Joint Strength Control Rod Insertion  !

3-8 Oll l

l l

Non Proprietary Table 3 Summary of Fuel Assembly Test Results Mark-BW Characteristic VANTAGE SH Prototvoe

[ l t

n D 3-9 j 1

_.4m_a.A.-AtM -*-ar.d-aA.aJ-_ .*.d a am Aed 4.4 e4-4--6 --+M--- *--4am4- J4444J"AJ~a -M ^^8--h--** a4 4m --Pm* **m-m+Ndd'dh'-m4 -hsm-ad-64A------

Non Proprietary Figure 3-1 Mark-BW Fuel Assembly O

3-10

1 Non Proprietary Figure 3-2 Mark-BW and VANTAGE SH Unrecoverable Pressure Drop Comparison

]

k 1

l 2

l 1

i l

1 i

l l

l i

1 s

i.

t 3-11 i

l 1

1

f I

Non Proprietary i

7'"Ng 4.0 Plant Description t i V

The Sequoyah Nuclear Plant (SON) uses nuclear steam supply l systems designed by Westinghouse that are representative of the

! standard Westinghouse four-loop, 3411 MWt design except for I certain internal structures provided for the upper head injection system. The emergency core cooling system (ECCS) provided for the plant consists of the conventional combination of high pressure pumped injection, pressurized water storage tanks, and l low pressure pumped injection all connected into the reactor l coolant system (RCS) piping just upstream of the reactor vessel.

l SON uses an ice condenser containment.

I 4.1 Physical DescriptiQD The RCS is enclosed entirely within a containment and is 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 four pipes (hot legs), cooled within the steam generator tubes, and returned to the reactor vessel through four cold leg pipes. Flow through the system is driven by four reactor coolant pumps, one per coolant loop. System pressure is maintained by a pressurizer es connected to one of the coolant loops in the hot leg.

Reactor Vessel The reactor vessel is basically a cylindrical shell with a hemispherical bottom head and a removable hemispherical upper head. Major regions of the reactor vessel are the inlet and outlet 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 downcomer to the lower plenum. From the lower plenum, coolant is directed upward passing through the core to the upper plenum.

Within the upper plenum, the coolant mixes with a small amount of flow that was bypassed directly from the downcomer to the upper head and exits the reactor vessel through the hot leg nozzles.

The baffle region at SQN is not connected in parallel with the core. Flow in the baffle region is directed down from an upper elevation of the downcomer to the reactor vessel lower plenum.

Thus, SON is still of the original downflow design typical of Westinghouse plants of similar vintage.

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 N' 4-1

l l

Non Proprietary l l

one instrument tube. Each fuel rod consists of stacked fuel  !

pellets contained in a Zircaloy-4 fuel rod with a gap between the .

fuel pellet and the fuel rod. Fifty three of the fuel assemblies l have control rod cluster assemblies used for power control and shutdown capability. SON has silver-indium-cadmium control rods. i SQN will be replacing the Westinghouse VANTAGE SH fuel assembly ]

with the Mark-BW fuel assembly. Both the Mark-BW and the VANTAGE ;

SH are 17 x 17 fuel rod arrays with an active length of j approximately 12 feet. A comparison of fuel rod geometries for I both fuel types is provided in Section 5.10. l Reactor Coolant Loops 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 horizontal plane and undergoes an upward bend as it attaches to the steam generator inlet plenum. The steam 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 outlet plenum and continues downward for about 10 feet. At this point, the piping is bent through a 180 turn to vertical and rises to meet the reactor coolant pump casing. Discharge from the reactor coolant pump iu 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 SQN steam generators are of the recirculating or U-tube design; Westinghouse model 51 steam generators. The operation of the steam generators for both units is the same. Feedwater enters the secondary side of the steam generator through a header outside the tube shroud. The feedwater is mixed with saturated fluid that has been recirculated from the steam generator moisture separators. The feedwater/ recirculation fluid mixture flows down the annulus formed by the shell and the shroud and enters the tube nest at the top of the tube sheet. The fluid then travels up the tube nest where it is heated to boiling by heat transfer from the primary fluid inside the tubes. The two-phase mixture from the tube nest then enters the separators wherein the steam is allowed to proceed to the steam generator upper dome and steamlines; the liquid is recirculated to the annulus.

4.2 Description of Emergency Core Cooling System The ECCS provided for the SON units consists of the conventional combination of high pressure pumped injection, pressurized water 4-2

Non Proprietary l

l ye~'s storage tanks, and low pressure pumped injection all connected into the reactor coolant piping just upstream of the reactor i (

l \_ /

) vessel. Originally, the ECCS also included a direct upper head l injection system. In 1989, the upper head injection system was l shown to be unnecessary for accident mitigation and an i encumbrance to plant maintenance that adversely affected plant l operability. It has, therefore, been removed from the active plant systems and is no longer considered in plant licensing.

l The high pressure injection capability of the plants is achieved through two systems: the Centrifugal Charging (CC) System and the I

Safety Injection (SI) System. The CC System is the high pressure portion of the ECCS, capable of injecting above normal operating system pressure, and is part of the Chemical and Volume Control 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 operation is activated automatically after receiving an SI signal, indicating low RCS pressure, low steam line pressure, or high containment pressure. The SI System operates in the middle pressure range, capable of injecting up to about 1300 psia. It has two separate pumping sources with sufficient redundancy in the number of components to provide the required flow rate assuming a single active failure. The system is also actuated by an SI signal.

,,- s Low pressure injection is achieved with the Residual Heat Removal

(

\ (RHR) System. Normally used for cooling when the reactor is not

\ 'j operating, the system also serves the low pressure ECCS injection function by providing borated water through four cold leg injection lines. 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 from the containment sump. During recirculation, the injection flow is passed through a heat exchanger before being returned to the RCS. The system contains sufficient redundancy such that one full train is available under a single active failure. Actuation is by an SI signal on low RCS pressure, low steam line pressure, or high containment pressure. In its recirculation mode, the RHR System provides for long-term core cooling.

In the recirculation mode, only the RHR pump is capable of taking suction from the containment sump. However, long-term, high pressure cooling is possible because both the CC pumps and the SI pumps can take suction from the RHR pump discharge and deliver coolant through their cold leg connections to the RCS.

The accumulator system consists of four tanks, each containing about a thousand cubic feet of borated water and four hundred

, cubic feet of nitrogen pressurized to about 640 psig. The tanks

! are connected to the RCS at the reactor coolant pump discharge I via pipes that are common to the accumulator, SI, and RHR l

i (A)

C 4-3

Non Proprietary systems. 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 reactor coolant system occurs whenever the reactor coolant system pressure falls below the tank pressure.

4.3 Plant Parameters The major plant parameters and operating conditions are presented in Table 4-1.

Table 4-1 Plant Parameters and Operating Conditions Reactor Power 3411 MWt Operating Pressure 2235 psig Highest Allowable Total Peaking (Fg) 2.5 System Flow 348 x 103 gpm

^

Core Heat Transfer Area 59,870 ft 2, Average Linear Heat Rate 5.43 Kw/ft*

Fuel Assembly Mark-BW, 17 x 17 array Fuel Pin OD 0.374" Hot Leg Temperature 609 F Average Loop Coolant Temperature 578.2 F Steam Generator Pressure 842 psig

• Calculated using the nominal active fuel stack length of 144 inches.

4-4

Non Prorncietary

/N 5.0 LOCA Analysis

/ \ -

In accordance with the requirements of 10CFR50.46 and 10CFR50, i Appendix K, an evaluation of the Emergency Core Cooling System (ECCS) performance has been performed for Framatome Cogema Fuels (FCF) reload fuel for the Sequoyah Nuclear Plant (SQN). In presenting that analysis, this section of the reload topical complements other topical reports that describe the Mark-BW fuel design; the mechanical, nuclear, and thermal-hydraulics methods supporting the design; and ECCS codes and methods. The analyses and evaluations presented in this section are intended to serve, in conjunction with these other topical reports, as a reference for future reload safety evaluations for SQN applicable to cores with FCF-supplied fuel assemblies.

SQN uses nuclear steam supply systems designed by Westinghouse j that are representative of the standard Westinghouse four-loop, 3411 MWt design, with the exception of certain internal structures. The ECCS provided for the plant consists of the conventional combination of high pressure pumped injection, l pressurized water storage tanks, and low pressure pumped injection, all connected into the reactor coolant piping just J upstream of the reactor vessel. SON containments are of the ice l

condenser type; thus, post loss of coolant accident (LOCA)

,s containment pressures are comparatively low.

, [

l \

l

\ _ ,/ The results of calculated predictions of LOCAs must meet the criteria imposed by 10CFR50.46. At the time of initial i operation, SQN was fueled with Westinghouse-supplied fuel and compliance was demonstrated by calculations performed by Westinghouse and TVA. This section of the topical report documents compliance to 10CFR50.46 when the plants are fueled by FCF-supplied fuel. In this report, possible LOCAs are divided i into two groups depending on the assumed break size. For breaks larger than 1.0 ft , compliance is demonstrated by calculations and analyses performed in accordance with Volume I of the B&W Nuclear Technologies Recirculating Steam Generator LOCA Evaluation Methodology (BWNT RSG LOCA EM) BAW-10168, Reference 5-1 and the BWNT RSG LOCA EM update of Reference 5-2. For breaks smaller than 1.0 ft 2 , compliance is demonstrated by calculations and analyses performed in accordance with Volume II of the BWNT l RSG LOCA EM.

~

Section 5.1 presents the computer codes utilized in the SQN reload LOCA analyses and a brief description of the calculational l methods. The analysis parameters used for the large break  !

I calculations are discussed in Section 5.2. Generic sensitivity studies are addressed in Section 5.3. Large break plant-specific sensitivity studies and spectrum analyses to determine the most limiting break configuration are documented in Section 5.4.

I w 5-1

Non Proprietary Section 5.5 presents the LOCA limit calculations that confirm adherence to the first two criteria of 10CFR50.46. The evaluation of maximum hydrogen generation, coolable geometry, and long-term cooling are presented in Sections 5.6, 5.7, and 5.8, respectively. SQN small break LOCA reload analyses are presented in Section 5.9. Section 5.10 addresses transitions cores. A summary of the results of the analyses is presented in Section 5.11. LOCA analysis references utilized in Section 5.0 of this report are included as Section 5.12.

5.1 Computer Codes and Methods For the evaluation of large break cladding temperature transients and local oxidation, the BWNT RSG LOCA EM consists of several computer codes. Figure 5.1-1 illustrates the interrelation of the computer codes used for the large break analyses. The RELAPS/ MOD 2-B&W code (Reference 5-3), hereafter referred to as RELAP5, calculates system thermal-hydraulics, core power generation, and core thermal response during blowdown. The thermal-hydraulic transient calculations are continued with the REFLOD3B code (Reference 5-4) to determine refill time and core reflooding rates for the remainder of the transient. BEACH (Reference 5-5) calculations are restarted directly from the RELAP5 run and is used to determine the core thermal response during reflood with core flooding rates from the REFLOD3B outputs.

O 5-2

Non Proprietary l{

t

\

Figure 5.1-1 RSG Large Break LOCA EM Computer Code Interface

[ Initial RC System and

( Core Parameters RELAP5/ MOD 2 - B&W i

i 0 < Time < EOB Core Stored Energy Containment System Mass and Energy Minimum Acc Mass and Energy Backpressure SG Mass and Energy REFLOD3B l Pumped ECC Flows EOB < Time < EOE Flooding Rates Core l Parameters At EOB )

l BEACH EOB < Time < EOE Metal Water Reaction i

Hot and Average Pin Thermal Response

{ Peak Cladding Temperature O 5-3

Non Proprietary 5.2 Inputs and Assumotions The major plant operating parameters used in the LOCA codes are:

1. Power Level - The plant is assumed to be operating in steady-state at 3479 MWt (102% of 3411 MWt).

2. Total System Flow - The initial Reactor Coolant System (RCS) flow is 348,000 gpm.

3. Fuel Parameters - The initial fuel pin parameters are taken from TACO 3 (Reference 5-6) runs performed at the fuel assembly burnup that produces the highest peak cladding temperature (PCT). Sensitivity studies discussed in Section 5.3.2 show that fuel conditions at the beginning of life (BOL) are the most severe for large break LOCA.  !

i

4. Emergency Core Cooling System (ECCS) - The ECCS flows  !

are based on the worst case between the assumption of a !

single active failure and the assumption of no failure, i Sensitivity studies discussed in Section 5.4.2 show that the condition of minimum ECCS is the most severe assumption.

l

5. Total Peaking Factor (F g ) - The maximum total peaking factor ausumed by this analysis is 2.5.

6. The moderator density reactivity coefficient is based i on BOL conditions to minimize negative reactivity.

7. The cladding rupture model is based on NUREG-0630, 5.2.1 RELAPS Modeling The RELAPS computer code is used to analyze RCS thermal-hydraulic behavior and core thermal response during the blowdown phase of a LOCA. RELAP5 permits the user to select a model representation that results in a suitable finite difference model for the fluid system being analyzed. The large break LOCA nodalization for the plant evaluation, shown in Figures 5.2-1 and 5.2-2, was developed in accordance with the BWNT RSG LOCA EM (References 5-1 and 5-2).

The control volume inputs generally consist of volume geometry (area and height), flow-related parameters (resistance, hydraulic diameter, surface roughness, etc.), 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 selected. Flow paths are defined between control volume 5-4

l Non Proprietary

/~'N geometric centers. RELAPS is run in steady state to assure proper initialization.

Core As shown on Figure 5.2-1, the reactor core model consists of a l hot channel (nodes 326-345), an average channel (nodes 426-445),

and a core bypass region (PIPE component 346). The hot channel simulates one assembly and the remaining 192 assemblies are modeled in the average channel. The fuel assemblies are axially i divided into 20 segments with variable nodal length such that I each grid is located at the bottom of a node and three nodes are used to cover a grid span. Radially, the fuel pellet is divided into 7 equally spaced mesh points and two equally spaced mesh points for cladding. With the exception of crossflow between the hot and average channel, core nodalization is equivalent to the BEACH model nodalization.

l Crossflow is allowed between the average core and the hot l assembly using crossflow junctions. The resistance for these l l

junctions is developed from the experimental correlation given in Section 2.2.7 of Reference 5-7. For the sensitivity studies and the break spectrum, the axial power distribution in the core is based on a symmetric chopped cosine with a total peak of 2.5.

For the LOCA Limits studies, the position of the axial peak gg varies with the case (refer to Section 5.5). Initial fuel ll ) temperature, fuel pin internal pressure, gas composition, and l ( ,/ dimensions are based on BOL TACO 3 calculations.

Core heat generation in the LOCA transient consists of combined l

fission power as calculated by the RELAPS point kinetics model and fission product decay heat. The action of control or safety rods is not credited by the BWNT RSG (large break) LOCA EM. The kinetics model, however, accounts for changes in reactivity due l

to the transient response of fuel temperature and coolant density. The details associated with kinetics inputs are described in Appendix C of Reference 5-8, beginning on page C.3.

j In short, Doppler temperature and moderator density feedback used t

in the LOCA analysis have been selected individually as bounding

, for the burnup at which the impact of reactivity on the LOCA l calculations produces the highest PCT. Reactivity feedback components are combined by flux-squared weighting and spatially integrated to provide a single, effective feedback for the point kinetics model.

Reactor Vessel l l The reactor vessel model consists of a downcomer (nodes 300-308),

the baffle region (PIPE component 350), the lower head (node ,

310), the core inlet plenum (nodes 312 and 314), the core outlet  !

plenum (nodes 352 and 354), the flow and support columns (node 4 i

\s 5-5

Non Proprietary 356), and the upper head (nodes 358 to 364). The core inlet is centered at the interface between downcomer nodes 300 and 302.

In the upper plenum, the vessel outlet is centered at the interface between nodes 352 and 354. Reactor coolant hypass between the downcomer and upper head is taken into account by mnnecting node 300 to node 360. Similar to the Reference 5-8

• cdel, the UHI flow columns and control rod guide tubes are modeled with a separate volume (node 356), and the upper plenum is divided into two volumes (nodes 352 and 354).

Reactor Coolant Loops The loop noding scheme is a result of the loop noding and break noding sensitivity studies performed in Appendix A, Volume I of the BWNT RSG LOCA EM (Reference 5-1). The four RCS loops are modeled as two by combining the three loops which will not contain the break as shown in Figure 5.2-2. The 100 series nodes model the unbroken loops, and the 200 series models the broken loop. Within each loop, the hot legs are separated into 4 nodes; the RSG inlet plenum (nodes 120 and 220) and RSG outlet plenum (nodes 130 and 230) are single nodes. The RSG tubes (PIPE components 125 and 225) are separated into sixteen segments. The tube flow area is based on the assumption that 15% of the tubes have 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 (nodes 160 and 260) is a single node. The 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.

Reactor Coolant Pumpa The reactor coolant pump performance is developed from homologous relationships adjusted for two-phase degradation based on the data in Table 2.1.5-2 of Reference 5-3. These are the same degradation data presented in NUREG/CR-4312 (Reference 5-9) . In accordance with the BWNT RSG (large break) LOCA EM , the reactor coolant pumps are assumed to trip at the time of break initiation.

Pressurizer The pressurizer model consists of three parts: the surge line (node 400), an eight-section pressurizer (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.

5-6

_ _ _ . - - ._- _. . . - . . .._~---.- . . . - _. __ -_- - - . _ _ ,

Non Proprietary l

Recirculating Steam Generator O In agreement with the loop noding arrangement, the three steam generators associated with the unbroken loops are modeled as a combined region (700 node series) with the broken loop containing a single generator (600 node series). The four-volume tube riser section (nodes 630 and 730) spans the height of the generator tubes. Two volumes (nodes 635 and 640, and 735 and 740) are modeled below the separators. Nodes 650 and 750 are separator ,

volumes. The steam dome is modeled by two nodes (660 and 670, and 760 and 770). The separator component in RELAP5 acts as a steam separator and dryer with two-phase fluid entering from the bottom, steam exiting upward, and saturated fluid going back to the downcomer through nodes 655 and 755. Nodes 625 through 665 and 725 through 765 form the steam generator downcomer. Main and auxiliary feedwater is supplied at nodes 620 and 720. Nodes 675 to 690 and 775 to 790 provide simulation of the steam lines and the safety valves.

The heat structures that represent the steam generator tubes are reduced in heat transfer area 3n accordance with the 15% tube  :

plugging assumption. Heat stru _mres of characteristic volume  :

and surface area are also includeu for the shell, the downcomer  :

walls, and separator components. i e-s Break Characteristics

\ The four-node break configuration is a result of the break noding sensitivity study in Appendix A of Volume I of the BdNT RSG LOCA EM (Reference 5-1). Figure 5.2-2 illustrates that a double-ended cold leg guillotine (DECLG) break is modeled with independent leak paths from nodes 270 and 275 to the containment. For a split-type break, the leak is modeled as a single path that leaves node 275. For the DECLG 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%.

Primary Metal Heat Model 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 system pressure boundaries are assumed insulated to maximize stored energy in metal slabs.  ;

Emercency Core Coolant System  ;

The accumulators are modeled by node 900. Injection is allowed only into the unbroken loops at the cold leg. The injection k'

5-7

i 1

Non Proprietary 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. Accumulator fluid, injected into the intact cold legs, is bypassed to the containment by the RELAPS model. The end of bypass is, however, determined by the model and generally occurs just prior to end of blowdown (EOB). RELAPS tracks accumulator fluid injected after the end of bypass and it is transferred to the lower plenum node of the REFLOD3B model. The procedure for the introduction of accumulator fluid to the vessel is described in Volume 1 of the BWNT RSG LOCA EM, Reference 5-1, beginning on page LA-215.

The pumped injection systems, although not explicitly modeled in the RELAP5 model are activated by the Safety Injection (SI)

Systems signal of the Engineered Safety Features Actuation System; the signal occurs during blowdown. Pumped injection is simulated in the REFLOD3B model, using the SI signal time generated by RELAP5 during blowdown. Conservative time delays for signal generation, electrical supply startup, and injection pump startup are used. Processes accounted for in delaying pumped injection are listed in Reference 5-8, beginning on page C.10.

5.2.2 REFLOD3B Modeling The REFLOD3B code, Reference 5-4, simulates the thermal-hydraulic behavior of the primary system during the core refill and reflood phases of the LOCA. The noding arrangement, shown in Figure 5.2-3, is consistent with the BWNT RSG (large break) LOCA EM (References 5-1 and 5-2) and consists of reactor vessel and loop models. The reactor vessel is represented by a four fixed-node model; nodes 1R and 2R are volumes above and below the steam-water interface in the inner vessel region, and nodes 3R and 4R represent liquid and steam volumes, respectively, in the downcomer region including the lower plenum. The primary system piping is represented by two loops similar to the RELAPS blowdown model with a reduced number of volumes. Values for volume geometry and flow path hydraulic parameters are developed from the RELAPS model.

RELAP5 results at the EOB define the starting point for the REFLOD3B calculations. The core and system initial conditions for REFLOD3B are derived from those of the associated RELAP5 run at EOB, taking into account the differing nodalization. Initial liquid inventory for REFLOD3B is equivalent to the liquid remaining in the reactor vessel at EOB added to the accumulator liquid bypassed to containment following the end of bypass in RELAPS. Liquid is placed in the lower plenum of the REFLOD3B 5-8

Non Proprietary model. The initial flow rates, gas volumes, liquid inventories, fe~%si g and pressures of the accumulator tanks are taken directly from

\s_ / RELAP5. The reactor vessel 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 side.

The primary metal heat structures are also taken from the RELAPS model. Initial conditions in REFLOD3B match those in RELAP5 at EOB. The secondary side metal structures include a representation of the shell material. In mostly stagnant regions, such as the downcomer or lower head, the heat transfer coefficient is based on pool boiling or natural convection to vspor. 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 ensure that the fluid leaving the steam generator is continuously dry steam, superheated to the secondary side temperature. Core thermal properties, such as fuel, gap, and gap thermal conductivities, are chosen to maximize heat transfer to the coolant, effectively reducing core flooding rates.

For a DECLG break, two leak paths are modeled, one from the RV I upper downcomer (node 4R) and the other from the pump side of the break (node 28). The pump rotor resistance is based on the

[g locked rotor condition. A 0.85 psi pressure drop is imposed on cold leg pipe junctions to account for momentum losses due to steam-ECC water interaction during the accumulator injection phase. This value is reduced to 0.50 psi for the pumped injection once the accumulators have fully discharged.

The containment backpressure as a function of time, from the SON Final Safety Analysis Report (FSAR), is used in the REFLOD3B calculations. The FSAR containment pressure produces conservatism in the prediction of reflooding rates. This containment pressure is justified as acceptable in Section 5.4.5.

5.2.3 BEACH Modeling The BEACH code is used to determine fuel rod cladding temperature response during the refill and reflooding phase of a LOCA. The BEACH model consists of separate hot and average fuel assemblies  !

and respective flow channels. Time-dependent inlet and outlet volumes permit inputs of boundary data from the REFLOD3B calculations. No crossflow is permitted between the hot and average channel of the BEACH model. Otherwise, the subdivision of the BEACH model is equivalent to the RELAPS core model described in Section 5.2.1. The nodalization is equivalent to that introduced in the BWNT RSG LOCA EM (Reference 5-2).

,b 5-9

1 Non Proprietary In accordance with the BWNT RSG LOCA EM, Reference 5-2, the BEACH l run is restarted from RELAP5 at EOB. Fuel assembly thermal parameters are transferred directly between the codes. The initial temperature of the steam surrounding the fuel assemblies is set equal to the adjacent cladding surface temperature at the EOB. Inlet and outlet pressures, flooding rate, inlet water temperature, and core decay heat boundary data are input from the REFLOD3B calculations.

The automated rupture modeling features, introduced in Reference 5-5, are used in the BEACH run. The process of accounting for a cladding rupture remains unchanged from previous applications of the BWNT RSG LOCA EM with the except4an that ennual input '

revisions by the user are no longer .4ecessary. Rupture-related blockage form loss, convective enhancement inputs, and droplet  :

1 breakup parameters are calculated and inputs adjusted internally at the predicted clad rupture location.

O l

1 5-10

i i

)

i Non Proprietary l l

r Figure 5.2-1 RELAP5 Large Break LOCA Model Sequoyah Noding for Reactor Vessel and Core l

l l =

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+

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O 5-11

Non Proprietary Figure 5.2-2 RELAP5 Large Break LOCA Model Sequoyah Noding for Primary Loops with Model 51 SGs r-d 158o % '

y >; Q 8 :: g m l l m t i t I = 7e5 855 " l 4 4 I 750 8 ss5 = l I I 7a I u,w __

u,w $45 I I I M u l 73s l m l 4,o p l eoo sao l s>s l y t 4,o f j N 73tHp4 4,0 s 630-04 N

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( ,ln Proprietary V I Figure 5.2-3 REFLOD3B Large Break LOCA Model Sequoyah Nodalization STEAM GENERATOR 1 STEAM GENERATOR 2

• SECO 6 RY SECONOMW 7 tt 7 g 20 21

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/ 31 30 im INTACT LOOP TOP OF M BROKEN LOOP ,

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5-13

Non Proprietary l

Figure 5.2-4 RELAP5/ BEACH Large Break LOCA Model Sequoyah Nodalization l 1

Segment Elevation 3 WDPVOL 4 TMDPVOL No. Feet

• l 3 g 34541 445 01 )

20 11.64 34s 3  % 44s 19 10.92 y 344 k 444 I 18 10.28 343 Q 3 443 17 9.71 342 h [ 442  !

16 9.14 y 341 Q 443 l

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% @ 440 14 8.00

@ 339 y 439 13 7.43 338 3 k 438 l 12 6.86 337 437

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8

% { k 434 4.58 k 333 3 433 7 4.01 3 332 432 6 3.44 331 5

% k 431 2.87 330 43o Q 3 4 2.30 329 h f 429 3 1.68 3 328 k 428 2 1.01 k 327 h 427 1 0.34 326 h 3 426 o a mopa mopm i 319 419 1 TMDPVOL 2 TMDPVOL

• Elevation from the bottom of the active core stack l l

l 5-14 O

Non Proprietary

,C} 5.3 Sensitivity Studies t /

\_ / LOCA evaluations require that a substantial number of sensitivity studies be performed with the evaluation model in order to establish model convergence and conservatism. Studies such as break spectrum and worst case ECCS configuration are considered plant-specific and are documented in Section 5.4 of this report.

Most of the studies upon which the evaluations in this report are based, however, are generic and were documented in the BWNT RSG LOCA EM report, Reference 5-1. Many of the generic studies, potentially affected by a subsequent BWNT RSG LOCA EM update, were justified as applicable in Reference 5-2. The current section provides a discussion of the generic sensitivity studies from the BWNT RSG LOCA EM report that are applicable to SON.

5.3.1 Evaluation Model Generic Studies Of the sensitivity studies presented in the original BWNT RSG LOCA EM topical report (Reference 5-1), the majority are generic and would apply to any RSG plant evaluated. Those studies considered generic each demonstrate results that are characteristic of the BWNT RSG LOCA EM--the codes and interfaces-

-and that are not plant dependent. An example of this is the RELAPS time step study, which demonstrated that the automatic time step selection in RELAPS would produce converged results.

fx This demonstration need not be repeated for plant-specific

( \ applications provided the modeling techniques used adhere to the

\ '- l general guidelines of the BWNT RSG LOCA EM. The following is a listing of the sensitivity studies considered to be generic, with l a discussion of why the conclusions of the study remain applicable for this applications report.

RELAPS/ MOD 2-B&W Time Step Study I This study was performed in Volume I of the BWNT RSG LOCA EM, Section A.2.1 of Reference 5-1 and was subsequently juctified for use with the BWNT RSG LOCA EM update, page 10 of Reference 5-2.

The RELAP5 time step study verified that for light water reactor geometry, the RELAPS time step controller governs the code solution sufficiently to assure converged results. Deviations of the SON system design from those simulated by the BWNT RSG LOCA l EM are not sufficient to change this result. Therefore, the '

study remains applicable with respect to blowdown thermal-hydraulic and fuel heatup calculations with the RELAPS code.

l RELAPS/ MOD 2-B&W Loop Noding Study J This study was performed in Volume I of the BWNT RSG LOCA EM, Section A.2.2 of Reference 5-1. The RELAP5 loop noding study verified the general noding requirements within the loop for

[A) 5-15

l l

Non Proprietary 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. ,

Deviations of the SQN system design from that simulated in the l BWNT RSG LOCA EM are not sufficient to change the noding l l

requirements. Therefore, the results of this study remain j applicable. i

l REFLOD3B Primary Coolant Pump Rotor Resistance Study This study was performed in Volume I of the BWNT RSG LOCA EM, Section A.2.4 of Reference 5-1. The rotor resistance study I determined that the limiting reactor coolant pump configuration, with respect to reflooding rate, is associated with a locked rotor assumption. This result applies to all RSG plants that the BWNT RSG LOCA EM is applicable to. Therefore, the results of this study remain applicable to SON.

RELAP5/ MOD 2-B&W Break Noding Study This study was performed in Volume I of the BWNT RSG LOCA EM, Section A.3.1 of Reference 5-1 and was subsequently justified for use with the BWNT RSG LOCA EM update, page 11 of Reference 5-2.

This study verified that hydraulic stability is achieved by providing at least one control volume in the pipe between any adjacent component and the break node. The break noding study is applicable to all plants covered by the BWNT RSG LOCA EM. The results of this study, therefore, remain applicable.

RELAP5/ MOD 2-B&W Pressurizer Location Study This study was performed in Volume I of the BWNT RSG LOCA EM, Section A.3.2 of Reference 5-1 and was subsequently justified for use with the BWNT RSG LOCA EM update, page 11 of Reference 5-2.

Although the assumption placing the pressurizer in one of the ,

I intact loops was somewhat conservative, this study showed that there is little difference in results when the pressurizer is modeled in the broken loop. The lack -f sensitivity to pressurizer location is expected to hold for all designs covered by the BWNT RSG LOCA EM. Therefore, the study remains applicable and will not be repeated for SQN.

RELAPS/ MOD 2-B&W Core Crossflow Study This study was performed in Volume I of the BWNT RSG LOCA EM, Section A.3.4 of Reference 5-1 and was subsequently justified for use with the BWNT RSG LOCA EM update, page 12 of Reference 5-2.

The core crossflow study verified that cross flow in a light water reactor is limited and does not alter the course of a LOCA evaluation substantially. The study is dependent only on the very basic aspects of the fuel design, which are consistent 5-16

! l Non Proprietary i l

g across the range of designs considered by the evaluation model.

Therefore, the study remains applicable to SQN.

l RELAPS/ MOD 2-B&W Core Noding Study l In conjunction with the core crossflow study, this study, Volume I of the BWNT RSG LOCA EM, Section A.3.4 of Reference 5-1 (Revision 0), verified that the modeling of a light water reactor ,

core in six axial segments with a hot and an average channel  !

provides sufficient spatial detailing for both model convergence and result accuracy. A subsequent, approved, application of the BWNT RSG LOCA EM in Reference 5-10 introduced a twenty segment core nodalization as an improvement to the blowdown model within the scope of the BWNT RSG LOCA EM. This improvement is adapted l in the update to the BWNT RSG LOCA EM, Reference 5-2, and the I current application for SON. Since the deviation in core i nodalization from the original BWNT RSG LOCA EM is an improvement on the model and is within the scope of the BWNT RSG LOCA EM, the core noding study is considered applicable to SON.

l 5.3.2 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 circumstances. Originally addressed in the base BWNT RSG LOCA EM, the confirmable studies have been verified as being applicable, with certain restrictions, for each subsequent revision to the BWNT RSG LOCA EM. The following is a listing of such sensitivity studies with a discussion of why the conclusions of the study can be applied to SQN.

RETAP5/ MOD 2-B&W Pump Degradation Study I

A sensitivity study was performed in Volume I of the BWNT RSG LOCA EM, Section A.3.3 of Reference 5-1, to establish reactor coolant pump two-phase degradation modeling criterion. As a result of this study, pump degradation relations contained in )

l NUREG/CR-4312 were chosen as being most representative at higher i

void fractions than the alternatives. The study can be applied to all plants which experience similar LOCA core flow histories during blowdown.

Figure 5.4-3, in the plant-specific sensitivity studies section of this report, illustrates blowdown core mass flux for the SON base case. Figure 6-1 of Reference 5-8 illustrates blowdown mass flux responses for both the McGuire/ Catawba and the BWNT RSG LOCA

, EM base cases. A comparison of the figures indicates similar

! core flow histories. The results of the reactor coolant pump i degradation study is, therefore, equally applicable to SON.

h 5-17 1

l

Non Proprietary REFLOD3B Loop Noding Study This study, Volume I of the BWNT RSG LOCA EM, Section A.2.3 of Reference 5-1, 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 SON. A separate study is required only for severely altered loop designs, such as the B&W or Combustion Engineering 2-by-4 designs. Therefore, the study remains applicable.

Time-in-Life Study A time-in-life sensitivity study was performed in Volume I of the BWNT RSG LOCA EM, Section A.3.9 of Reference 5-1, and was subsequently justified for use with the BWNT RSG LOCA EM update, page 15 of Reference 5-2. The study describes, in some detail, factors to be considered in determining the limiting initial conditions for LOCA analysis within a given fuel cycle. The study concludes that, for fuel pin thermal conditions that do not produce a blowdown rupture, the most restrictive burnup for LOCA analysis is the burnup producing the highest initial fuel temperature. For conditions that lead to a cladding rupture during blowdown, a more detailed study is required to determine the most severe thermal conditions. This conclusion is independent of the fuel and plant design that is determined not to encounter a blowdown rupture. Both the conclusion of this study and the conditions associated with it are applicable to SON LOCA analysis.

[ ]

1 1

1

l Non Proprietary

/N [ ]

( l

\m / A RELAPS blowdown run was conducted that utilized a composite of the worst fuel pin thermal conditions predicted by the TACO 3 study to determine the likelihood of a blowdown rupture. [ ]

The composite blowdown case did not result in cladding rupture.

The studies for burnup in the referenced BWNT RSG LOCA EM report support a basis for the selection of the worst time in life to evaluate a LOCA given that a blowdown rupture cannot occur. A bounding blowdown analysis was performed that indicated that the occurrence of a blowdown rupture does not occur for the reload of Mark-BW fuel at SON, provided fuel peaking is maintained within the assumed limits. The BWNT RSG LOCA EM time-in-life study determined that BOL fuel conditions are the most conservative conditions to assume for the LOCA analysis. Thus, the LOCA-limiting time in life for the SON LOCA analysis is established as j

[ ). )

5.3.3 Break Location I

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 i considered generic. There are three break locations to consider 1 7-ss for large break LOCA: the hot leg piping, the cold leg piping at

[

i the pump suction, and the cold leg piping at the pump discharge--

('~'/ 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 Volume I, Appendix A, of Reference 5-1 concluded that the blowdown and reflood cooling for the pump suction break were ,

sufficient to dismiss the suction break as a potential limiting i case strictly on the basis of the blowdown and reflood system analyses.

An examination of the consequences of the break location explains these observations. With the break at the pump discharge, one accumulator and a portion of the available pumped injection will be bypassed directly to the containment, not providing blowdown

(N

\

V 5-19

Non Proprietary 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, however, is not as strong as for a hot leg break.

Therefore, in conjunction with the better blowdown cooling, the

, suction break produces a lower peak cladding temperature than does the pump discharge break. While the lower peak temperature alone precludes the suction break from being limiting, there is further reason to 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%), that condition represents a.more limiting ECCS flow than does the pump suction break. For these reasons, and based upon the discussions of the foregoing paragraphs, the spectrum and LOCA limits cases performed for SON are for large breaks in the reactor coolant pump discharge piping.

O 5-20

, _ _ . _ . . _ - ~ ~ . _ _ -

l 1

4 l Non Proprietary I

l_ Figure 5.3-1 Time-In-Life Study Burnup Limit

[ ]

i l

l l

Figure 5.3-2 Time-In-Life Study Max Average Fuel Temperature

[ ]

Figure 5.3-3 Time-In-Life Study Hot Pin Internal Pressure

[ ]

i l

I 5-21

l 1

Non Proprietary 5.4 Plant-Soecific Studies and Soectrum Analysis i

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 i studies. These studies and the spectrum analysis are performed l to identify a worst case break to use in calculating the loss of coolant accident (LOCA) limits. Figure 5.4-1 shows the order in l which the plant-specific sensitivity studies and the spectrum I analysis cases were performed. This chapter presents the results of the studies leading to the final configuration used in the LOCA Limits cases.

5.4.1 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 DECLG break, with a discharge coefficient (Cd) of 1.0, located between the reactor coolant pump and the reactor vessel. Full operation of the emergency core cooling system (ECCS) is assumed for the base case. All of the sensitivity studies utilize an axial power profile that is peaked i at mid-core. I Figures 5.4-2 through 5.4-9 present key parameters resulting from the base case analysis. In general, the results compare well to those documented in the most recent BWNT RSG LOCA EM report, Reference 5-2, for the same class of plant. Aside from differences in plant-specific inputs, the cases differ in that the BWNT RSG LOCA EM demonstration case was run with a total allowed peaking factor of 2.32 while the current case assumes a total peaking of 2.50.

5.4.2 Minimum Emergency Core Cooling System Analysis Prior to the break type or spectrum studies, a study is conducted to determine if an assumed single failure of the ECCS produces more severe cladding temperatures relative to the base case.

Under a single failure assumption, only one train of pumped ECCS ,

injection is available. With no failure, two full trains are I 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.

The calculations for the base case assumed no failure of the ECCS i or supporting systems. This is normally referred to as the " Max ECC" case. To evaluate the assumption of a failure, a calculation was performed with the condition that one of the diesel gens.rators of the emergency power supply failed to 5-22

Non Proprietary p\

operate, resulting in a loss of electrical power to half of the pumped ECCS. This is generally referred to as a " Min ECC" case.

\s_ / Table 5.4-1 presents a comparison of the results of the Min and Max ECCS analyses. Figures 5.4-10 through 5.4-16 present relevant parameters for the' Min ECC case. Comparable base case results are shown in Figures 5.4-2 through 5.4-9.

In past applications of the BWNT RSG LOCA EM the Max ECC case has resulted in higher cladding temperatures (References 5-8, 5-10).

This is a consequence of lower reflooding rates relative to the Min ECC case. The exceptionally high resistance to flow characteristic of the SQN accumulator surge line, however, causes Min ECCS cladding temperatures to be the most severe.

Peak clad temperatures (PCTs) occur in the period of time just j after the accumulators empty. During this period the effect of '

the difference between the Min and Max pumped ECCS flows on core reflood rate and, subsequently, the PCT is minimized by the comparatively high accumulator injection flow rate. Beginning of core recovery, and the end of fuel adiabatic heatup, occurs slightly sooner for the Max ECC case. The Min ECCS PCT is approximately 20*F higher as a result. The remainder of the evaluations will, therefore, proceed under the assumption of a 1

single failure of the ECCS system.  ;

s 5.4.3 Discharge Coefficient Study

\ The break spectrum analysis is performed to determine the worst case break size and the worst break configuration. Break location studies presented in Volume I, Section A.3.8, of Reference 5-1 and Section 6.3 of Reference 5-8 conclude that the hot leg and reactor coolant pump suction breaks are assured to be less limiting than the pump discharge breaks because of the lack of ECCS spillage for these events. Noting the similarities in the McGuire/ Catawba and SQN plant configurations, it is apparent that this conclusion applies to SON as well. The differences between the split and guillotine bro + and the range of break sizes are more difficult to generalize, however, and those studies were re-run for SON. The break size study was performed first, followed by the break type study.

For the BWNT RSG LOCA EM, the break size study is interchangeable with a discharge coefficient study since the break flow is directly proportional to the product of the break area and the discharge coefficient. The Min ECC case establishes the results of a DECLG break at the pump discharge with a Cd of 1.0. Cases characterizing Cds of 0.8 and 0.6 are conducted to complete the discharge coefficient study. Key parameters for these cases are shown grouped by case in Figures 5.4-12 through 5.4-30. Table 5.4-2 presents a comparison of the results of the discharge coefficient study.

(s- 5-23 i

i Non Proprietary l

There are no major differences among the sequences of < events for the three cases that make up the discharge coefficient study. As expected, the blowdown is extended as the break flow is decreased. Slightly more heat is removed from the fuel during blowdown in the 0.8 and 0.6 discharge coefficient cases, as indicated by the hot spot cladding temperatures at the end of blowdown (1229 F, 1131 F, and 1026 F for the cd = 1.0, 0.8, and 0.6 cases, respectively). All three cases leave somewhat over 60 cubic feet of liquid in the reactor vessel at the end of blowdown, and the adiabatic heatup periods during lower head refilling are nearly the same. The reflooding transients are similar for the three cases. Reflooding rates are increased slightly with lower discharge coefficients, however, and reflood cooling is improved as a result.

Since the discharge coefficient study cases show increasingly better blowdown and reflood cooling with reduced discharge, the peak cladding temperature predicted for the Cd = 1.0 case is the highest of the three cases. Therefore, a Cd = 1.0 was selected for the remaining sensitivity studies and as the limiting discharge coefficient.

5.4.4 Break Type Study Following the selection of the discharge coefficient, the type of break, split or guillotine, was considered. The DECLG break is modeled as a complete severance of the pipe, allowing separate discharges through the full area of the cold leg piping from both 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 the size of the cold leg piping cross section. Although the flows from the two sides, reactor coolant pump and reactor vessel, must still pass through limiting pipe areas, they are allowed to mix at the break location. The blowdown rates and system flow splits are somewhat different for the two types of breaks, and that can lead to differences in cooling response.

Figures 5.4-31 through 5.4-37 present the results of the split type break case. Comparable results for the limiting JECLG break case are shown in Figures 5.4-12 through 5.4-16. Both case types are double-area 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 5.4-2.

As indicated in Table 5.4-2, the period of blowdown for the split break is relatively short. Although the blowdown PCTs are higher for the split case, relative to the DECLG break, ultimately the blowdown cooling is better. Increased heat removal from the fuel during blowdown is indicated by the hot spot cladding temperatures at the end of blowdown (1155 F for the split case 5-24

l l

Non Proprietary 1

1 i /'~Ng and 1229 F for the DECLG break) . The end of blowdown liquid l( ) inventory in the reactor vessel lower head is nearly the same for ,

l \_ / the two break types; adiabatic heatup periods are nearly the same I as a result.

The limiting DECLG break case produces a peak cladding i temperature which is 1 F greater that the split case. This l result would imply that the sensitivity of a large break to break '

type is not very great. The peak cladding temperature for the l break type sensitivity cases occurs during a numerical spike, I

however, and bears some examination at this juncture.

Numerical spikes in cladding temperature are produced early in I the reflooding period associated with large break. The spikes coincide with the combined quench and liquid filling of ,

individual nodes in the bottom of the core. A node fills and the I boiling mechanics responsible for liquid entrainment and carry-out is momentarily interrupted. This phenomenon briefly interrupts the droplet cooling of upper-core nodes adjacent to grids and rupture locations. The effect is particularly pronounced at the rupture location because both the inner ar.d outer cladding surface at this node are subject to metal-water reaction.

In the split case clad temperatures are, in general, about 80 F f- s lower than those resulting from the DECLG break case. A spike in t i the cladding temperature at the rupture location, however,

('"') represents an abrupt reversal in this trend. The rupture location cladding temperature spike peaks at 1946 F for the split case compared to a peak of 1835 F for the guillotine break case.

The high rupture node temperature adds heat to the hot channel fluid and promotes a higher-than-expected cladding temperature in the adjacent (FCT location) node, pushing PCTs for the two cases to within 1 F of each other. The end result is an 80 F clad temperature penalty associated with the discretization of the core model.

The DECLG break case produces the highest general cladding temperatures, relative to the split case. The PCT for the guillotine break case exceeds that fo: the latter case by only 1 F. Based on these observations, the guillotine break is the most limiting break type.

5.4.5 Containment Pressure Study This study was performed to determine the sensitivity of the large LOCA analysis to containment pressure. Figure 5.4-4 illustrates the "bace" containment backpressure profile utilized l

in all of the above sensitivity studies. The base pressure profile is taken from the SON FSAR and is characteristic of the l

(

[\ 4 l \ /

, L' 5-25

Non Proprietary SQN worst case LOCA analysis results as determined by Westinghouse methodology.

Figure 5.4-38 shows the base containment backpressure profile in comparison with pressure curves that were generated for specific SQN large break cases. These case-specific backpressure predictions were made by TVA with the MONSTER computer code. By implementing the TVA/ MONSTER containment pressure into the large break analysis, direct feedback between the mass and energy release associated with a break and the containment response to the release is credited.

The limiting large break configuration is the DECLG break in the reactor coolant pump discharge piping, Cd=1.0, with minimum pumped ECCS injection. The MONSTER containment pressure related to these limiting parameters is used as a boundary condition in the current sensitivity study. The results of this study are tabulated in Table 5.4-3 and plots of important parameters for this study are presented as Figures 5.4-39 through 5.4-43.

Table 5.4-3 shows that the blowdown and refill characteristics for TVA/ MONSTER containir.ent pressure case differ little from the results of the base containment pressure case. Rupture timing and the physical characteristics of the rupture are the same for both cases. The PCT for the base containment pressure case is 2 F higher than the predicted PCT for the current case (2034 F vs. 2032 F) .

Sensitivity of the LOCA results to containment pressure is greater than might be concluded by a PCT comparison. A numerical spike in rupture node temperature is predominant for the current case. This non-physical result translates up the core and is, in effect, responsible for a PCT penalty of nearly 30 F. In general, the TVA/ MONSTER containment pressure case indicates cladding temperatures that are about 30 F lower than the case utilizing the base containment pressure as a boundary condition.

The use of the base (SQN FSAR) containment backpressure in the large break LOCA analysis produced overall cladding temperature responses that are higher than those predicted with a case-specific TVA/ MONSTER containment pressure. PCT for the former case is slightly higher. Use of the base containment backpressure, therefore, ensures conservatively high cladding temperatures and will be used as a boundary condition in the SQN LOCA limits analysis.

5.4.6 Gadolinia Puel Pin Study The SON reload cores with Mark-BW fuel will contain some fuel pins with less than 10 wt.% gadolinia in the UO 2 fuel pellets.

In the study, the U-235 enrichment in the gadolinia-bearing fuel 5-26

Non Proprietary D. pins was reduced from that of the nongadolinia-bearing fuel pins

( within the assembly. Then a SQN-specific time-in-life fuel performance analysis was performed with the GDTACO computer code (Reference 5-11) that defined the power history envelope for gadolinia-bearing fuel pins. Evaluations were subsequently performed that confirmed that the UO2 fuel pins at BOL are LOCA-limiting relative.to gadolinia-bearing fuel pins for any time frame within the fuel cycle. For cycle-specific application, the U-235 enrichment in the gadolinia-bearing fuel pins is further reduced to assure the gadolinia-bearing fuel pins remain nonlimiting.

G ,

\

l

(

5-27

l Non Proprietary Table 5.4-1 Min vs. Max ECC Flow Comparison l

Item or Parameter Max ECC Min ECC End of Blowdown, s 25.7 25.7 PCT at EOB, F 1229 1229 RV Liquid at EOB, ft' 74.1 74.1 Bottom-of-Core Recovery, s 45.1 45.5 Rupture, s 46.2 46.2 Rupture Elevation, ft 6.3 6.3 PCT, Ruptured Segment, F 1774 1835 PCT Elevation, ft 5.1 6.9 PCT, F 2014 2034 Table 5.4-2 Spectrum and Break Type Comparison Guillotines Split Item or Parameter Cd => M M M M End of Blowdown, s 25.7 26.7 31.3 24.5 PCT at EOB, F 1229 1131 1026 1155 RV Liquid at EOB, ft 3 74.1 65.3 65.0 60.8 Bottom-of-Core Recovery, s 45.5 47.5 50.0 44.6 Rupture, s 46.2 51.8 66.2 49.3 Rupture Elevation, ft 6.3 6.3 6.3 6.3 PCT, Ruptured Segment, F 1835 1846 1684 1946 PCT Elevation, ft 6.9 6.9 6.9 6.9 PCT, F 2034 1992 1865 2033 Table 5.4-3 Containment Backpressure Study Comparison Item or Parameter SON FSAR TVA/ Monster End of Blowdown, s 25.7 25.7 PCT at EOB, F 1229 1229 RV Liquid at EOB, ft 3 74.1 74.1 Bottom-of-Core Recovery, s 45.5 45.7 Rupture, s 46.2 46.2 Rupture Elevation, ft 6.3 6.3 PCT, Ruptured Segment, F 1835 1905 PCT Elevation, ft 6.9 6.9 PCT, F 2034 2032 5-28

Proprietary Figure 5.4-1 Plant Specific Studies Analysis Diagram CONTA!NMENT PRESSURE ECC STUDY SPECTRUM STUDY STUDY DISCHARGE BREAK COEFFICIENT TYPE SON FSAR LOCA PCT = 2034 F UMITS GUILLOTINE PCT = 2034 F Cd = 1.0 TVA/ MONSTER

[ PCT = 2034 F PCT = 2032 F MIN ECC Cd = 0.8 SPUT PCT = 2034 F PCT = 1992 F PCT = 2033 F BASE MODEL Cd = 0.6 PCT = 2014 F PCT = 1865 F MAX ECC PCT = 2014 F 5-29

Non Proprietary Figure 5.4 Sensitivity Study - Base Model System Pressure During Blowdown 2400 I ,

I DECLG Break, Cd = 1.0, Ma4 ECC i

2000 -

- - * - - - l- -- i

+ -

l l

1600 '

-l b y - -- --

l I 1200 3

i \

800- - --- 4--

400- -

g g I  !

O 5 10 15 20 25 30 TIME, s Figure 5.4 Sensitivity Study - Base Model Mass Flux During Blowdown at Peak Power Location 160 DECLG Break Cd=1.0, Mar ECC 120- -~~- L b J--

80- -- -

- ~ ~ -

D o r 40- - - - - - - -

- m g

m 40 --

(v 1--

a y

v

• - ~ ~ ~ - ~ ~ ~ ~ ~ ~~~ - ~-

120- --

160 0 5 10 15 20 25 30 TIME, S 5-30

Non Proprietary Figure 5.4 Sensitivity Study - Base Model Minimum Containment Backpressure 25 20- -- }---- d-- - - - - - l ~~

f 154 -a

[ ~~- -- -~~-.

10< * +---

• l i l 1

5- * ' -

4- -- - ~~

} } l 0 100 200 300 400 500 600 TIME, S l Figure 5.4 Sensitivity Study - Base Model l

\

Reflooding Rate i 10 l DECLG Break, Cd - 1,0, MarECC a- ---- - -

-l - - .

I y e. - -

7.---.-----

! E w - . . -. -- --

- --4 _

4 2 .---

L .- -- --

k ,

i 0 l I I

O 100 200 300 400 500 600

! BMES 4

' ^

5-31 1

l

l Non Proprietary Figure 5.4 Sensitivity Study - Base Model Hot Channel Quench Front and Collapsed Liquid Level II j LEGEND Quench Front Location 10- -+ + - - - - - - - - - . -- Collapsed uguid Level .

b 8- 4 --

5 '

6 L - - - -

E i 8 4- --

-L

-g L, v. . ,~ w i ---

' g v"'-# i l 1 <- .-

DECLG Break, Cd -1.0, Max}. ECC l l

e 1

o O 100 200 300 400 500 600 TIME, S Figure 5.4 Sensitivity Study - Base Model Hot Pin Cladding Temperature at PCT Segment 9 2400 i

DECLG Break, Cd =1.0, Max ECC 2000- - -k - - - - - -

i-- -

l i

1600-g---- -- - -- - -

E  !

1200- -

f- - =~

800- * --

. I a

400 1 , )

1 0 -

I O 100 200 300 400 500 600 j TIME, S G

5-32 l

l l

Non Proprietary Figure 5.4 Sensitivity Study - Base Model

\ 2400 Hot Pin Cladding Temperature at Rupture Segment 11 l DECLG Break, Cd = 1.0, Ma ECC 2000- - - - - -

y b .-.

I

- .,I- .

l i1600 O

1200<

g 800 e_

L -- - - - - .

l 4o0 _.[ -7 4' .

l a' i 1

0 '

O 100 200 300 400 500 600 TIME, S Figure 5.4 Sensitivity Study - Base Model Hot Pin Cladding Temperature at Segment 12 2400 DECLG Break, Cd=1.0, ManECC 2000- - - - - - - - - - - - - -

g 1600-p - --- - -

1200 --

V h800- - - - -l -

400- - -- I -----

l i 0

O 100 200 300 400 500 600 VMEs 4

s i

5-33 i

I i

I l

'?on Proprietary Figure 5.4 ECC Study

.oumped ECCS Injection Flow Rate 1000 <

l 800 - -- ---

, sng ..h_

- - ~ ~ - _. .-

Y g

"- 400 7

-> k b /'

200 /

/ 1 - -

LEGEND Max ECC .

_._.---- Min ECC .

0 0

100 200 300 400 500 600 '

TIME, S Figure 5.4 ECC Study DownComer Water Level 20 LEGEND Max ECC .

, ---~~~ Min ECC .

15 -

5 10- -

54 -

O O 100 200 300 400 500 600 TIME. S 5-34

.-- . .-- . - . - . - . - = .--- .-_.. .._. . - - . _ - -

Non Proprietary

-s Figure 5.4 ECC Study Reflooding Rate 10 l i i DECLG Break, Cd -1.0, Min ECC i

8' ~ ~ ~ ~ ~ - - - -~~-- ~ ~ ~ - - * - - - 4 3 E. 6 ~~~ -

- - ~ ~ ~ ~ ~~l - - - -- -- - - - - - - - - ~ ~ - - -

4 . .. .

..........._... ....4 .....J......_...

? l I

- 1 2<

t b  !

I I I i 0

O 100 200 300 400 500 600 TIME, S m

Figure 5.4 ECC Study Hot Ch8nnel Quench Front and Collapsed Liquid Level ,

LEGEND Quench Front Location 10 - - -- --- -- - - Coltspsed Uquid Level, 8- - --+ ---- ~

]

h H

8- - -

.. _ /

._ .-~ sA. %

• v -- ---- ~ - ~%~

g-- ./**-*v'n 1, ._

1. p DECLG Break, Cd =1.0, Miri ECC l

0 100 200 300 400 500 600 TIME, S 5-35 i 1

l

Non Proprietary Figure 5.414 - ECC Study hot Pin Cladding Temperature at Segment 9 2400 I i OECLG BreaV,, Cd-1.0, Mio ECC i 1 2000- -

i i

1600- g T i l

l 1200< ~ ~ ~ ~ ~ ~ - ~~~~~~ ~~~ ~ - -

b <

c .

-- - b h 800<

l l

400< --~ ~ 4-l 0

0 100 200 300 400 500 600 ,

TIME, S Figure 5.4 ECC Study H0t Pin Cladding Temperature at Rupture Segment 11

. 2400 DECLG Break, Cd=1.0, Mi ECC i

2000- ---

1 u.

g1600- f-- -:----- -

l1200<

5 800< - - - - -

4 400- - +~--- ----- - . .

1 0

O 100 200 300 400 500 600 TIME, S 5-36

Non Proprietary Figure 5.4 ECC Study Hot Pin Cladding Temperature at PCT Segment 12 2400

. DECLG Bros , Cd =1.0, MiriECC 2000- --

- r- ~---- ~~

i l

16001 ---

4-- ---

\

\

i o

1200<

800-T N

i '

l 400- -t- - t l ~ ~'! ,'  !  ! -

1 i

0 0 100 200 300 400 500 600 TIME, S 1

Figure 5.4 Discharge Coefficient Study - Cd = 0.8 System Pressure During Blowdown 2400 l DECLG Break Cd=0.8, Mio ECC 2000, _ _ .__ - _ . - .

t 1

1600- --

l 1200- ---- -- - - - - ~~ -

l 800- - ~ ~ - \- --

e 4004 -------- -- -- '

l- 1 0

0 5 10 15 20 25 30

TIME, S 5-37

Non Proprietary Figure 5.4 Discharge Coefficient Study - Cd = 0.8 Mass Flux During Blowdown at Peak Power Location 160 , , , ,

i 1 i i I i DECLG Break, Cd =0.8 Mis ECC 120 f- - }-- -h y [

l I . I  !

l  !  !  !  !

80< ' - - - "

- -  ;- -- t-b m  !  !

ao. - .s.~. . _

- . . . - L. . - .... . ! .-

1 S o i

1 R I  ! i o '

d  :

i m

M

~40 -~---+------~~+--~--"- ---i- +-

i 1 5 l l 80- ---------t -

- - i--- + --+

i i I l l l

.,2a, --.

..--4.--.

.]..._-.......z.--

Ji .f.

1 i  !  !

1 I i

-1 c - -

0 5 10 15 20 25 30 TIME, S Figure 5.4 Discharfooding Ref Ratee Coefficient Study - Cd = 0.8 10 DECLG Brea , Cd =0.8. Mi ECC t

8' ------- H ---+-

I i

! I ut G' ---- -

+-- t --

I e

Z 4 . . . . , . . _ . . . .

d i l i i i ,

l i l 1 2< - - - --

4-- - - - ' -

• I

i i 8 I i 0

~

O 100 200 300 400 500 600 TIME, S 5-38

Non Prc.prietary ,

Figure 5.4 Discharge Coefficient Study - Cd = 0.8 Hot Channel Quench Front and Collapsed Liquid Level

'O 12 LEGEND Quench Front Location 10< --- + - - -' *. ------- Collapsed uquid Level .

k 8 + - - - - - - ---

l 6- ~~~----t-- ---------+--- b -

n 4< ~~- --

f 4 ~ - . ~

2'  ; k--

j 3 r t DECLG Breald, Cd -0.8, Mio ECC

' ' i 0

O 100 200 300 400 500 600 TIME, S

[ Figure 5.4 Discharge Coefficient Study - Cd = 0.8 Hot Pin Cladding Temperature at Segment 9 2400 DECLG Break, Cd =0.8, Mi ECC i

2000< ----7 i t h -

)

l1600-3200<

800 400< ~ - - - - -

1 0 100 200 300 400 500 600 TIME, S 5-39

Non Proprietary Figure 5.4 Discharge CoeftlCient Study - Cd = 0.8 Hot Pin Cladding Temperature at Rupture Segment 11 2400 '

l DECLG Brea Cd =0.8, Mio ECC l

2000- --~~ --- -

-t y l  !

' 1600- -- - - - - - -- -

• 1200- ------ I L Q

j 800- - - - - ---

I 400- +

0 100 200 300 400 500 600 TIME, S Figure 5.4 Discharge Coefficient Study - Cd = 0.8 Hot Pin Cladding Temperature at PCT Segment 12 2400 .

i l I DECLG Break, Cd =0.8, Mi ECC 2000- -

t t-1600- - -

I It' 1200 - -

I- t- 1 h800  ;

400-l 0 100 200 300 400 500 600 TIME, S 5-40

Non Proprietary Figure 5.4 Discharge Coefficient Study - Cd = 0.6 I System Pressure During Blowdown l 2400 i

l G I DECLG Break l, Cd i= 0.6, Miri ECC 2000- - - - - - ~ * - - - ~ " ~* ~~ " ^ - - - -

l \ '

!  ! 1 1

1600- - - d--- l l

l 1200 k - - -

i 800 - --- L- -

r -- -- --}-' '

400<

t

! l l g i l t  ; ,

! O 5 10 15 20 25 30 l TIME. S l

O Figure 5.4 Discharge Coefficient Study - Cd = 0.6 l Mass Flux During Blowdown at Peak Power Location l 160 DECLG Brea Cd =0.6, Mi ECC 120- - - - + - - --f - - - - - ~~-

80- \ - -

-t-8 40< -\/ ---- --' ----

l 1 v> 40 -- -

i

- . L .--- /m 1

80 ~~ ~ - - - - ~ ~ - - - - - - -*

• 120' -- - - -- - - -

160 l I 0 5 10 15 20 25 30 TIME, S

't 5-41

Non Proprietary Figure 5.4 Discharge Coefficient Study - Cd = 0.6 Reflooding Rate 10 ,

DECLG Break, Cd = 0.6, Mi ECC i

8 -- 4 -

7 I I 6- -- - -

t i l

E 4 --

+

j c

I i

2< -- -- + - - -

0 O 100 200 300 400 500 600 TIME, S Figure 5.4 Discharge Coefficient Study - Cd = 0.6 Hot Channel Quench Front and Collapsed Liquid Level 12 LEGEND Quench Front Location 10 - - - - ----~~ Collapsed Uguid Level .

1 I

[;- 8 + ---I- - - -

6< - - - e -----+-

4- - Y

/ -- - - - - -

q.ewN ~. w . l 2< - - - - -

p -

i T DECLG Breal<. Cd =0.6, Miri ECC 0 I O 100 200 300 400 500 600 TIME, S 5-42

Non Proprietary Figure 5.4 Discharge Coefficient Study - Cd = 0.6 2400 Hot Pin Cladding Temperature at Segment 9

( l I DECLG Break, Cd =0.6, Min ECC 2000

- T- -

T 1600- - - - - - -- - - -

1200- - - - - - N ~~~ ~~-- ~~-

O

)J  !

h 800- - - - -

---

400- -

0 0 100 200 300 400 500 600 TIME, S (V) 2400 Figure 5.4 Discharge Coefficient Study - Cd = 0.6 Hot Pin Cladding Temperature at Rupture Segment 11 DECLG Break, Cd=0.6, Mi ECC 2000 -

1600- - - - -

6-~ 4-- - ~~~

j 1200 J - - - - - - - - - - - - - - - - - - - _ - - . --

fa00- ---- - - - - - - . .

400- -- - - - . --

l 1

0 '

O 100 200 300 400 500 600 TIME, S 5-43

Non Proprietary Figure 5.4 Discharge Coefficient Study - Cd = 0.6 Hot Pin Cladding Temperature at PCT Segment 12 2400

! l

, DECLG Break, Cd =0.6, Mi.9 ECC l

2000- ,

1600 t -

I .

I s i

)--

1200- '-

v ---

4 o i  !

f- - --- - ---i - - - -

800- ,

i i  !

-- - i 400 -- [ - -]

i i 0

0 100 200 300 400 500 600 TIME, S Figure 5,4 Break Type Study - Split System Pressure During Biowdown 2400 .

l l Split Break', Cd =1.0, Min l ECC l

)

2000- r -- - -T  !

I l i

l 1600 f- l --

g,200 .. _ . . _

$,00 N y __ _._.

400- --

j t

0 0 5 10 15 20 25 30 TIME, S 5-44 O

Non Proprietary

( 40 Figure 5.4 Break Type Study - Split Mass Flux During BIOwdown at Peak Power Location l Split Brea Cd = 1.0, Mi ECC

• ' --- l[

y- - - -

• -------r l --

l g 120-

$ {

en 160- ~ - - - *-

-& ~

l 200- --- -~ 1 -

7----

i 240- - - - - - - - + - - - - - -

7--- ---}--

280 0 5 10 15 20 25 30 TIME. S T

J Figure 5.4 Break pe Study - Split Reflooding ste 10 i

Split Brealt Cd-1.0, Min ECC 8- - - - - ~ ~ -

1 6' ---

- v- - - -

l O

ses= =essese - . me=oes w

d" ,

0 0 100 200 300 400 500 600 TIME, S l v 5-45

Non Proprietary i

Figure 5.4 Break Type Study - Split hot Channel Quench Front and Collapsed Liquid Level 12

LEGEND i  ;

j Ouench Front Location i 10- + -* .! . --

Collapsed uquid level .

l .,

2 8< - - - -~~

6 - - - - - --- -

, 4< --- --

ggwW~ve j 2-

/' T; Split Break, Cd = 1.0, Mir) ECC l 0

O l' 100 200 300

\

400 500 600 t

l TIME, S Figure 5.4 Orcok Type Study - Split Hot Pin Cladding Temperature at Segment 9 O ;

1 2400 Split Brea Cd = 1.0, Mi ECC 2000' l t- l 1600</

1200< -

+

\

g,00 _

400< -

• 1 l L.

0 O 100 200 300 400 500 600 TIME, S 5-46

Non Proprietary Figure 5.4 Break Type Study - Split O 2400 Hot Pin Cladding Temperature at Rupture Segment 11 Split Break, Cd-1.0, Mi ECC 2000

• 6- - --

i  !' - - - -

- - j --

1600 Y -------4-- '

-> -- -L ~~n----

1 1200- - - -

---

+-- -- ~

-d -

800- l---- -- e I-----..

. I k 400, _ . . .

,._.........p-......y__.__... ..

i l i 0

I

~

l l 0 100 200 300 400 500 600 TIME, S I

k/j Figure 5.4 Breck Type Stud - Split Hot Pin Cladding Temperature at PCT egment 12 2400 l

! l Split Break, Cd =1.0, Mi ECC 2000- -- -

1600-f w -- -

1200- -

- - - - - * - - L - --

800- - - - - - - ~

• A -- --

400- --

6-0 0 100 200 300 400 500 600 i

TIME, S 5-47

Non Proprietary Figure 5.4 38 - Containment Pressure Study Minimum Containment Backpressure 25 l LEGEND I'/A/ MONSTER Presstre 20 3 .

- Base Containment Pressure

-r- - J - - - - - --

f 15 ---

~%  :/----- --

7 10 - - . ~ - <

i  :

5-  !-- - - l 0 '

O 100 200 300 400 500 600 TIME, S Figure 5.4 Containment Pressure Study Reflooding Rate 1 10 i' l

l DECLG Break, Cd=1.0, Mi ECC TVA/ MONSTER Containment Pre'ssure a- --

6- ~~ n ~~ . - ~ ~ - - - - -

es Of---

j 2 --- -'-- -

, I I .

0 100 200 300 400 500 600 TIME. S 5-48

Non Proprietary Figure 5.4 40 - Containment Pressure Study Hot Channel Quench Front and Collapsed Liquid Level 12 LEGEND l Quench Front Location 10 J-. '

--~~~- Collapsed uguld Level.

l;" 8- - - ~ ~ -- ~~~- - - - - - - ~~

l 6< --- ---- -- -

g, -l / -

1 - . - - - . . .), . .

DECLG Break, Cd = 1.0, Mirp ECC TVA/ MONSTER Containment Pressure O i I O 100 200 300 400 500 600 TIME, S Figure 5.4 Containment Pressure Study '

Hot Pin Cladding Temperature at Segment 9  ;

2400 DECLG BremyCd=1.0, MiriECC TVA/ MONSTER Containment Pressure ,

2000- - - - - --

l 1600- y

--- - - - = - - - -

1200< - - -- - - - - - -

800 ~~- - -

400 - - - - - - - - ~~-

0 0 100 200 300 400 500 600 TIME, S 5-49

r Non Proprietary Figure 5.4 Containment Pressure Study Hot Pin Cladding Temperature at Rupture Segment 11 2400 I

DECLG Break, Cd-1.0, Mih ECC TV// MONSTER Chntainment Pre'ssure 2000-

• b -*

l 1600- f -- --

1200- - -- - -

O  !

h 800- ---

l l 400- .- _ _ , , , , , , , ,

O O 100 200 300 400 500 600 TIME, S Figure 5.4 Containment Pressure Study

! Hot Pin Cladding Temperature at PCT Segment 12 l 2400 DECLG Break, Cd = 1.0, Mi ECC TW/ MONSTER Contalnment Pressure 2000 -

- f- -

N 1600

}

,200- . - - - - .

7 -- -

800< -- -

O O 100 200 300 400 500 600 TIME, S 5-50

[

l Non Proprietary l

l 5.5 LOCA Limits i \s_ / The LOCA evaluation is completed with a set of analyses done to l show compliance with 10CFR50.46 for the core power and peaking l that will be taken as the limiting LOCA conditions for core l

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 10CFR50.46. Five runs are made at differing axial i 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 core operating limits, and plant operation is controlled such that the local peaking and power do not exceed the allowable values.

l 5.5.1 LOCA Limits Dependencies

! The absolute LOCA limits to power and peaking for each elevation l in the core can be determined through repeated calculations at l each elevation, with successively higher local power levels, until the analysis shows one or more of the applicable acceptance l/ -s'

'\

criteria to be exceeded. The highest linear heat rate for which the criteria are not exceeded is the absolute LOCA limit for a particular elevation. The more practical approach, the one l 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 i analyses are then done to confirm that the assumed limits will meet the applicable criteria.

Figure 5.5-1 shows the axial power peaking limit selected and confirmed as applicable to SQN for 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 limitations comply with l 10CFR50.46. Should the results not comply, the allowed peaking is reduced, and the analysis is repeated until acceptable results are obtained. Likewise, if the results show large margins of compliance, the peaking may be increased to provide additional operational flexibility. Neither of these steps is necessary for l SON LOCA analyses based on the peaking limit illustrated in Figure 5.5-1.

An additional condition assumed in these analyses is that the allowable peaking will be dependent on fuel assembly burnup in accordance with Figure 5.5-2. This limitation is made necessary

~- 5-51 l- . . -

1 Non Proprietary because, at burnups approaching 50,000 mwd /mtU, the initial fuel enthalpy and internal pressure can become a more severe combination than the BOL values. By assuring that the local heating rates will be limited to those represented by the product of the peaking limit in Figure 5.5-1 and the burnup limit, shown l in Figure 5.5-2, the reduction in power compensates for the j increases in fuel temperature and pin pressure such that the BOL l conditions remain the most severe. (This is discussed in greater detail in the time-in-life sensitivity studies, Section 5.3.2.) i Therefore, Figure 5.5-2 can limit the operation for the Mark-BW fuel. However, at the high burnups where this limit is imposed, restrictions on core operation are improbable because the highly depleted fuel is unlikely to reach the limit within the operational envelopes of the plant Core Operating Limits Report.

5.5.2 LOCA Limits Results To validate Figure 5.5-1, five separate LOCA calculations were performed. The assumed power peaks were centered at the middle of the second through the sixth grid spans. Figure 5.5-3 shows the hot rod power shapes for the LOCA limits cases. The maximum of each power shape corresponds to a different elevation / peak combination defined by the peaking limit of Figure 5.5-1. The containment pressure response used in the LOCA limits calculations is illustrated in Figure 5.5-4. The SON FSAR is the source of this curve. Use of this curve is justified on the basis of its conservatism in the plant-specific sensitivity studies, Section 5.4.5.

The results of the calculations are tabulated in Table 5.5-1 and shown in Figures 5.5-5 through 5.5-34. The figures comprise five sets with six figures in each set. The six figures of each set show (1) the pressurizer pressure, (2) the mid-core mass flux, (3) the core flooding rate, (4) core collapsed liquid level and quench front elevation, (5) cladding temperature at the hot spot and rupture locations on the pin, and (6) the distribution of local cladding oxidation along the pin.

Pressurizer pressure is shown for each of the LOCA limits cases and indicates insensitivity of the general RCS response during blowdown to the chosen axial power profile. This conclusion is supported by the core mass flux responses that are also plotted for each case. Core mass flux is also relatively insensitive to l core location, similar to the results reported in a separate BWNT RSG LOCA EM application, Reference 5-8. Differences in REFLOD3B-generated core flooding rates are difficult to gauge from the plotted results. Reflood rates are, however, retarded slightly '

for the core inlet-peaked power profiles. l Hot channel quench front progression is superimposed on the hot I channel collapsed liquid level for each case. The plots have no j 5-52 l

l l

Non Proprietary published predecessors in previous BWNT RSG LOCA EM applications C) but do give a good indication of the extent to which the core is flooded at any point in time. The quench front plots also characterize modifications in quench front progression modeling introduced in the BWNT RSG LOCA EM update, Reference 5-2. Both the hot channel quench front and collapsed liquid level increase most slowly and to a lesser elevation for the core inlet-peaked power profiles, an expected trend considering the higher relative integrated heat rates for those cases.

To demonstrate the cladding temperature results, two curves are presented for each case. Temperature histories are shown for the rupture location and for the PCT node. Cladding rupture occurs in the vicinity of the peak power. PCT occurs at a location of lesser power, generally somewhat removed from (below or above) the rupture location because the rupture promotes localized cladding cooling. l The last figure in each LOCA limit set shows the local oxide thickness as a function of elevation for the fuel pin. Each figure shows total oxidation including that assumed prior to the start of the accident. Oxidation accumulates up to the time the cladding falls below 1000 F or the elevation has been quenched, as predicted by REFLOD3B for the average channel. The large l variations of the resultant curve reflect the relatively lower cladding oxidation in the vicinity o.5 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. The relatively high integrated heat addi~t ion in the lower region of the hot channel delays cladding quench and slows the reflooding rate. Cladding temperatures for this case are quite high as a result. Rupture occurs at the pre-grid node adjacent to the peak power location. PCT occurs in the center node of the grid span above the rupture location. For the 2.9-ft LOCA limit, the PCT is 2112 F. The highest local oxidation, 5.5%, occurs at the PCT location. The whole-core oxidation calculated for this case is 0.77%.

4.6-ft Peak Power Case With the power peaked at 4.6-ft, the cladding temperature l responses resemble closely those resulting from the 2.9-ft LOCA '

limits case. Rupture occurs at the pre-grid node adjacent to the peak power location. PCT occurs at the peak power location. For the 4.6-ft LOCA limit, the PCT is 2080 F. The highest local oxidation, 4.4%, occurs at the PCT location. The whole-core oxidation calculated for this case is 0.67%.

s- 5-53

l l

l Non Proprietary I 6.3-ft Peak Power Case  ;

For a peak power situated near the core midplane, the integrated l heat addition in the entrance of the hot channel is reduced and  !

reflooding occurs more readily than it did for the entrance-skewed power shapes. Quenching of the lower nodes occurs more readily. Additionally, the peak power is in an optimum elevation for cooling from liquid droplet carryout. The lowest PCT, 2034 F, of all LOCA limits is identified with this case.

Rupture occurs at the peak power location. PCT, although closely rivaled in the grid span below the rupture, occurs in the node adjacent to the rupture. The maximum local oxidation, 3.8%,

occurs at both the rupture and PCT location. The whole-core oxidation calculated for this case is 0.66%.

8.0-ft Peak Power Case In accordance with the power peaking limit shown in Figure 5.5-1, this case is run at a slightly lower total peaking than the previous cases. Temperature responses for this case are similar to, albeit more severe than, those resulting from the 6.3-ft LOCA limits case. Rupture occurs at the location of peak power and the peak cladding temperature, 2115 F, is predicted to occur in the grid span nelow the peak location. The PCT location, therefore, receives no benefit from rupture-related cooling. The 8.0-ft LOCA limits case produces the highest PCT of all LOCA limits cases. The highest local oxidation is 5.0% at the PCT location, and the whole-core oxidation is 0.68%.

9.7-ft Peak Power Case Like the 8.0-ft LOCA limits case, this case is run at a reduced total peaking defined by the power peaking limit curve in Figure 5.5-1. Rupture occurs at the location of peak power. A PCT of 2108 F results in this case, occurring in the grid span below the peak power location. Again, the node producing the PCT for this case does not benefit from rupture-related cooling. The highest local oxidation is 5.4% at the PCT location, and the whole-core oxidation is 0.70%.

5.5.3 Compliance to 10CFR50.46 The LOCA limits calculations directly demonstrate compliance to two of the criteria of 10CFR50.46 and serve as the basis for demonstrating compliance with two others. As seen in the figures and in Table 5.5-1, the highest peak cladding temperature, 2115 F, and the highest local oxidation, 5.5%, are well below the 2200 F and 17% criteria. Section 5.6 documents compliance with the whole-core oxidation limit based on the local oxidations calculated for these analyses. Section 5.7 documents the 5-54

l l

Non Proprietary 1

l compliance of the core geometry based on the deformations I

predicted for the LOCA. Compliance with the fifth criterion, long-term cooling, is demonstrated in Section 5.8.

l l

l Table 5.5-1 LOCA Limits Results Elevation of Peak Power. ft l l l

2.Jit LJi fl B J1 M  !

l Item or Parameter l l

End of Blowdown, s 25.71 25.72 25.73 25.79 25.80 RV Liquid at EOB, ft' 76.1 76.1 74.1 76.7 76.6 Bottom-of-Core-Recovery, s 45.25 45.23 45.53 45.45 45.49 Rupture, s 48.30 47.41 46.20 49.14 58.39

( Rupture Elevation, ft 3.4 5.1 6.3 8.0 9.7

\

PCT , Ruptured Segment, F 1848. 1902. 1835. 1954. 1722.

Oxide at Rupture Node, % 3.9 3.4 3.8 3.6 2.9 PCT Elevation, ft 4.6 4.6 6.9 6.9 8.6 PCT, F 2112. 2080. 2034. 2115. 2108.

I Oxide at PCT Node, % 5.5 4.4 3.8 5.0 5.4 l

l Whole-Core Oxidation, % 0.77 0.67 0.66 0.68 0.70 i

\

5-55 l

Non Proprietary Figure 5.5-1 LOCA Limits Study Peaking Umit 3.5 3

3.0 ------+-~-----~-i r " ~ ~--- -?,-- -- - --~l 8

1.2 g 2.5 1.0 N

t' 2.0 - -- - - - - -

-g- -- -- --

0.8 a h i i h 8

$ 1.5- ----

h-------- ~ ~ - - - - - -

0.6 Elevation K(z) Fq 0.000ft 1 000 2.500 1.0 8.285.ft 1000_.2.500_ __

o,4 7.995 ft 0 986 2.415 9.705 ft 0 920 2.300 12.000 ft 858 2.146:

o.5 '

0 2 4 6 O.2 8 10 12 ELEVATION, FT Figure 5.5-2 LOCA Limits Study Burnup Limit M>

5-56

1

Non Proprietary 1

i Figure 5.5-3 LOCA Limits Study 3.5 Axial Power Profiles LEGEND 3.o.----------l---- - - - - -------

2-ft Peaked ProlBe 4-ft Peaked Prose 6-ft Peaked Prose 6-ft Peaked Prose 2.5< 104t Peaked Profue 7 --4g - 3 r-- -t s

i /

/ ~

1. /  %

f

1. g s 2.0 ---

--s

1. / - - - -

- f- --- A q--~\t-- -- -

,s / '

's ,

F '

1.5 ------y --/ r/- - - - -- - - - - -- - - - - - --as

's \ (

i z  %-A --

p / s \

s '

1. /

\

1.0 - - gle / *- -

r- ---

.--m-

%% ~s ~s--

o.5 i o 2 4 6 t

8 10 12 ELEVATION, FT l

Figure 5.5-4 LOCA Limits Study Containment Backpressure l 40 LEGEND X FSAR Figure 15.4.1 1 32 '

i

.easonage messessosomsente eneessagesessenaue

• 88**e86 16- --

-c ---- r 8' - - - - - - -~~

-ee .

l 0 50 100 150 1 200 250 300 TIME, S 5-57

Non Proprietary Figure 5.5 5 LOCA Limits Study - 2.9-ft Case i 2400 Pressurizer Pressure 1 I

! I i i 2000 -- - - - -- . - - -

. ft ... . . . .. d .. _ .,, ,, ,, _, ]1 ,

^ - - + - - -

g 1600- - - - - - - - b. .. . , -.--..j,..... ,

E i d

3 1200< - - - - - ~ ~ - - - + - ~ - - . . 3 n g . T. -

i

' l i .

t 800 -- --

-..-]............_.. . . . , . ,,_},,,,.,,,,,,,,,,,,,,_

l 1

400- - -

- . - . . ~ . . . . . . . . l f  !

I i 0 I O 5 10 15 20 25 30 TIME. S Figure 5.5-6 LOCA Limits Study - 2.9-ft Case 160 Core Mass Flux I

l l LEGEND I

i 120- --- - - - - - - 4---- - . . . ,

1 Mid-Core Location .

I 80 , --

7 - --- t- . . . . - ..-

l .3

• i 40- N .~~ - - - - - .w . ----

$ \

s .i '

x

. O. - -. i . _ . . _ . _ . - -

i  ; I 40' - -- --

t--'- -

t- - -- - l- -

I 80 .-

1

}

1 1

120<

h. - - -

I -

160

, i

' I O 5 10 15 20 25 30 1

TIME, S 5-58 l

l I

l

1 Non Proprietary Figure 5.5 7 LOCA Limits Study- 2.9-ft Case l Reflooding Rate

(] l .. LEG END Core Entrance Location s- .- - - -

. - -- - -. l.- -

-.1- --.-- .

I j- -_ _. -

g 4, - - _

l g V

l O- -

-- A ---- -.-- -

g i

l 1

.a . - -

I

\

f \

8 '

\

i 0 100 200 300 t 400 500 600 I

1 TIME, S i

l Figure 5.5-8 LOCA Limits Study - 2.9-ft Case Hot Channel Quench Front and Collapsed Liquid Level  !

1 l

LEGEND j l

t l Quench Front Location 3o , - . . _

Collapsed Liquid Level k 8- --- ~ ~ - - -

-.~~~--+--~~-- -

g, _ - - - .

_ _ --,.- . _ -- -.~. - -

bl l 4- = - -

1

1. ----~

1,

m. m 3 .. .

~~*wva % -

I e j l l 1 0

0 100 200 300 400 500 600 TIME S t l

5-59 1

Non Proprietary Figure 5.5-9 LOCA Limits Study - 2.9-ft Case Cladding Temperature 4

LEGEND i  ! ,

I i l , -_  ! Rupture Location - Segment 6 2000- -- f'!- -- %,h-k'- - - -- ~~--~~-- PCT Location - Segment 8 r I N

/

1600 --  !/ N'N'% s -- +  !

$ \ N  :

l N

1200 f y -

l i

\. . 3 i .

80o. . . _ . . _ ....... (. .  ;  ; .L _ . .

. i i  ;

} I I '

i i 4oo. ._ . . .!. _... _. _ ; . . . _ . _ . . . . _ . _ . .

i i l i

1 i n

0 100 200 300 400 500 600 1

TIME, S l l

O!

l Figure 5.5-10 LOCA Lirnits Study - 2.9-ft Case Hot Channel Local Oxidation 10 l

8 -j- t {

l l l

$ s. :_. _ . . . . - . -. --._ f .. -4.-

i k

< i n 4- ~~-1-~---~~~-- - - - ---'

, r- - -- -- - -- J l

2' - - - #

0 2 4 6 I in 8 10 12 ELEVATION, FT 5-60

Non Proprietary i

l Figure 5.5-11 LOCA Limits Study - 4.6 ft Case Pressurizer Pressure 2400 e

2000, i I

.) 4. ..4.-.....-

1

• 1600 - - - - - - - - - -----

' l i

E '

1 1200- -- -- -- - - --

800< - - * - - - - ! - -- --L -

k----< l I  !

l 400- - - - - - - - l

---k-

)

i i ( i i

0 O 5 10 15 20 25 30 TIME. S ,

1 l  % i i

s Figure 5.5-12 LOCA Limits Study - 4.6-ft Case Core Mass Flux

! 160 f LEGEND t

120< -- - - - -- - - - - - - - -- - - - - . Mid-Core Locatiort l

80 1 - - - - ~ ~ ' ~ --

- h - - ----

40 . __ ___ ~ __- .-.._ . - - __,._ . . _ _ -

0< ~~~--- - - - - - -- & --

t i

I l 40 --

L 1 i

i 80 - - - - - ---- - L-- -i 120 - ---

I

160 0 5 10 15 20 25 30 i

TIME. S 5-61

Non Proprietary Figure 5.5-13 LOCA Limits Study - 4.6-ft Case Reflooding Rate LEGEND Core Entrance Location 8' - - - -

-h -

I I 4 -

W 0- d- A - -- I -

4 -

l l

8 0 100 200 300 400 500 600 TIME, S l Figure 5.5-14 LOCA Limits Study - 4.6-ft Case Hot C1annel Quench Front and Collapsed Liquid Level O1 12-LEGEND l

Quench Front Location I 10' 6- - 5..

- Conapsed uquid Level

]

1 t;- 8 ---- ! L

g. -

r___ q. _=

l4 ,

f ,,,,, - /*.* s . -^~*ge ,e.,a..w%

,. s 2< -

j 0

0 100 200 300 400 500 600 TIME, S 5-62

Non Proprietary Figure 5.5-15 LOCA Limits Study - 4.6 ft Case Cladding Temperature 2400 l LEGEND 1

1 Rupture Location Sege wnt 9 2000'

/N N ---- ---- PCT Location Segmen: 4 1600-y "wiN__ g 1200< '

• \ - - - - - - - - -- -- -

\

800<

j --- -

h- ---+-------

l i

i l

400- ---- 4- ~ ~ - ~~ - - - -

- - b"

--f- k. <

h' h . T ^

~"

o 0 100 200 300 400 500 600 TIME, S Figure 5.5-16 LOCA Limits Study - 4.6-ft Case to Hot Channel Local Oxidation l

8 - - - - - - - - -

( 6' -

G i 5 4 --

t --- --

2< --- -

1 3 l '

0 '

0 2 4 6 8 10 12 ELEVATION, FT 5-63

Non Proprietary Figure 5.5-17 LOCA Limits Study - 6.3-ft Case Pressurizer Pressure 2400 1

2000- -- * ~

1600- - - - - -- -

, 1200< --- ---- -- -

l l

i 800-400- ----

h

\ T i

0 -

0 5 10 15 20 25 30

, TIME. S Figure 5.5-18 LOCA Limits Study - 6.3-ft Case Core Mass Flux 160 LEGEND 120 , 4 Mid-Core Location, 80< -- + - ---- - -

40- ---- -- - -- -- -

o. - ----

j y y -. .

, 80- ---- - ~ .

-120  ; ~ ~ - - - - - - - - - -

-160 O 5 10 15 20 25 30 TIME. s 5-64

Non Proprietary Figure 5.5-19 LOCA Limits Study - 6.3-ft Case Reflooding Rate 12 LEGEND Core Entrance Location g, _

-t ___. 4 . -

! i t

<. . - - . _ . = - - - l n

o _

l--- . L. L _.. . _ . . . . - _ _ _ . ._

4 .

...,_._....._p- ._4 _ -4

I I

i i

~8 i

0 100 200 300 400 500 600 TIME, S b F

\ Hot hIure aannel 5.5-20 Quench LOCA Limits Front andStudy Collapsed - 6.3-ftLiquidCase Level 12 LEGEND Quench Front Location 10< - - - - l----- L -

Collapsed uquid Level b -

8- - - -

--4 - -

\

l< #

e- ---- -- d- -

4

• l

__ A.% -

-a v .

1 j' - -

O 0 100 200 300 doo 500 600 TIME, S 5-65

Non Proprietary l

Figure 5.5-21' LOCA Limits Study - 6.3-ft Case Cladding Temperature l LEGEND I

l Rupture Location Segment 11 i- ------ - PCT Location SeDment 12 2000-f 1 --

1600 3-- '"vs- -- -

\

1200- -

--73~

'k.,

(- 'N t

800- e = - - -

i i i i l l 400- - -- --- ; -

7~ j - --- -- t -

~~~~-

j .

i t 1

! i O

O 100 200 300 400 500 600 TIME, S 6.3-ft Case Figure 5.5-22 LOCA Limits Hot Channel Local Study Oxidation 10 .

l 8- - - - - - -

6- n T~~~ t +-~~

4 -- - + t +

o 2- t "

I \ ,

l i

i l o

O 2 4 6 8 10 12 ELEVATION, FT 5-66

Non Proprietary a

Figure 5.5 23 LOCA Limits Study - 8.0-ft Case Pressurizer Pressure 2400 1

2000 4------- - +-- -

i l

1600 - - - - - - - - - - - -

1 1200- - + - - - - - -- --~~ -

1 1

l J

l a00 -- --- - - --

i l

1 400< -

--;~~~---r--- l = - -

)

1

{ l I  !

0 I i i 1 0 5 10 15 20 25 30 1

)

TIME, S l

s . .

Figure 5.5-24 LOCA Limits Study - 8.0-ft Case Core Mass Flux 160 LEGEND 120 -- - - - - - -~ ~ - . Mid-Core Location.

80- ~~~d - - - ~

k o

40- ---- - - - - - - ~ ~~~~-~~ ~ - --

f 0- -

---}----- ---- -

1- ---<

2

-4 0 - -

,,r ~~~

2 .

80 ~~< ~ ~ - ~ ~ - ~ ~ - -- - - - - - - - - - - ~~4 -----

120- - - - -

160 0 5 10 15 20 25 30 TIME, S

\

5-67

4 1

Non Proprietary I

Figure 5.5-25 LOCA Limits Study - 8.0-ft Case i Reflooding Rate i 12 LEGEia I

Core Entrance Location i

8 _ _ , . _ . p _

l I

4- --

p i  ;

8  ;

_m- '

J i ,

Io.

._ t.. _. _ --

l 4 4._

8 i O 100 200 300 400 500 600 '

TIME. S l l

Fit ure 5.5-26 LOCA Limits Study - 8.0-ft Case Hot C.lannel Quench Front and Collapsed Liquid Level O

II LEGEND Quench Front Location 10 . .,

-- Conapsed uquid Level k 8< b --

2 Y--

B s< 4 -

t

/ - - - - - .

g , ,

a 4 r r

f_,,,a...  % # ^## r-.a/**-W"""~"-~~*--

N '

,p~/,

2< --

t O

O 100 200 300 400 500 600 TIME. S 5-68

Non Proprietary

~

4 f Figure 5.5-27 LOCA Umits Study - 8.0-ft Case Cladding Temperature 2400 LEGEND I

g,, Rupture Location . Segment 14 2000- J '.aql -. .. ---- PCT Location - Segment 12 )

{

%g.

1600- L --

dA ---- ---

i 4

1200- -

- - - - - - - - - . ~ ~ - -

4 - - - - - - -

i 1

800' 3--- - --

i k

i x

4 l' t 400- - - - - - - -

--' h----

id ~ - . . . _

j O

O 100 200 300 400 500 600 4.

TIME, S i

!O

' 8.0-ftion Case Figure 5.5-28 Hot Channel LOCA Local Limits Oxidat Study j 10 4 .

1 1

3 s 8< --

g . . - - ...

a 4

l a

• - 6- --~~~ ------

t i

1 i

0 4' ~~ ~ A ~"

i i 2 - - --- e ,

! O l ~

0 2 4 6 4 8 10 12 i p. ELEVATION, FT i

!(

5-69 4

1 i

Non Proprietary Figure 5.5-29 LOCA Urnits Study - 9.7-ft Case 2400 Pressurizer Pressure 2000<--- - -

-l 1600- - - - - - - - -

---{ --t E I 1200- - --- -- - ---- --

E 800' \- -

r-

--]

400 --

l  !

O '

O 5 10 15 20 25 30 TIME, S Figure 5.5-30 LOCA Limits Study - 9.7-ft Case O

160 Core Mass Flux LEGEND 120- -- -

Mid-Core Location.

80 -

-- ~

~

40- --- - --- -- - - - - -

f

^

O<

g - - - - - - - -

(v' d

us -40 g I .

80- b * - - - -

-120- -- -----

- +---

-160 0 5 10 15 20 25 30 TIME, S 5-70

1 I

Non Proprietary Figure 5.5-31 LOCA Umits Study - 9.7 ft Case Reflooding Rate 12 LEGEND Core Entrance Location a -- --

I E

4 . _ . _ . _ . _ _ . _ . _ _ _ . . _ _ . ___._

1 i

g  %-

0 '

4 L L

- I- - - - - -

1 i

i 4 I

-~~~t T

~ ~ ~ ~ ~'~ ' - '

t '~~'

.a I O 100 200 300 400 500 600 TIME. S Figure 5.5-32 LOCA Limits Study - 9.7-ft Case 12 Hot C1annel Quench Front and Collapsed Liquid Level 10 ------ - - - -

k 8- ------- ---

---

~- ---

l 6- ---

---

~

~~; 7 --

--t-~-------

u 4- - - - - -

l

-.,r~

,gayag.sda agu--.Au----

2 I -

.a .- -

LEGEND

/ Quench Front Location 0

j --

---

Collapsed Uquid Level 0 100 200 300 400 500 600

! TIME. S 5-71

Non Proprietary Figure 5.5 33 LOCA Limits Study - 9.7-ft Case Cladding Ternperature 00 '

LEGEND Rupture Location - Segment 17 2000. '---- -- PCT Location Segment 15

[( bs N'%

/. s 1600- - - - - - - ----

% ---.i =_.

l %

j

, N..

1200- - - - - - ----- y- .- -

1 G

\N

\*.

800-400 _ . . . - - _ . . . _ . . _ _ _.

r .

_. 7 i L.

1 0

O 100 200 300 400 500 600 TIME. S Figure 5.5-34 LOCA Limits Study - 9.7-ft Case H0t Channel Local Oxidation O

10 8 ' -

1. 6- t- +-

-+-

i 9

5 4 j-

. m ,___

..m.pa.me..e. ..

o O

b 2 4 M

6 8 10 12 ELEVATION, FT l 5-72

l Non Proprietary es 5.6 Whole-Core Oxidation and Hydrogen Generation j ks _, The third criterion of 10CFR50.46 states that the calculated 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 method provided in the evaluation model, Reference 5-1, has been applied to determine core-wide oxidation for each of the LOCA limits cases. Specifically, a discussion of the logic utilized to calculate the whole-core oxidation begins in Volume I, page LA-229 of Reference 5-1.

In the calculations, local cladding oxidation continues as long as the cladding temperature remains above 1000 F, and the REFLOD3B analysis did not show that the cladding, at a given location, is quenched. These local oxidations are summed over the core to give the core-wide oxidation. The figures in Chapter 5.5 illustrate the local oxidation, predicted in each of the LOCA limits analyses, 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 as a result of a large break 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 0.77 4.6-ft Peak 0.67 6.3-ft Peak 0.66 8.0-ft Peak 0.68 .

9.7-ft Peak 0.70 l

Because these cases represent a range of the possible power distributions that can occur in the plant, the maximum possible oxidation that can occur during a LOCA at the SON plants is calculated to be less than 0.77%. Thus, the third criterion of 10CFR50.46, which limits the reaction to 1% or less, is met with considerable margin.

5.7 Core Geometry The fourth acceptance criterion of 10CFR50.46 states that calculated changes in core geometry shall be such that the core remains amenable to cooling. The calculations in Chapter 5.5 directly assess the alterations in core geometry that result from the LOCA at the most severe location in the core. These calculations demonstrate that the fuel pin is cooled successfully. As discussed in Volume I, Section 7 of the BWNT 5-73

Non Proprietary RSG LOCA EM topical (Reference 5-1), clad swelling and flow blockage due to rupture can be estimated based on NUREG-0630.

For SON, the hot assembly flow area reduction at rupture is less  ;

than 60% for all LOCA limits cases. The upper limit of possible channel blockage, based on NUREG-0630, is less than 90%. Neither 90% blockage nor 60% blockage constitutes total subchannel obstruction. 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 being cooled.

The effects of fuel rod bowing on whole-core blockage are  !

considered in the FCF fuel assembly and fuel rod designs, which  !

minimize the potential for rod bowing. The minor adjustments of i fuel pin pitch due to rod bowing do not alter the fuel assembly l flow area substantially and the average subchannel flow area is i preserved. Therefore, due to the axial distribution of blockage i caused by rupture, no coplanar blockage of the fuel assembly will occur and the core will remain amenable to cooling. Deformation of the fuel pin lattice at the core periphery can occur from the combined mechanical loadings of the LOCA and a seismic event.

These loads have been analyzed separately to ensure that they have no adverse effect on the core cooling processes. The loadings and effect on the Mark-BW assembly are presented in the Mark-BW fuel mechanical design report (Reference 5-12). Although deformations can occur, they are limited to the outer two or three points on the lattice structure of the core and do not cause a subchannel flow area reduction larger than 35%. 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 10CFR50.46 have been met and the core has been shown to remain amenable to core cooling.

5.8 Long-Term Cooling The fifth acceptance criterion of 10CFR50.46 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 l in the core. Successful initial operation of the ECCS is shown )

by demonstrating that the core is quenched and the cladding  :

temperature is returned to near saturation temperature. )

Thereafter, long-term cooling is achieved by the pumped injection l 5-74 91 l

1

Non Proprietary

' 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 crystalline precipitation that could prevent the coolant flow from reaching certain portions of the core. This section l presents the evaluation of the final stages of the initial operation of the ECCS, a discussion of the long-term supply of water to the core, and a discussion of the procedures to prevent the build-up of boric acid in the core.

5.8.1 Initial Cladding Cooldown The BWNT RSG LOCA EM heat transfer models used to predict cladding temperatures following the initiation of cladding cooldown are conservatively biased and cannot be used directly to predict cladding quench. Core quenching is, after a fashion, predicted by REFLOD3 in the calculation of whole-core oxidation.

After quenching, dominant core heat transfer is by pool nucleate boiling or by forced convection to liquid, depending on the location of the break in the RCS (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 degrees of the saturation temperature of the coolant. Thus, within ten to fifteen minutes following a large break LOCA, the

- core has been returned to an acceptably low temperature level.

[s)

(v,f 5.8.2 Extended Coolant Supply Once the core has been cooled to low temperatures, maintaining 4 that condition relies upon the systems available to provide a i continuous supply of coolant to the core. Detailed descriptions )

of the plant systems and functions are provided in the FSAR for i SQN. Provision for long-term core cooling with the ECCS, as {

demonstrated in the FSAR, is independent of the fuel design. l Thus, the licensing basis for previous operation remains valid 0 for Mark-BW reload fuel, i 1

5.8.3 Boric Acid Cpncentration ]

The long-term cooling mechanism for a hot leg break is forced l convection to liquid. Once cooling is established, the l coolability of the core is assured and need not be further j considered. For the cold leg break, however, there is no forced ]

flow through the core. The liquid head balance between the core  :

and the downcomer prevents ECCS water from entering the core at a f 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 l

/

v)

\

\

5-75

Non Proprietary recirculation mode of operation within 15 hours1.736111e-4 days <br />0.00417 hours <br />2.480159e-5 weeks <br />5.7075e-6 months <br /> of the initiation 5 of the accident.

In this mode, injection paths are aligned so that injection takes place in both the hot and cold legs. By doing so, the amount of injection to the hot leg becomes a through-flow that can contro3 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 solubility limits for the dissolved solids. Therefore, there is no dependency on the iuel element design as the concentrations depend only on the injection rate, the RCS geometry, and the core power level.

Since none of these factors have been altered by the fuel change, the evaluation in the SON FSAR remains valid for the plant. The operator actions and procedures to establish the operation are also described in the FSAR.

5.8.4 Adherence to Lono-Term Cooling Criterion Compliance to this criterion is demonstrated for the systems and components specific to SON in the FSAR 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 capable of recirculation is available and operated by the plant to provide extended coolant injection. The concentration of 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 has been established, and compliance to 10CFR50.46 has been demonstrated.

5.9 Small Break LOCA In addition to large break analyses, small break LOCA analyses were performed in support of the loading of Mark-BW fuel at SON.

Those breaks with an area less than 1.0 ft2 are examined. The small break LOCA is initiated from the same operational state as the large break. Calculations follow the methods outlined in Volume II of Reference 5-1. The results of the small break studies demonstrate compliance with the regulatory criteria of 10CFR50.46.

5.9.1 Small Break LOCA Transient Description The small break LOCA transient can be generally characterized as i developing in the following distinct phases: (1) subcooled depressurization, (2) RCS loop saturation and flow coastdown, (3) loss of RCS loop circulation and reflux mode cooling, (4) loop seal clearing and core refill, and (5) long-term cooling provided by high head ECCS pumps and accumulators.

5-76 O1 1

Non Proprietary Following the break, the primary system rapidly depressurizes to the saturation pressure of the hot leg fluid. During the initial

s. depressurization phase, a reactor trip is generated on low pressurizer pressure, and the turbine is tripped on the reactor trip. An assumed loss of offsite power concurrent with the l reactor trip results in a reactor coolant pump trip.

In the second phase of the transient, the reactor coolant pump

. coasts down. Natural circulation flow in the RCS loops is l sufficient to provide continuous core heat removal via the steam generators. The occurrence of critical heat flux in the core is, therefore, not likely during this phase.

l The third phase in the transient is characterized by RCS loop l draining and the resulting interruption of natural circulation flow. During this period, heat transfer by reflux boiling in the l

steam generator tubes takes place. This mode of cooling is not, however, sufficient to remove all of the core decay heat. RCS pressure stabilizes at an equilibrium pressure above the steam generator secondary side pressure. In this phase of the small i break LOCA transient, the RCS reaches a quiescent state in which I

core decay heat, leak flow, steam generator heat removal, and system hydrostatic head balances combine to control the core inventory.

l s Loop seal clearing and core recovery occur in the fourth phase of a small break LOCA. RCS inventory continues to decrease. Steam venting from the core to the break is, however, blocked by the .

I presence of a liquid loop seal in the reactor coolant pump l

suction piping. The core level is momentarily depressed and a '

I short-lived temperature excursion may occur. Steam build-up at I the core exit and core level depression combine to clear the loop J seal and re-establish a path for steam relief to the break. j Steam venting causes a rapid depressurization, and the RCS pressure decreases below the secondary side pressure. Break mass

flow decreases substantially as well, in the transition from a i liquid to a steam discharge. Once the loop seal is cleared and steam is vented through the break, steam pressure at the core t exit is reduced, the hydrostatic heads associated with liquid l standing in the various RCS components readjust, and the core l recovers.

l Following loop seal clearing, high head ECCS flow may be insufficient to make up for the mass lost through boiling in the core. The duration and magnitude of this imbalance dictates the liquid inventory in the core. The core could, during this period, uncover and experience a temperature excursion. ,

i l' The final phase of the small break LOCA transient is characterized as a long-term cooling period. Steam continues to be relieved and the RCS continues to depressurize; ECCS flow 1

5-77

Non Proprietary 1

increases as a result. At some point in time, energy removal I associated with the break flow in addition to liquid inventory I replacement by the ECCS balance core decay heat addition. RCS l inventory increases, the core is recovered, and any core i temperature excursion is terminated.

5.9.2 Small Break LOCA Evaluation Model I RELAPS is used to predict the reactor coolant system thermal- i hydraulic responses to a small break LOCA. The code has been I approved by the NRC for licensing application and is documented in detail in Reference 5-3. RCS nodalization is based on the I model described in Volume II of the BWNT RSG LOCA EM, Reference 5-1. Nodal diagrams of the SQN small break LOCA model are presented in Figures 5.9-1 and 5.9-2. The small break LOCA model is similar to that used in the large break analyses.

The reactor core is divided radially into two regions similar to that of the LBLOCA model; one region represents the hot fuel assembly and the other represents the remainder of the core. The core is further divided into twenty axial segments. Cross-flow junctions connect hot assembly fluid nodes with the adjacent

" average" assembly nodes. This arrangement allows the computation of hot assembly cladding and vapor temperatures with limited influence by coolant from the average core and provides resolution of the mixture level to within approximately 0.5 foot.

Initial fuel pin parameters are calculated with the TACO 3 computer code (Reference 5-6). The reactor vessel downcomer and upper plenum regions are represented in finer axial detail than those of the large break LOCA model to give a better representation of the void distribution that affects the system hydrostatic balance.

In the small break LOCA model, the RCS is subdivided into two flow loops. One loop represents the broken loop and the other represents the three intact loops. The pressurizer is attached to the hot leg of the composite intact loop. The nodalization is similar to that of the large break LOCA model.

The steam generator tube region is divided into two radial regions. One region represents the shortest half of the tubes and the other region represents the remainder of the tubes. This provides sufficient modeling accuracy to simulate tube draining effects; tube draining can be sensitive to tube length.

The reactor coolant pump suction nodalization has been altered relative to the large break LOCA model to produce an accurate hydrodynamic representation of loop seal clearing. Two additional nodes are added to the downside of the pump suction pipe and one node to the riser section. This allows finer 5-78 I

Non Proprietary fs resolution of the void distributions and elevation heads that

• control the occurrence and timing of loop seal clearing.

The bottom elevation of the lowest node of the intact loop pump suction piping is artificially extended one foot below the corresponding node in the-broken loop. This preferentially promotes the clearing of only the broken loop. An RCS configuration characterized by a single clear loop and three -

intact loops conservatively restricts steam flow to the break.

The added restriction can result in worsened core conditions and the potential uncovering of the core.

Both the broken loop and the intact loop reactor coolant pump discharge piping are modeled as four nodes. In the large break model, the intact loop is modeled as one node. Using four nodes provides an accurate simulation of the hydrodynamic effects of the ECCS injection.

The computer code options and generic input requirements used in the small break portion of the BWNT RSG LOCA EM are summarized in Volume II, Tables 9-1 and 9-2, of Reference 5-1. In addition, a minor BWNT RSG LOCA EM change is made to alleviate nonphysical conditions in the artificial leak sink volume: the selection of phase-equilibrium modeling in the leak volume. This change permits relatively stable conditions in the leak volume and has no significant impact on transient results.

ps I The small break LOCA model is constructed within the guidelines of the BWNT RSG LOCA EM established in Reference 5-1. It has been developed utilizing the RELAPS large break LOCA model, described in Section 5.2 of this report, as a basis. The small break model adheres to the requirements of 10CFR50 Appendix K and contains demonstrated conservatism for the evaluation of ECCS mitigation of a postulated small break at SON.

5.9.3 Inputs and Assumptions The major plant operating parameters used in the SQN small break LOCA analyses follow. These inputs are similar to those utilized (

in the large break analyses.

1. Power Level - The plant is assumed to be operating in steady-state at 3479 MWt (102% of 3411 MWt).

2. Total System Flow - The initial Reactor Coolant System (RCS) flow is 348,000 gpm.

3. Fuel Parameters - The initial fuel pin parameters are taken from TACO 3 (Reference 5-6) runs performed for BOL fuel conditions.

%-- 5-79

i Non Proprietary i

4. ECCS - The ECCS flows are based on the assumption of a ,

single active failure. A single train of ECCS is modeled as described in Volume II, Section 4.3.2.2, of the BWNT RSG LOCA EM, Reference 5-1. For the case of a centrifugal charging line break, charging flow is assumed to be spilled to the containment.

5. Total Peaking Factor (Fo) - The maximum total peaking factor assumed by this analysis is 2.5. The hot assembly peaking for small break analysis is illustrated in Figure 5.9-3.

6. The moderator density reactivity coefficient is based on BOL conditions to minimize negative reactivity.

7. The cladding rupture model is based on NUREG-0630.

5.9.4 Analysis Results In the SON small break LOCA analyses, five break cases were considered independently to predict core and system responses over a spectrum of break sizes. Small break spectrum results in Volume II, Appendix A, Reference 5-1, indicate that break areas corresponding to 2- to 6-inch diameters produce the most severe core depression. Breaks of 2 , 3, 4- and 6-inch diameters in the bottom of the reactor coolant pump discharge piping were analyzed for SON. In addition, a 1.34-inch diameter centrifugal charging line break, located in the top of the piping, was analyzed.

Table 5.9-1 presents time sequence of events for each of the small break LOCA cases. Fuel thermal responses for the hot pin are included in Table 5.9-2. Parameters of interest to the small break analyses are shown in Figures 5.9-4 through 5.9-28. There are five sets of figures, each set contains five plots. The five figures of each set show (1) the RCS pressure, (2) the break flow rate, (3) loop seal levels in the pump suction downflow and upflow pipes, (4) core collapsed level, and (5) hot spot cladding temperature.

A relatively slow depressurization rate occurs in the 3-inch break case The core does uncover, making the 3-inch break the most limiting case of the small break spectrum analysis. The resulting peak cladding temperature (PCT) is 1072 F. Core metal-water reaction for the 3-inch break is negligible because the cladding oxidation rate is not significant below about 1500 F.

For breaks smaller than 3 inches in diameter, core cooling is maintained by a combination of steam relief at the break and reflux cooling in the steam generator. The core does not uncover  !

for these smaller breaks. For larger breaks such as 4- and 6-5-80 I

Non Proprietary l

inch diameters, the rapid depressurization rate following loop

(~~ seal clearing has two positive effects on the core level. One is ,

increased ECCS flow, and the other is increased core level swell, i No core heatup was predicted for these break sizes.

The centrifugal charging injection line break is postulated to allow examination of a small break LOCA'that is characterized by '

I a degradation of high pressure injection. The break size is insufficient to allow significant depressurization of the RCS.

l Coolant addition in the progression of the transient is, i l therefore, governed by the high pressure injection alone. With a broken injection line, a large portion of the ECCS flow associated with the centrifugal charging flow is directed to the break. To ensure a conservative result, all of the charging flow is assumed lost to the break, and the transient is mitigated by safety injection pumps only. The results of the charging line break indicate that the core remains covered by the mixture level and that no core heatup occurs.

The small break analyses are terminated when the break flow rate is exceeded by the ECCS flow rate. Note that the collapsed liquid level at the end of the transient may still be below the top of the core. The core mixture levels at the end of the l analyses are, however, above the top of the active core and the H I

I RCS pressure is still falling. A steady increase in ECCS injection and continued core cooling is therefore assured.

5.9.5 Compliance to 10CFR50.46

\

The small break calculations directly demonstrate compliance to two of the criteria of 10CFR50.46 and serve as the basis for demonstrating compliance with two others. As seen in the figures and in Table 5.9-2, the highest peak cladding temperature, 1072 F, and the highest local oxidation, about 0.002%, are well below the 2200 F and 17% criteria.

The whole-core oxidation criterion of 1% cladding reaction is met as well in the small break LOCA analyses. Whole-core oxidation will be much less than the peak local oxidation figure of 0.002%.

Whole-core oxidation associated with small break LOCAs, utilizing the assumptions and inputs as documented above, is negligible.

The fourth acceptance criterion of 10CFR50.46 states that .

calculated changes in core geometry shall be such that the core remains amenable to cooling. The calculations in Section 5.9 directly assess the alterations in core geometry that result from the LOCA, at the most severe location in the core. These calculations demonstrate that the fuel pin cooled successfully. j Further, for SQN, no hot assembly cladding ruptures occurred in '

any of the small break LOCA cases. Therefore, the assembly i I

l

/%g l 5-81 l

i Non Proprietary l retains its pin-coolant channel-pin-coolant channel arrangement ,

and is capable of being cooled. l The fifth acceptance criterion of 10CFR50.46 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 I

by demonstrating that the core is quenched and the cladding temperature is returned to near saturation temperature.

Compliance to the long-term cooling criterion is demonstrated for l the systems and components specific to SON in the FSAR and is not I related to the fuel design. The initial phase of core cooling  !

has been shown to result in low cladding and fuel temperaturec.

A pumped injection system capable of recirculation is available and operated by the plant to provide extended coolant inj ecc. ion .

Therefore, compliance with the long-term cooling criterion of 10CFR50.46 has been demonstrated. l Ol 1

)

5-82 GI a

..__..m..

i Non Proprietary r"' Table 5.9-1 Small Break LOCA Time Sequence Of Events As_ /

Events. Sec 2-inch 3-inch 4-inch 6-inch CCI Break Initiation 0.0 0.0 0.0 0.0 0.0 Reactor Scram 51.4 23.3 14.1 8.3 114.5 Coolant Pump Coastdown 51.4 23.3 14.1 8.3 114.5 Steamline Isolation 51.4 23.3 14.1 8.3 114.5 Feedwater Isolation 61.4 33.3 24.1 18.3 124.5 ECCS Injection 98.0 65.5 51.9 45.1 165.9 Loop Seal Clearing 2235.7 710.6 364.8 158.6 NA Peak Cladding Temperature NA 1883.1 NA NA NA Accumulator Injection NA 2205.0 1195.0 430.0 NA Table 5.9-2 Small Break LOCA Results l

l Results 3-inch l

Peak Cladding Temperature, F 1072 Peak Temperature Location, ft 10.9 Rupture Time, sec NA Rupture Location, ft NA I

I Maximum Local M/W Reaction, % -0.002 Total M/W Reaction, % <0.002 l 1

i' 1

\% 5-83 l

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3 i

l l 4 ---- l---- E---- = -  !  !

2< - ~r - - - - - - -

0 0 400 800 1200 1600 2000 2400 2800 3200 TIME, S Figure 5.9-18 4-inch Pump Discharge Break Hot Rod Clad Temperature 2400 2000< -

3

,1600< ~ -

1 ----- --- -L _ . _ . ,

l1200<

800< - - - - - + .-

1 L

400- M -

0 0 400 800 1200 1600 2000 2400 2800 3200 TIME, S i

5-93

Non Proprietary Figure 5.9.-19 6-inch Pump Discharge Break Primary System Pressure 2400 2000 - - + - - - + - - - *- -

1600- . - - - - - - -

---7 E

.k1200- --- --

800' t- f **

400- -

\ N O

O 400 800 1200 1600 2000 2400 2800 3200 TIME, S 1600 Figure 5.9-20 6-inch Pump Discharge Break Leak Flow Rate O

j 1400 - - - - -

---

. r-l l

1200 ----- - -

l

' 3 1000 - .

E' l k 800 -

l l

I m 600- - - -

1- ---

p .

400- - ~ ~~-

s- - -

200' 4 p- + - - - - --

0 0 400 800 1200 1600 2000 2400 2800 3200 TIME, S I i

j 5-94

Non Proprietary FIGURE 5.9-21 6-inch Pump Discharge Break Pump Suction Loop Seal Levels 20

_ LEGEND 15 -

^

I----- INTACT LOOP SEAL DOWN.

BROKEN LOOP SEAL DOWN

- G-- INTACT LOOP SEAL UP 10 -+-

I ------ BROKEN LOOP SEAL UP ,

l 5 t

k 5< 9 -- -

h i 0- I i

' L- -

l> i

- l ' I I

5 _ - . -

. .. --__I__

i 10 -

L.A- I 15 -- -- -

20 0 400 800 1200 1600 2000 2400 2800 3200 TIME, S FIGURE 5.9.-22 SEQUOYAH Core Collapsed Level 6 INCH PUMP DISCHARGE BREAK 14 -

12- - - - - -

I" 10- ---- --- - - - - - - - -

a< -

u 6- -- ---4 '

4 __ . _ _ .

2< ~ - - - - - -

0 0 400 800 1200 1600 2000 2400 2800 3200 TIME, S 5-95 i

Non Proprietary Figure 5.9.-23 6-inch Pump Discharge Break Hot rod Clad Temperature 2400 1 -

2000 -

r' , - -- r L I

l 1

,1600 j ~~ --

d l 1 I

i

--1 ~ - - -t ---r 3-l1200- 800- ~ ~ ~

• 1 i -

i

~~4-----.

1 1

i l

N 400 N. ,. l 0

0 400 800 1200 1600 2000 2400 2800 3200 TIME, S Figure 5.9-24 CCI Line Break Primary System Pressure 2400 2000 -

t -

1600- - -- - - - - - - -

9 12 .

i i 1200- - - -

t  ! 2 +- r 8 e h

800 _ -

.. _ ._ p .. . _ _ . _ _ .

400 _ __

t i

i 0

0 400 800 1200 1600 2000 2400 2800 3200 TIME, S 5-96

i Non Proprietary Figure 5.9-25 CCI Line Break (n

\

• 1soo ,

I Leak Flow Rate 140o. . _ _ _ _ _ _ _

! -- b-_ x _

.__ 4_ _. _ ,

i 1200 <-- -- J- --. - - - - -

100o .-

f. 800 -------

i e 800 -

l L. -- -- -.l I 400< ,

l  !

200

• \ l _

l -

o 0 400 800 1200 1600 2000 2400 2800 3200 TIME, S Figure 5.9-26 CCl Line Break Pump Suction Loop Seal Levels LEGEND 15 --L--- -__. f INTACT LOOP SEAL DOWN.

BROKEN LOOP SEAL DOWN G --

10 INTACT LOOP SEAL UP

,-- , , , _ _.p, ------ BROKEN LOOP SEAL UP k 6 -.

d

$ o*-- 0  :  : -

-e ..e--  :

t

-e---.e -

l 5- ~~ -

.g o - ---~~ _ . . _ - -

-15 ' * ---- - -

20 -

0 400 800 1200 1600 2000 2400 2800 3200 TIME, S n

f 5-97

Non Proprietary Figure 5.9-27 CCI Line Break Core Collapsed Level 16 .

i i s I '

! t 14 e b-~ ---y b i-12 Nh - - - -* -

t jo ._._ _.--

% I 4

8 ,

4- --- I 8 s- ---+ -

t b --j--  ;-

i

[-

4 -

1

"- i  !

2 - - - - - --

7 0

0 400 800 1200 1600 2000 2400 2800 3200 TIME, S Figure 5.9-28 CCI Line Break Hot Rod Clad Temperature 2400 l

2000 --

i -j -

, 1600 - - - -

1200' - --- - - - -

800- L-- -

f------ - -- - - - -

I C

i 400- ~+---

0 0 400 800 1200 1600 2000 2400 2800 3200 j TIME, S 5-98

Non Proprietary f ~x 5.10 Evaluation of Transition Cores

! \

\s_,/ During the period of transition from a core fueled with combined Westinghouse VANTAGE SH and Standard fuel to a full Mark-BW core, fuel assemblies may reside next to each other in various mixes for several cycles of operation. This section addresses the LOCA aspects of all types of fuel assemblies in the mixed core configuration. As will be shown, the mixing of the fuel assemblies will not alter the LOCA evaluations of any of the fuel types and the evaluations of the individual fuels, performed as full cores, remain valid for the mixed core condition.

Table 5.10-1 presents a comparison of the most important design parameters for the Mark-BW, VANTAGE SH, and Standard fuel assemblies. Entries in the table indicate geometric dimensions and unrecoverable pressure drops for the three fuel designs. The pressure drops are determined relative to the actual SON core flow.

l An evaluation of transition cores for combined Westinghouse  ;

Standard and Mark-BW fuel designs is presented in Appendix A of j Reference 5-10. The evaluation concludes that LOCA cladding temperatures for mixed core operation will not vary from those calculated for the two designs in whole-core operation. This conclusion also applies to the SON transition core wherein there is a potential for the side-by-side residence of both Standard

[,,sj and Mark-BW fuels.

)

Table 5.10-1 indicates that external fuel pin dimensions are the same for the Mark-BW and VANTAGE SH fuel types. Guide tube and c

instrument tube dimensions do vary somewhat. Unre'overable pressure drops across the two types of fuel assemblies vary as well.

Because the external fuel pin dimensions do not vary for the co-resident fuel types, there will be no effects on the thermal-hydraulic predictions associated with the LOCA analyses. The fuel in the LOCA models is treated as if it were physically uniform and is unaffected by the small differences in pellet length. The very small variations in pellet OD (at most, 3 mils) will not significantly affect the stored energy of the fuel and so would not have an affect on core reflooding rate predictions.

Guide and instrument tube differences do not substantially affect the LOCA evaluation. Depending on core elevation, the internal i flow area of the guide and instrument tubes is larger for the Mark-BW than for the VANTAGE SH. The shift in flow area from I inside the guide and instrument tubes to the heated channels is

! small and inconsequential.

A b

V 5-99

i Non Proprietary The only difference between the assemblies, that can produce a discernable change in LOCA results is the decreased assembly pressure drop of the Mark-BW. The change in fuel assembly pressure drop, VANTAGE SH to Mark-BW at SON conditions, is [ ].

In Appendix A of Reference 5-8, the effects on large break LOCA I

results of mixing fuel assemblies with up to a [] psi difference in pressure drop (Westinghouse OFA-to-Mark-BW) were evaluated.

The effect of assembly pressure drop differences during blowdown was to divert some flow away from the high pressure drop assemblies and toward the low pressure drop assemblies. During reflood the impact was on the whole-core pressure drop, which allows a gradual increase in the flooding rate as the core transitions from high pressure drop to low pressure drop assemblies. The high pressure drop assembly was shown to experience a possible 30 F temperature increase due to the mixed core during blowdown and a compensating 30 to 50 F decrease in temperature during reflood. The low pressure drop assembly l experienced the opposite effects. Since FCT occurs during l reflood, the study concluded that there were no adverse l I

consequences during the mixed core period and that no mixed core penalty is required. The differences in measured pressure drop are about the same for the switch from VANTAGE SH to Mark-BW fuel as the transition from OFA to Mark-BW. The resultant LOCA calculational conclusion is also essentially the same.

l The effect of fuel assembly pressure drop on small break LOCA was addressed in Appendix C to Reference 5-8, page C.84. Because of the abundant coolant flow available during the pump coastdown phase of the transient, and because gravity heads rather than friction flow losses control system evolutions during the core uncovery phase of a small break, a difference in fuel assembly frictional pressure drop even much larger than that between the Westinghouse and Mark-BW fuels will not adversely effect small break LOCA results.

As mentioned previously, there are no differences in the external fuel element dimensions and little difference in guide and instrument tube dimensions between the two fuel types. The minimal differences have an insignificant impact on core liquid inventory. As a result, geometric design differences between the VANTAGE SH and Mark-BW fuel types will not affect ' core mixture levels predicted in the small break studies.

In conclusion, an assessment of the design differences between the VANTAGE SH and the Mark-BW assemblies has shown that the LOCA cladding temperatures for mixed core operation will not vary substantially from those calculated for the various designs in pure core operation. Additionally, the relative average power limitation assessed by Westinghouse for the side-by-side 5-100

-. . ~ - - - . . . - . - - - -. .. -.

Non Proprietary i s operation of VANTAGE SH and Standard fuel assemblies is

\ unaffected by this evaluation and continues to be applicable to

\s_,/

those fuel types. Finally, the calculations for each of the fuel designs provide margin to 10CFR50.46 criteria. Large break LOCA is limiting in terms of PCT. The large break PCT reported for the Mark-BW fuel in Section 5.5 of this report is 2115 F. The PCT reported in the SON FSAR for VANTAGE SH fuel is 2069 F.

The full core LOCA analysis results 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 use of the Vantage SH assemblies, and the technical specifications applied to those assemblies will be based on those analyses. The analysis presented in Section 5.5 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.

Table 5.10-1 VANTAGE 5H/ Mark-BW Design Differences A_/

s c 3 5-101 d'

Non Proprietary 5.11 Summary and Conclusion 10CFR50.46 specifies that the ECCS for a commercial nuclear power plant must meet five criteria. The calculations and evaluations documented in this report demonstrate that the Tennessee Valley Authority's SQN units continue to meet these criteria when operated with Mark-BW fuel. LOCA calculations performed in concurrence with an approved evaluation model (References 5-1 and 5-2) demonstrate compliance for a full Mark-BW core for breaks up to and including the double-ended severance of the largest primary coolant pipe.

The coexistence of the Mark-BW fuel and the Westinghouse Standard and VANTAGE SH fuel in the same fuel cycle is shown to be inconsequential and does not cause the calculated temperatures for either assembly to approach the limitt of 10CFR50.46. The results of the analyses assuming a whole core of Mark-BW fuel are also applicable to the Mark-BW fuel in a mixed core. In addition, the results of LOCA analyses performed by Westinghouse for the VANTAGE SH fuel still pertain for that fuel. Moreover, PCT penalties assessed by Westinghouse for the mixed core operation of VANTAGE SH and Standard fuels, remain applicable to those fuel types.

Specifically, this report demonstrates that when the SON plant is operated with Mark-BW fuel:

1. The calculated peak cladding temperatures for the limiting cases are less than 2200 F, (Sections 5.5 and 5.9).

2. The maximum calculated local cladding oxidation is less than 17.0 %, (Sections 5.5 and 5.9).

3. The maximum amount of core-wide oxidation does not exceed i 1.0 % of the fuel cladding, (Sections 5.6 and 5.9).

4. The cladding remains amenable to cooling, (Sections 5.7 and 5.9).

5. Long-term cooling is established and maintained after the l LOCA, (Sections 5.8 and 5.9).

Large break sensitivity studies and the large break spectrum i studies performed with the BWNT RSG LOCA evaluation model show l that the double-ended guillotine break at the pump discharge with a discharge coefficient of 1.0 and minimum ECCS is the most limiting case. Table 5.11-1 shows the results of this accident on the Mark-BW fuel design when the assumed axial location of peak power is varied along the length of the pin and when the local power for the Mark-BW fuel is controlled such that it can not exceed the Mark-BW peaking limit curve, Figure 5.5-1. These 5-102

Non Proprietary 7

s s results demonstrate compliance with the first four criteria of

( ) 10CFR50.46. Compliance with the long-term cooling criterion is

< \ _,/ achieved through the use of the ECCS in recirculation, drawing water from the reactor building through a heat exchanger to provide extended energy removal. The concentration of boric acid is held below its solubility limit by starting hot leg injection within 15 hours1.736111e-4 days <br />0.00417 hours <br />2.480159e-5 weeks <br />5.7075e-6 months <br /> of the accident.

The small break LOCA spectrum runs were also conducted in accordance with the BWNT RSG LOCA evaluation model. Table 5.11-2 shows the results of the small break study. Of the five break sizes analyzed, the 3-inch diameter break was the only case to result in the core uncovering. The peak clad temperature for this case was 1072 F. A very small amount of cladding oxidation results in this case because the clad temperature is so low.

Adequate core cooling is substantiated by the reaults of the small break LOCA analyses. Cladding rupture dee.a not occur, indicating an unaltered core geometry. The ECCS supply is sufficient to continue core cooling for an extended period and recirculation of spillage from the sump is available thereafter.

For small breaks, compliance with the five criteria of 10CFR50.46 is demonstrated.

During the transition from the Westinghouse VANTAGE SH and Standard fuels to the Mark-BW fuel, all fuel types may reside in

,_ the core simultaneously. Section 5.10 demonstrates that the

/s\ results and conclusions presented above are also applicable to

() the Mark-BW assemblies in the transition core. Section 5.10 demonstrates that the LOCA licensing analysis results associated with the Westinghouse fuels are not adversely affected by the introduction of Mark-BW fuel. Thus, the original Westinghouse small break calculations showing that the Westinghouse Standard and VANTAGE SH fuels met the criteria of 10CFR50.46 remain valid for these fuel types through the transition period.

l

,m l\m e) l 5-103

l Non Proprietary l

l Table 5.11-1 Summary of Large Break LOCA Limits Analyses Elevation of Peak Power, feet M M M M M Parameter PCT, F 2112. 2080. 2034. 2115. 2108.

Maximum Local Oxidation, % 5.5 4.4 3.8 5.0 5.4 Whole Core Oxidation, % 0.77 0.67 0.66 0.68 0.70 1

Table 5.11-2 Summary of Small Break LOCA Analyses l I

Break Diameter, inches M

Parameter i PCT, F 1072.

Oxide at PCT Node, % -0.002 Whole Core Oxidation, % <0.002 l

l 5-104

1 Non Proprietary 5.12 References f'])

(

N~ /

5-1 BAW-10168P Revision 02, BWNT Loss-of-Coolant Accident Evaluation Model for Recirculating Steam Generator Plants, October 1992.

5-2 BAW-10168P Revision 03, BWNT Loss-of-Coolant Accident Evaluation Model for Recirculating Steam Generator Plants, November 1993.

5-3 BAW-10164P Revision 03, RELAPS/ MOD 2-B&W - An Advanced Computer Program for Light Water Reactor LOCA and Non-LOCA Transient Analysis, October 1992.

5-4 BAW-10171P Revision 02, REFLOD3B - Model for Multinode Core Reflooding Analysis, January 1989.

5-5 BAW-10166P Revision 04, BEACH - A Computer Program for Reflood Heat Transfer During LOCA, October 1992. l 5-6 BAW-10162P, TACO 3 - Fuel Pin Thermal Analysis Code, ,

October 1989. l 5-7 BAW-10092P, CRAFT 2 - FORTRAN Program for Digital Simulation

,s of a Multinode Reactor Plant During Loss-of-Coolant, April

) 1997.

\' /

5-8 BAW-10174P Revision 1, Mark-BW Reload LOCA Analysis for the Catawba and McGuire Units, September 1992.

5-9 V. H. Ransom et al., RELAPS/ MOD 2 Code Manual, Volumes 1 and 2, NUREG/CR-4312, EGG-2396, 8/85.

5-10 BAW-10177P, Mark-BW Reload LOCA Analysis for the Trojan l Plant, October 1990. l 5-11 BAW-10184P, GDTACO - Urania Gadolinia Fuel Pin Thermal Analysis Code, February 1995.

5-12 BAW-10172P, Mark-BW Mechanical Design Report, July 1988.

l l

l

\ /

U 5-105 l

than Proprietary f-'s 6.0 Non-LOCA Transient Analysis And Evaluation k )

v Chapter 15 of the Sequoyah Nuclear Plant (SQN) Final Safety Analysis Report (FSAR) is the primary point of reference for the evaluation presented in this section. Each of the accident transients is examined with recpect to continued applicability, of the sequence of events and bounding results developed in the reference FSAR, to reload cycles with Framatome Cogema Fuels' (FCF) Mark-BW fuel. Those transients that are affected by  !

operation with Mark-BW fuel are reanalyzed (Table 6.1-1). Other transients are evaluated, identifying the relevant core-related .

parameters and bounding values to be confirmed for consistency I with the reference safety analyses. I Consistent with Reference 6-2, the following accidents were '

analyzed:

1. Rod Cluster Control Assembly (RCCA) Withdrawal at Full I Power l

2. Loss of Electric Load j

3. Four Pump Coastdown

,, - - 4. Main Steam Line Break (MSLB)

5. Locked Reactor Coolant Pump Rotor The RCCA withdrawal accident is analyzed because it is the  !

limiting Condition II (moderate frequency) reactivity insertion event. Loss of electric load is analyzed because it is the limiting overpressure event for the unit. Four pump coastdown is a Condition III accident that Framatome Technologies, Inc. (FTI) analyzes to Condition II acceptance criteria. As such, it is the limiting loss of flow accident. FTI also analyzes the MSLB to Condition II acceptance criteria. Consequently, the MSLB is the worst case overcooling accident for the unit. The last event,  !

locked reactor coolant pump rotor, is analyzed by FTI because l this design basis accident typically results in some fuel pins experiencing DNB. Therefore, the Mark-BW fuel performance is assessed.

In addition to the events listed above, steam line break with coincident rod withdrawal at power is analyzed because it is considered part of the SQN licensing basis. All other accidents are evaluated to determine that the existing analyses remain valid with Mark-BW fuel or that they are bounded by the accident analyses listed above.

A l \

\x - 6-1

Non Proprietary l

l The analyses and evaluations performed for this section confirm l

that operation of the SQN unit for reload cycles with Mark-BW fuel will continue to be within the previously reviewed and licensed safety limits. Reanalysis of the transients affected by i the fuel reloads demonstrates that the acceptance and design criteria specified in Regulatory Guide 1.70 continue to be met.

! The plant modeling, analysis methods, and input assumptions are described in Section 6.1. The evaluations of Condition II events, Condition III events, and Condition IV events are presented in Sections 6.2, 6.3, and 6.4, respectively. This is consistent with the presentation of these events in the SON FSAR.

Conclusions are provided in Section 6.5, and references are listed in Section 6.6.

6.1 Plant Modeling and Input Assumptions This section describes the plant modeling and input assumptions used to perform the reload safety analyses of non-Loss-of-coolant-accident (non-LOCA) transients for the SQN plant. The plant model is that used with the RELAP5/ MOD 2-B&W computer code (Reference 6-1) to calculate core power, primary system, and secondary system responses during accident transients. The application to SON described here is based upon the generic ,

modeling and application of RELAP5/ MOD 2-B&W to non-LOCA transients reported in Reference 6-2.

The fuel thermal response, including the approach to departure from nucleate boiling (DNB), is predicted using the LYNXT computer program (Reference 6-5) and by application of the statistical core design methodology described in Section 7. When necessary, localized power peaking effects are calculated using the NEMO computer code (Reference 6-7). These computer codes are described in Appendix A. The specific applications of these codes are described with the transient analyses and in their respective code topical reports.

6.1.1 RELAPS Model Description The RELAP5 modeling of the SON plant is essentially the same as that described in Referecne 6-2. Figure 6.1-1 shows the noding arrangement. The SQN plant application model represents a four-loop plant of Westinghouse design operating at 3479 MWt (102% of 3411 MWt). As shown by Figure 6.1-1, a single loop with a single recirculating steam generator, a combined triple loop with triple recirculating steam generator, pressurizer, and reactor vessel are modeled. The recirculating steam generators reflected in the model are the Type 51 steam generator having 51,500 ft2 of heat transfer area.

l l

l 6-2

Non Proprietary Cs No heat structures were included except for heat transfer heat

(\

structures (RSG tubes and fuel clad). Excluding other heat structures is conservative because these tend to mitigate both I

transient heatup and cooldown rates. Cold plant dimensions were adjusted to obtain full power volumes. Water volume has little effect on safety analysis transients, but the larger volumes are conservative for the loss of coolant accident.

The RELAPS/ MOD 2 nonequilibrium option was used throughout, thus j the model can calculate nonequilibrium fluid conditions anywhere two-phase conditions exist in the system.

Fuel Pin and Nucleonics The fuel pin model consists of 5 radial nodes, 1 gap node, and 2 clad nodes. In addition, there are 6 axial nodes. The specific l heat and thermal conductivity for all materials are input as l parameter versus temperature tables. The gap specific heat and thermal conductivity are calculated by RELAP5/ MOD 2 as described in Section 2.3.2 of Reference 6-1.

l RELAPS/ MOD 2 has a point kinetics model that has six delayed l neutron groups. The built-in delayed neutron constants were used because the values are input as delayed neutron precursor yield

" ratios" which do not vary significantly from beginning-of-life l to-end-of-life. RELAP5/ MOD 2 has provision for water temperature, fuel temperature, and water density reactivity feedback. Other

). reactivity feedbacks such as boron are modeled by the use of

! control variables and general tables. Tripped rod reactivity versus time was input as a generalized reactivity versus time table.

The 1973 ANS Standard Fission Product Data built into RELAPS/ MOD 2-B&W were used with a fission product yield factor of 1.0. Also, a 0.4 U-239 yield factor was used. For the steam line break analysis, no decay heat was input.

Reactor Core The reactor core consists of an average channel triple loop core, average channel single loop core, triple loop core bypass, and single loop core bypass. Splitting the core into triple and single loop segments allows the affected loop temperature response to be isolated from that of the triple loop during asymmetric events (e.g. steam line break events). Crossflow l paths are provided at the core inlet and outlet to represent I various flow mixing situations. Both the inlet and outlet plenum I are broken into triple loop and single loop segments.

This core modeling is used for all non-LOCA accident analyses. .

l s

It differs from that described in Reference 6-2 for the full l

! 6-3

Non Proprietary power accidents. The full power model described in Reference 6-2 has a single hot channel, average core and core bypass regions.

Modeling the hot channel is unnecessary because RELAP5/ MOD 2 is not used to predict hot pin response. Consequently, the " split core" model used in the zero power analyses is adopted for all accidents. This simplifies the reload accident analyses because the same control volume and flow path arrangement is used for all non-LOCA analyses.

Reactor Vessel The reactor vessel model consists of a triple downcomer and a single downcomer. The core inlet plenum is modeled as two nodes, one for mixing and the other for core flow distribution. The core outlet plenum has been split into two volumes. The lower volume is below the center line of the reactor vessel nozzles and the upper volume is above the center line of the nozzles. Above the internals upper support plate, the upper head is split into two volumes. The height of the first node corresponds to the top of the guide tubes, and the other node represents the rest of the upper head.

Reactor Coolant Loops The hot leg is separated into 4 nodes. The RSG inlet plenum and RSG outlet plenum are single nodes. The RSG tubes are divided into sixteen segments. The ccid legs are modeled using 5 nodes in the RC pump suction piping, a single node representing the RC pump, and 4 nodes to represent the RC pump discharge piping.

Reactor Coolant Pumps For this model, the RELAP5/ MOD 2 built-in Westinghouse pump data were used. The model is adequate for the reactor coolant flow variations due to transient pressure and temperature changes during transients. Reactor coolant flow variations for pump coastdowns and locked rotor analysis were calculated as described in the transient analysis discussion.

Pressurizer The pressurizer model consists of three parts, the surge line, the eight section pressurizer, and a valve model. The model does not include heaters or spray because they tend to minimize the pressurization and depressurization rates during transients.

Recirculating Steam Generator (RSG)

The single RSG model represents a Westinghouse Type 51 steam generator with a reduction in heat transfer surface area and an increase in flow resistance consistent with 20% tube plugging.

6-4

Non Proprietary The triple RSG model simply has volume and heat transfer capacity fs\ three times greater than that of the single RSG. In either case, subcooled feedwater combines with saturated fluid from the separator to produce a subcooled fluid that enters the four-section downcomer. The four-section tube riser accepts the heat transferred from the RSG tubes.

In the drum region, two liquid volumes are modeled below the separator and two steam volumes are modeled above the separator.

The separator component in RELAPS/ MOD 2 acts as a steam separator and dryer since two-phase fluid enters from the bottom, steam goes out the top, and saturated fluid goes back to the downcomer.

The steam dome is represented by a single control volume.

Feedwater System The feedwater system was modeled as a time-dependent volume and a time-dependent junction. The time-dependent volume can also be used to adjust feedwater temperature versus time, and the time-dependent junction can be used to adjust mass flow' rate versus time. This type of modeling is adequate for the transients of this section where the feedwater flow is terminated early in the transient and can also be used for quite complex feedwater flow situations.

Auxiliary feedwater was not modeled as a separate system since it

[\

usually is not used in the short duration full power transients. l Auxiliary feedwater can be represented as continued flow through '

the feedwater time-dependent junction.

Steam Line l The steam lines run from the steam generators to the turbine, which is modeled as a time-dependent volume to maintain the proper back pressure. Motor operated valve components represent the main steam isolation valves. Main steam stop valves are also represented. Separate junctions and time-dependent volumes are used to model main steam safety valves.

Emergency Core Cooling System Time-dependent volumes and junctions were used to model injection of boron into the reactor coolant system during the steam line break event. A mass flow rate versus pressure table was used to represent the high pressure safety injection (HPSI) system that pumps highly borated fluid into the reactor coolant system. Only the HPSI system was modeled because the reactor coolant pressure does not reach the accumulator tank pressure during the steam line break accident.

p 6-5

Non Proprietary 1

6.1.2 Core and Plant Parameters )

Nominal values of initial conditions were typically used for system parameters except the core power was set to 102% for the at-power accidents. For the DNB calculations, the statistical core design (SCD) methods described in Reference 6-3 were used.

That methodology accounts for control band and measurement uncertainties on those parameters that influence the MDNBR. The reactor protection system setpoints used in this section included the maximum steady-state errors and the maximum trip delay times.

Where specific conservative initial conditions can have an adverse effect on time to trip, valve opening, and so forth, the I effect can be accounted for in the trip value, valve opening  !

setpoint, etc. These methods minimize double accounting for the same error.

l Table 6.1-2 lists the values of the pertinent parameters for the RELAPS/ MOD 2 SON full power and low power plant models. Also listed are the corresponding plant parameter values from the SQN FSAR. This listing of input parameters and initial conditions is consistent with the requirements of l Regulatory Guide 1.70, Table 15-2. 1 6.1.3 Reactor Protection System Setpoints Table 6.1-3 lists the limiting reactor protection setpoints and total trip delay time used in this transient analysis. The trip i delay is defined as the total time delay from the time that trip  !

conditions are reached to the time the rods are free and begin to  !

fall. The difference between the limiting trip point assumed for the analyses and the nominal trip point represents.an allowance for instrumentation channel error and setpoint error. The nominal trip setpoints are specified in the Plant Technical Specifications. The overtemperature aT (OTAT) points used in the accident analyses correspond to the Technical Specification values because the SCD methodology used to calculate the minimum DNBR limit includes measurement and control band uncertainties.

6.1.4 Power Distribution The transient response of the reactor system is dependent on the initial power distribution. The nuclear design of the reactor core minimizes adverse power distributions through operating l instructions and the placement of control rods. Power distribution may be characterized by the radial factor (Fa) and the total peaking factor (Fo) . The peaking factor limits are l given in the Technical Specifications.

For transients that may be DNB limited, the radial peaking factor is of importance. The radial peaking facter increases with decreasing power level due to rod insertions. This increase in 6-6

Non Proprietary l

l

(' N) DNBFa is included in the core limits. All transients that may be r limited are assumed to begin with a Fa consistent with the

\d initial power level defined in the Technical Specifications.

6.1.5 Reactivity Coefficients The transient response of the reactor system is dependent on reactivity feedback effects, in particular the moderator temperature coefficient and the Doppler power coefficient. In the analysis of certain events, conservatism requires the use of large reactivity coefficient values, whereas in the analysis of other events, conservatism requires the use of small reactivity coefficient values. Table 6.1-2 lists the reactivity coefficients used in the transients. Figure 6.1-2 shows, as functions of power, the upper and lower bound Doppler power coefficients used to determine the constant value for Doppler feedback applied in the transient analyses. In some cases  ;

conservative combinations of parameters are used to bound the effects of core life, although these combinations may not represent possible realistic situations.

6.1.6 Rod Cluster Control Assembly Insertion Characteristics l The negative reactivity insertion following a reactor trip is a function of the position versus time characteristic of the rod

- cluster control assemblies and of the variation in rod worth as a

[ ) function of rod position. With respect to accident analyses, the l

( _ ,/ critical parameter is the time of insertion up to the dashpot entry, approximately 85% of the rod cluster travel. Figures 6.1-3 through 6.1-5 show the RCCA position and reactivity insertion ,

characteristics assumed for the accident analyses in this section. The heatup accidents in this section were analyzed i using a conservative insertion time to dashpot of not less than 2.7 seconds. For the cooldown event (MSLB), the event begins with all rods but the most reactive inserted.

l Figure 6.1-5 shows the total negative reactivity insertion versus l

time for a core where the axial distribution is skewed to the lower region of the core. An axial distribution which is skewed to the lower region of the core can arise from an unbalanced l xenon distribution. This curve is used to compute the negative l reactivity insertion versus time following a reactor trip, which is input to all point kinetics core models used in transient analyses unless otherwise noted.

The bottom skewed power distribution itself is not input into the l point kinetics core model. A total negative reactivity insertion ,

following a trip of 4% ak/k is assumed in the transient analyses I except where specifically noted otherwise. This assumption is I

conservative with respect to the calculated trip reactivity worth available.

4

/

s' 6-7

Non Proprietary Table 6.1-1 Summary of Non-LOCA Assessment for Reload With Mark-BW Fuel EVENT ANALYZED EVALUATED Uncontrolled Rod Cluster Control Assembly Bank X Withdrawal From a Subcritical Condition Uncontrolled Rod Cluster Control Assembly Bank X Withdrawal At Power Rod Cluster Control Assembly Misalignment X Uncontrolled Boron Dilution X Partial Loss of Reactor Coolant Flow X Startup of An Inactive Reactor Coolant Pump NA Loss of External Electrical Load and/or Turbine Trip X Loss of Normal Feedwater X Loss of Non-Emergency AC Power to the Station X

Auxiliaries Excessive Heat Removal Due to Feedwater System X Malfunctions Excessive Increase in Steam Flow X Accidental Depressurization of the Reactor Coolant X

System Accidental Depressurization of the Main Steam System X Spurious operation of the Safety Injection System at X

Power Inadvertent Loading of a Fuel Assembly into an X Improper Position Complete Loss of Forced Reactor Coolant Flow X Single Rod Cluster Control Assembly Withdrawal at Full X

Power i l

Rupture of a Main Steam Line X Major Rupture of a Main Feedwater Pipe X Steam Generator Tube Rupture X Single Reactor Coolant Pump Locked Rotor X Rupture of a Control Rod Drive Mechanism Housing (Rod X

Cluster Control Assembly Ejection)

Steam Line Break With Coincident Rod Withdrawal at X

Power 6-8

Non Proprietary Table 6.1-2 (S

Typical Input Parameters and Initial Conditions For Transients Full Pow Low Pow SON PARAMETER Model Model FSAR Neutron Power, MW 3 4 's .' 17 3411 (102%) (0.5%) (100%)

Moderator Temp BOL 7 --

0 Coefficient, pcm/ F EOL -45 -- --

Moderator Density BOL -- -- --

Coefficient EOL --

(1) (1)

Power Doppler Coeff, pcm/% BOL(MIN) -6.5 --

-6.5 EOL(MAX) -12.5 (2) -12,5 Effective Neutron Lifetime, 10-' sec BOL 18.10 18.10 19.40 EOL 18.10 18.10 18.10 Delayed Neutron Fraction BOL 0.0075 0.0075 0.0075 EOL 0.0044 0.0044 0.0044 Minimum DNBR 2.52 --

Axial Power Distribution 1.55 --

N Radial Power Distribution 1. 62 m __

g Reactor Coolant Flow Rate, Mlb/hr 136.2 139.3 138.0 Loop Average Coolant Temp, F 578.2 547 578.2 Average Fuel Temp, F 1419 547 Coolant Volume in Pressurizer, ft 8 Full Pow 1242* --

1080 Lcw Pow --

500 500 Reactor Coolant Pressure, psia 2250 2250 2250 Steam Pressure, psia Full Pow 866 --

910 Low Pow --

1090 1110 Feedwater Temp, F 440 70 440 Total Control Rod Worth, %Ak/k 4 --

4 (1) Function of moderator density.

(2) Function of power level.

(3) Used in statistical core design analysis.

(4) Corresponds to nominal level of 60% plus 10% for measurement uncertainty.

d 6-9

Non Proprietary Table 6.1-3 Trip Points and Time Delnys to Trip TRIP FUNCTION EVAL MOD SON HIGH NEUTRON SETPOINT 118 118 FLUX TRIP (%) DELAY 0.5 SEC 0.5 SEC HIGH PRESSURIZER SETPOINT 2425 2425 PRESS TRIP (PSIA) DELAY 2 SEC 2 SEC LOW REACTOR SETPOINT 87 87 COOLANT FLOW DELAY 1 SEC 1 SEC TRIP (%)

UNDERVOLTAGE SETPOINT 68 68 TRIP (%) DELAY 1.5 SEC 1.2 SEC OVERTEMPERATURE AT SETPOINT ATo = 65.56 F T' = 578.2 F P' = 2235 psig K1 = 1.15 K2 = 0.011 K3 = 0.00055 F/ psi i t1 = 33 see t2 = 4 sec t, = 5 see ts = 3 sec f1 (AI) = 0.0 DELAY 8 sec OVERPOWER AT818 SETPOINT ATo = 65.56 F  ;

T" = 578.2 F i K. = 1.1556 K3 = 0.02 for T > T" K3 = 0.00 for T 5 T" K. = 0.0011 for T > T" ,

K. = 0.00 for T 5 T" l t3 = 10 see

t. = 5 see t3 = 3 sec f2 (AI) = 0.0 DELAY 8 sec LOW SG PRESSURE SETPOINT 447.9 psia SI REACTOR TRIP LEAD 50 sec LAG 5 see DELAY 2 sec (1) Used only for steam line break with coincident rod withdrawal.

6-10

m i

d >

V Jn Proprietary ,

Figure 6.1-1. RELAPS/ MOD 2-B&W Model Used for Non-LOCA Analysis of the SON Plant

8.ssv M m4$!=> !4 = 8 E M l= i =+i t--

n. l m -

[

t t m

l reo ese s l

+

+

l 7so 7ss ess ese j s

I

+

748 1 745 8 ***

+

s A i m x = **s I I ;

n. m i A l l n. l T.. I= m.

t H e l

t I

" ~

a T

1 ls ~ "

=

=

a.. W .= = "

X = ...  ;

a n, .="_ _~_ n.., . .,.  ; . _m _ai _ a

-12. 45. e = -t a fasee d3 . . 44 . 73See 43 *TO R g g, pag .,,,44_ ,,-13_. g,geg

-14 . 43 48

.i. _ 254 73641 73401 7 - -.4 44 83641 p --

830 01

i 4 i , I 4 L

W W--l i's H N _ M=.H = kid-EU--W W a

r m -

m jN 1Ss 255 i

,,, i_.o., - _ --- --,_l - ,,,

l TRIPLE LOOP SINGLE LOOP 6-11

Non Proprietary Figure 6.1-1. RELAPS/ MOD 2-B&W Model Used for Non-LOCA Analysis of the SON Plant (Continued) l =

1

+

1 = I

+ + +

1=l I m I I ,= l w I ,=, l l = l m 9 - ,

i ._ g a= +

I ~

l

+

==

l=l l = H = l l = l

+ +

1x1 + *

= m j

=

1 T' '*

jo m j 2 z:

m l

=

g l=

+ = lI I; ui +

j= m j

+ :I I '

+

306 f 318 319 372

{

= 1 I,' "'

lm m j r'

t t '~7X t -

l 314 315 l 300 4 4 374 l 312 313 l l l i + + i t

l m I 6-12

Non P7_oprietary  ;

Figure 6.1-2. Doppler Power Coefficient

[

r

.ts -

NOTE I l

.te -

14 -

13 -

,I- A 4-4" N0ft I: VPEA CufWE'WT P40ATNE 00PPLEA ONLY P0wtA DEPECT = .t.8% K

.e - .

Norts wwEAcuamaastNEoarNs o a .e a e ion PEACENrPOWEA Figure 6.1-3. RCCA Position Versus Time 1.2 1 * +

h J O., . _. ..

I

{ 0.6 l0.< , -

e 0.2 -, + - 4 -

0 ' '

O O.5 1 1.5 2 2.5 3 3.5 4 TIME (SEC) .

6-13

1 Non Proprietary Figure 6.1-4 Normalized Rod Worth Versus Rod Fraction Inserted i --.. -,

0.8 +

1 g 0.s + 4 ,

l 2

0.4 =

--+ +- 1 i

0.2 + , . .

o D

m. .

0.2 0.4 0.6 0.8 1 1.2 I ROD POSmON (FRACTION INSERTED) 1 F!gure 6.1-5. Tripped Rod Worth Versus Time 4 t:- +:- + +-

p-I 3 +- + -*-- +

e f2 6

4' 1 .. 1 a:

i l

j . . . . .....- ,

l l

0 'I ' ' '

O 0.5 1.5 2 1 2.5 3 3.5 4 I TIME (SEC) 6-14 O'

Non Proprietary l 6.2 Condition II - Faults of Moderate Frequency '

,, Accidents in this category are described in Section 15.2 of the SQN FSAR. Each of the accidents is discussed in the following l sections.

l 6.2.1 Uncontrolled Rod Cluster Control Assembly Bank Withdrawal l From a Subcritical Condition

! The objective of a normal startup is to bring a subcritical  ;

reactor.to the critical condition, or slightly above critical;

increasing power in a controlled manner until the desired power

! level and system operating temperatures are obtained. The ,

increase in power can be achieved by either diluting the .

moderator boron concentration or withdrawing the RCCAs. If, during the startup, an accidental dilution or withdrawal were to occur, an uncontrolled reactivity addition would result. The maximuu rate of reactivity increase in the case of boron dilution is less than that achievable when two overlapping RCCA banks are ,

being withdrawn (see Section 6.2.4) . Thus, the accidental withdrawal of two RCCA banks bounds the reactivity increase of an accidental boron dilution.

The accidental withdrawal of two RCCA banks could be caused by a malfunction of the rod control system. This could occur with the i reactor either suberitical, at hot zero power, or at power. The "at power" case is discussed in Section 6.2.2. This section is I concerned with the subcritical and hot zero power conditions.

The SON FSAR shows that the largest power excursion and the most  !

severe transient conditions occur when the reactivity increase from the uncontrolled RCCA withdrawal is a maximum. The FSAR and the results in Reference 6-4 (BAW-10173P, Mark-BW Reload Safety Analysis for Catawba and McGuire, March 1989) show that the l neutron flux rapidly increases in response to the increasing i reactivity from the continuous withdrawal of two RCCA banks. The neutronic excursion is terminated by the reactivity feedback effect of the negative Doppler coefficient. The reactor protection system, specifically the " Power Range High Neutron l Flux Reactor Trip (Low Setting)" terminates the accident. The j modeling of the protection system includes the most adverse combination of instrument and setpoint errors, as well as delays i for trip signal actuation and RCCA release. In addition, a 10% l increase is assumed for the power range flux trip setpoint, raising it from the nominal value of 25% to 35%.

Framatome Jogema Fuels (FCF) analyzed the accidental uncontrolled withdrawal of two RCCA banks in Reference 6-4. The analysis was l based on bounding RELAPS/ MOD 2-B&W calculations to determine the i neutronic power excursion. The heat flux and temperature 6-15 l '

1 i

I Non Proprietary transients were determined by fuel rod heat transfer calculations in LYNXT. Bounding axial and radial power shapes (associated with having two control banks in their highest worth positions) were used in the analysis. The calculations showed that the fuel temperatures are less than those at nominal full power conditions and the minimum DNBR is always above the limiting value.

Comparison of the pertinent core parameters (e.g. core flow, core j inlet temperature, peaking factors, reactivity coefficients, l differential RCCA bank worths) for Catawba with those for SON l indicates that the analysis in Reference 6-4 bounds SON with FCF fuel. Therefore, for operation of SON with FCF fuel, the normal cycle specific checks of core parameters will ensure that the Reference 6-4 analysis remains bounding.

6.2.2 Uncontrolled Rod Cluster Control Assembly Bank Withdrawal At Power Uncontrolled RCCA bank withdrawal at power results in an increase in the core heat flux. Since the heat extraction from the steam generator lags behind the core power generation until the steam generator pressure reaches the relief or safety valve setpoint, there is a net increase in the reactor coolant temperature.

Unless terminated by manual or automatic action, the power mismatch and resultant coolant temperature rise could eventually result in DNB. Therefore, in order to avert damage to the fuel clad the reactor protection system is designed to terminate any such transient before the DNBR falls below the limit value.

The automatic features of the reactor protection system that prevent core damage following the postulated accident include the following:

1. Reactor trip is actuated if any two-out-of-four power range neutron flux instrumentation channels exceed an overpower setpoint.

2. Reactor trip is actuated if any two-out-of-four AT channels exceed an OTAT setpoint. This setpoint is automatically varied with axial power imbalance, coolant temperature and pressure to protect against DNB.

3. Reactor trip is actuated if any two-out-of-four AT {

channels exceed an overpower AT (OPAT) setpoint. This i setpoint is automatically varied with axial power imbalance to ensure that the allowable heat generation j rate (kw/ft) is not exceeded. l 1

4. A reactor trip is actuated if any two-out-of-four l pressurizer pressure channels exceed a fixed setpoint.

6-16

. .-.. . .- - - - - -- . .- . ~ . . . . -- - ---

1 l

Non Proprietary l

This set pressure is less than the set pressure for the l pressurizer safety valves.

O,V

5. A reactor trip is actuated if any two-out-of-three pressurizer level channels exceed a fixed setpoint when the reactor power is above approximately 10%.

In addition to the above listed reactor trips, there are the l following RCCA withdrawal interlocks:

1. High neutron flux (one-out-of-four power range)

2. Overpower AT (two-out-of-four) t

3. Overtemperature AT (two-out-of-four) 6.2.2.1 Analysis of Effects and Consequences l Method of Analysis The system transient is analyzed using RELAPS/ MOD 2-B&W and the model described in Section 6.1. RELAP5 simulates the neutron kinetics, reactor coolant system, pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generator, and steam generator safety valves. The pertinent system variables including flow, temperature, pressures, and power level are

[\ inputs to a LYNXT thermal-hydraulic analysis of DNBR performance.

In order to obtain conservative results for an uncontrolled rod withdrawal at power accident, the following assumptions are made:

1. Initial reactor power, pressure, and RCS temperature in '

the system analysis are as discussed in Section 6.1. For the DNBR analysis, uncertainties in initial conditions i are included as described in BAW-10170P-A, the topical  !

report on the statistical core design method.

2. Reactivity coefficients are assumed that are consistent with the limiting DNBR result for the maximum rod withdrawal rate at full power. A least positive moderator density coefficient of reactivity is assumed.

A conservatively small Doppler coefficient is assumed.

3. The reactor trip on high neutron flux is assumed to be actuated at a conservative value of 118% of nominal full power. The delays for trip actuation are assumed to be the maximum values.

4. The RCCA trip insertion characteristic is based on the assumption that the highest worth assembly is stuck in its fully withdrawn position. >

4 6-17

Non Proprietary S. The maximum positive reactivity insertion rate is greater than that for the simultaneous withdrawal of the combinations of the two control banks having the maximum combined worth at maximum speed.

Results Figures 6.2-1 through 6.2-6 show the transient response for a rapid RCCA withdrawal incident starting from full power. The calculated sequence of events for this accident is shown in Table 6.2-1. Reactor trip on high neutron flux occurs shortly after the start of the accident. Since this is rapid with respect to the thermal time constants of the plant, only small changes in Tava and pressure result, and margin to DNB is maintained.

Since the RCCA withdrawal at power incident is an overpower transient, the fuel temperature rises during the transient until after reactor trip occurs. A spectrum study of reactivity insertion rates, with minimum and maximum reactivity feedback, was performed. The study showed that for high reactivity insertion rates, the overpower transient is fast with respect to the fuel rod thermal time constant, and the core heat flux lags behind the neutron flux response. Due to this lag, the peak core heat flux does not exceed 118% of its nominal value (i.e., the high neutron flux trip setpoint assumed in the analysis). Taking into account the effect of the RCCA withdrawal on the axial core power distribution, the peak fuel temperature will remain below the fuel melting temperature.

The spectrum study also showed that for slow reactivity insertion rates, the core heat flux re~ mains more closely in equilibrium with the neutron flux. The overpower transient is terminated by the OTAT reactor trip before a DNB condition is reached. The peak heat flux again is maintained below 118% of its nominal value. Taking into account the effect of the RCCA withdrawal on the axial core power distribution, the peak centerline temperature will remain below the fuel melting temperature.

Since DNB does not occur at any time during the RCCA withdrawal at power transient, the ability of the primary coolant to remove heat from the fuel rod is not reduced. Thus, the fuel cladding temperature does not rise significantly above its initial value during the transient.

6.2.2.2 Environmental Consequences The reactor trip causes a turbine trip, and heat is removed from the secondary system through the steam generator power operated relief valves or safety valves.

6-18

-l l

Non Proprietary r- 6.2.2.3 Conclusions The high neutron flux and OTAT trip channels provide adequate protection over the entire range of possible reactivity insertion rates. The minimum value of DNBR is always larger than the limit t value. The plant can be brought to a stable condition with no overpressure of the primary coolant or secondary systems; compliance with General Design Criteria 10, 15, and 26 is assured.  ;

6.2.3 Rod Cluster Control Assembly Misalignment Accidental operation of the RCCAs can be caused by either system malfunction or operator error. RCCA misoperation accidents l include:  ;

1. A dropped RCCA

2. A dropped RCCA bank

3. Statically misaligned RCCA

4. Withdrawal of a single RCCA ,

Each RCCA has a position indicator channel which displays the position of the assembly. The displays of assembly positions are grouped for the operator's convenience. Fully inserted assemblies are further indicated by a rod at bottom signal, which actuates a local alarm and a control room annunciator. Group demand position is also indicated.

1 Full length RCCAs are always moved in preselected banks, and the J banks are always moved in the same preselected sequence. Each bank of RCCAs is divided into two groups. The rods comprising a group operate in parallel through multiplexing thyristors. The two groups in a bank move sequentially such that the first group is always within one step of the second group in the bank. A definite schedule of actuation (or deactuation of the stationary gripper, movable gripper, and lift coils of a mechanism) is required to withdraw the RCCA attached to the mechanism. Since the stationary gripper, movable gripper, and lift coils associated with the four RCCAs of a rod group are driven in parallel, any single failure which would cause rod withdrawal s would affect a minimum of one group. Mechanical failures are in the direction of insertion, or immobility.

The dropped RCCAs, dropped RCCA bank, and statically misaligned RCCA events are classified as ANS Condition II incidents, a fault of moderate frequency. The single RCCA withdrawal incident is classified as an ANS Condition III event, an infrequent fault.

- 6-19

- - _ ~ _ _ _ .

-- - - .\

l Non Proprietary l

The single RCCA withdrawal incident is described in Section I 15.3.4 of the SON FSAR. The event evaluation is discussed here with the Condition II events for convenience.

No single electrical or mechanical failure in the control rod drive system could lead to the accidental withdrawal of a single RCCA from the inserted bank during power operation.

Only the operator could withdraw a single RCCA in the control bank. (This feature is necessary in order to retrieve an assembly should one be accidentally dropped.) However, for the operator to withdraw a single RCCA would require deliberate action. It is extremely unlikely that the operator could accidentally perform the procedure and continue to accidentally ignore or disregard (1) the rod deviaticn annunciator, (2) the rod control urgent failure display on the plant annunciator, and (3) the rod position indicators of the relative positions of the assemblies in the bank. The urgent failure alarm would also inhibit automatic rcd motion in the group in which it occurs.

Without operator action, a single RCCA withdrawal event could only result from multiple wiring failures. If there were multiple wiring failures, the single RCCA withdrawal would also result in the activation of the same alarms and indicators just described.

The probability of such a combination of operator actions or wiring conditions occurring is considered to be so. low that the limiting consequences of this event may include slight fuel damage. (The probability for single random wiring failure is on the order of 10"/ year. )

It has been shown that single failures resulting in RCCA bank withdrawals do not violate specified fuel design limits.

Moreover, no single malfunction can result in the withdrawal of a single RCCA. Thus, consistent with the philosophy and format of ANSI N18.2, and in accordance with General Design Condition 25, this event is classified as a Condition III event.

The withdrawal of a single RCCA from its inserted bank results in both a reactivity increase and increased power peaking in the region of the core surrounding the withdrawn RCCA. The reactivity increase causes the neutron flux to increase and thereby the thermal power, temperatures and system pressure to increase. The automatic termination of this event is provided by the reactor protection system tripping the reactor with the OTAT trip. The increased asymmetric peaking caused by the withdrawal of the single RCCA will be proportional to the difference in the nominal position of bank D and that of the withdrawn RCCA. If the peaking increase is small when the OTAT trip occurs, there is a small probability that DNBR criteria will be exceeded.

6-20

Non Proprietary i

However, if the single RCCA has been fully withdrawn and the lI remaining RCCAs in bank D are deeply inserted, it is possible that in some instances the increase in the peak power density

! will cause the DNB core safety limits to be violated.

I 6.2.3.1 Analysis of Effects and Consequences

1. Dropped RCCA, Dropped RCCA Bank, and Statically Misaligned RCCA Method of Analysis

a. Dropped RCCA:

Nuclear and system evaluations are used to define the bounding system statepoints for the dropped RCCA. The NEMO code (BAW 10180-A) is used to calculate the dropped  !

l rod worths, the maximum control group inserted worths,  ;

l power Doppler defect, and power peaking. LYNXT is used j to calculate the DNBR.

b. Statically Misaligned RCCA:

Steady-state power distributions for the statically misaligned RCCA are calculated with the NEMO code. The ,

ss DNBR is calculated with LYNXT.  !

l Results

a. Dropped RCCA:

Evaluations are presented for three different conditions for i j the system response, l (1) A dropped rod (2 dropped bank included) inserts negative reactivity that may be detected by the power l range negative neutron flux rate trip function. The core is not affected adversely since the power is decreasing in these circumstances.

l (2) For those dropped RCCAs that do not result in a reactor trip, power may be reestablished either by reactivity feedback or control bank withdrawal.

Following a dropped rod event in manual rod control, the plant will establish a new equilibrium condition. The equilibrium process without control system interaction is monotonic, thus removing power overshoot as a concern, and establishing the automatic rod control mode of operation as the limiting case.

! (r~s l

[

l 6-21 l

Non Proprietary (3) Power overshoot after a dropped RCCA incident can only result from the action of the automatic rod controller. For a given PWR system, the power overshoot is essentially a function of the rod controller characteristics. Large power overshoots can result if the rod controller is designed to restore primary system coolant temperature or secondary system steam pressure without regard for the core power level.

The peaking factors for the DNBR calculations are obtained for the most limiting conditions in each cycle.

The minimum DNBR is calculated at maximum power overshoot, pressure, and temperature with measurement and control errors. This calculation is performed on a cycle-by-cycle basis to verify that the minimum DNBR is greater than the design limit.

b. Statically Misaligned RCCA:

The maximum misaligned RCCA occurs when a single rod in the regulating bank is either fully withdrawn or fully inserted. The DNBR for the resultant peak will be calculated and is less than the design limit. This calculation is performed on a cycle-by-cycle basis.

2. Single RCCA Withdrawal Method of Analysis The analysis of the single RCCA withdrawal accident is performed in three discrete steps. These steps involve (1) determining the bounding peak power distributions throughout the core, (2) bounding the system conditions at the OTAT trip, and (3) determining the number of fuel rods that have exceeded the limiting DNBR criteria. The bounding peak power distributions are calculated using the NEMO computer code. The analysis to determine the bounding peak power densities for the worst case of a single RCCA withdrawn from bank D assumes that bank D is inserted beyond the insertion limits for the reactor at full power.

The peaking distributions are used in the thermal analysis to determine the maximum number of fuel rods that may have possibly experienced DNB. Thermal calculations are performed with the LYNXT code series for both the hot channel analysis and to determine the DNBR values.

Previous parametric analyses have shown that this incident is most severe if it occurs at the beginning of life when 6-22

l l

Non Proprietary gN the least negative value of the moderator temperature

coefficient occurs.

l C Using this assumption in the analysis maximizes the power rise and minimizes the tendency of the increased moderator temperature to flatten the power distribution.

Results For the single rod withdrawal event, two cases have been considered as follows:

l a. If the reactor is in the manual control mode, continuous l withdrawal of a single RCCA results in an overall system response similar to that presented in Section 6.2.2.

However, the increased local power peaking in the area of the withdrawn RCCA results in lower minimum DNBRs than for the withdrawn bank event. Depending on the initial bank D insertion and location of the withdrawn RCCA, utomatic reactor trip may not occur fast enough to prevent the minimum DNBR value around the peak fuel rods from falling below the limiting criteria.

Evaluation of case (a) at the system conditions at which the OTAT trip would be expected to trip che plant shows

, - ~ . that an upper bound for the number of rods with a DNBR

[

i less than the limiting value is 5%.

'~'

b. If the reactor is in the automatic control code, the withdrawal of a single RCCA will result in the immobility of the other RCCAs in the controlling bank. The transient will then proceed in the same manner as described in case (a) above.

For case (b), a reactor trip will ultimately occur as in case (a); and again, the trip may not be fast enough to prevent DNBR values in the core from being less than the design limit.

6.2.3.2 Conclusions For cases of dropped RCCAs or dropped banks, for which the reactor is tripped by the power range negative neutron flux rate trip, there is no reduction in the margin to core thermal limits, and consequently the DNB design basis is met. It is shown for all cases which do not result in reactor trip that the DNBR remains greater than the design limit value and, therefore, the

! DNB criterion is met.

l i

I

'l

\ l L/ 6-23

Non Proprietary l l

l For all cases of any RCCA fully inserted, or bank D inserted to l its rod insertion limits with any single RCCA in that bank fully l withdrawn (static misalignment), the DNBR remains greater than l the design limit value. )

For the case of the accidental withdrawal of a single RCCA, (1) with the reactor control either in the automatic or manual l mode, and (2) initially operating at full power with bank D below the insertion limit, fully inserted, an upper bound of the number of fuel rods experiencing DNB is 5% of the total fuel rods in the core. This is consistent with the conclusions presented in Section 15.3.6.3 of the SON FSAR.

6.2.4 Uncontrolled Boron Dilution Boron, in the form of boric acid, in the reactor coolant controls excess reactivity. The boron content of the reactor coolant is periodically reduced during normal operation to compensate for fuel burnup and is reduced after refueling to normal concentrations for startup and full power operation. Deborated water is supplied to the reactor coolant system through the reactor makeup portion of the Chemical and Volume Control System.

Inadvertent boron dilution would add reactivity to the core; and, if not terminated, would cause a reactor trip and/or loss of shutdown margin, depending on the plant operating mode.

Boron dilution is a manual operation under administrative control with procedures calling for a limit on the rate and duration of dilution. A boric acid blend system is provided to permit the operator to match the boron concentration of reactor coolant makeup water during normal charging to that in the reactor coolant system. The Chemical and Volume Control System is l designed to limit, even under various postulated failure modes, l the potential rate of dilution to a value which, after indication l through alarms and instrumentation, provides the operator sufficient time to terminate the dilution and correct the situation.

The opening of the reactor water makeup control valve provides j makeup to the reactor coolant system which can dilute the reactor '

coolant. Inadvertent dilution from this source can be readily terminated by closing the control valve. In order for makeup water to be added to the reactor coolant system at pressure, at i least one charging pump must be running in addition to a reactor makeup water pump.

1 6-24

Non Proprietary fs In order to dilute, two separate operations are required:

ks _s/ 1. The operator must switch from the automatic makeup mode to the dilute mode.

2. The start button must be depressed.

Omitting either step would prevent dilution.

Information on the status of the reactor coolant makeup is continuously available to the operator. Lights are provided on the control board to indicate the operating condition of the pumps in the Chemical and Volume Control System. Alarms are actuated to warn the operator if boric acid or demineralized water flow rates deviate from preset values as a result of system malfunction.

Discussion of Effects and Consequences The cycle specific effects of this transient will be evaluated for the SQN unit for all applicable modes of operation. The reference analyses to be used in future reload cycle evaluation will be those in effect for the cycles just prior to the insertion of Mark-BW fuel. A comparison of boron worths and initial and critical boron concentrations for subsequent cycles to those of the reference analysis will be performed to evaluate ,

/ \ this transient.

In the event that the cycle repecific values do not meet the assumptions of the reference analyses, the transient will be reanalyzed to verify that sufficient minimum operator action time is preserved.

6.2.5 Partial Loss of Reactor Coolant Flow I A partial loss of forced reactor coolant flow accident can be j caused by a mechanical or electrical failure in a reactor coolant pump, or by a fault in the power supply to the pump or pumps supplied by a reactor coolant pump bus.

If the reactor is at power at the time of the accident, the immediate effect of the partial loss of coolant flow is a rapid increase in the coolant temperature. This increase could result in DNB with subsequent fuel damage if the reactor is not promptly tripped.

Normal power for the pumps is supplied through buses connected to the generator. Each pump is supplied from a different bus. When a generator trip occurs, the buses continue to be supplied from

\s ! 6-25

Non Proprietary external power lines, and the pumps continue to circulate coolant through the core. The necessary protection against a partial loss of coolant flow accident is provided by the low primary coolant flow reactor trip signal which is actuated in any reactor coolant loop by two-out-of-three low flow signals.

Evaluation of the partial loss of flow event for SQN shows that the consequences are bounded by those of the complete loss of flow. The complete loss of flow is discussed and analyzed in Section 6.3.2.

6.2.6 Startup Of An Inactive Reactor Coolant I4 Rop SON Technical Specifications normally require that all four reactor coolant pumps be operating while the reactor is critical.

Consequently, this is not a credible event, and it is not evaluated.

6.2.7 Loss of External Electrical Load and/or Turbine Trip A major load loss on the plant can result from loss of external electrical load or from a turbine trip. For either case, offsite power is available for the continued operation of plant components such as the reactor coolant pumps (RCPs). The case of loss of nonemergency AC power is discussed in Section 6.2.9.

For a loss of external electrical load without subsequent turbine trip, no direct reactor trip signal would be generated. Since there is no full load rejection capability, the plant would be expected to trip from the Reactor Protection System (RPS).

In the event the steam dump valves fail to open following a large loss of load, the steam generator safety valves may lift and the reactor may be tripped by the high pressurizer pressure signal, the high pressurizer water level signal or the OTAT signal. The steam generator shell side pressure and reactor coolant temperatures will increase rapidly. The pressurizer safety valves and steam generator safety valves are, however, sized to protect the RCS and steam generator against overpressure for all load losses without assuming the operation of the steam dump system, pressurizer power-operated relief valves (PORVs),

automatic RCCA control or direct reactor trip on turbine trip.

1 The steam generator safety valve capacity is sized to remove the steam flow at the engineered safeguards design rating (105% of I steam flow at rated power) from the steam generator without i exceeding 110% of the steam system design pressure. I The pressurizer safety valve capacity is sized based on a '

l complete loss of heat sink with the plant initially operating at 6-26 l l

l

1 Non Proprietary l the maximum calculated turbine load along with operation of the l steam generator safety valves. The pressurizer safety valves are I .\s then able to maintain the RCS pressure within 110% of the RCS design pressure without direct or immediate reactor trip action.

i A loss of external electrical load is classified as an ANS Condition II event or fault of moderate frequency. The turbine trip event and the loss of external electrical load is

, differentiated in the manner by which steam flow is terminated.

l Termination of steam flow to the turbine following a loss of l external electrical load is completed by the automatic fast l

closure of the turbine control valves. Steam flow termination l

due to a turbine trip event is accomplished by the closure of the ,

turbine stop valves. The response times of the respective valve I closures are not a factor in the analyses since instantaneous steam flow termination concurrent with loss of load is assumed, ,

and credit is taken only for RPS trip functions. The analyses of both transients are therefore similar.  !

l 6.2.7.1 Analysis of Effects and Consequences l Method of Analysis The total loss of load transient was analyzed by employing the detailed digital computer program RELAP5/ MOD 2-B&W (Reference 6-

1) . The program simulates the neutron kinetics, RCS, (es) pressurizer, steam generator and steam generator safety valves.

The program computes pertinent plant variables, including

\ ,/

temperatures, pressures and power level.

In this analysis, the behavior of the plant is evaluated for a complete loss of steam load from full power without direct reactor trip, primarily to show the adequacy of the pressure-relieving devices and also to demonstrate core protection margins.

The following assumptions are made:

1. Initial Operating conditions - Initial core power and reactor coolant pressure are assumed to be at their nominal values consistent with steady-state full power operation. Uncertainties in initial conditions are included in the limit departure from nucleate boiling ,

ratio (DNBR) as described in Reference 6-3. This event j is not the limiting transient with respect to minimum DNBR. It is the limiting transient for primary and secondary system overpressurization. Therefore, to i maximize the pressure response, the initial primary system average temperature was set to the nominal value

( \' ' 6-27

Non Proprietary minus four degrees for control band and measurement uncertainty. A low value for average temperature results in the lowest secondary pressure, providing the longest possible time from termination of steam flow to main steam safety valve lift.

2. Event Initiation - The event is initiated by instantaneous closure of the turbine stop valves.

3. Reactivity Parameters - The total loss of load is analyzed using beginning of life (BOL) reactivity coefficients. A positive moderator temperature coefficient is used. The least negative Doppler power coefficient is used. An end of life b rr o is used in the point kinetics calculation. This combination of parameters yields the greatest peak primary system l pressure.

4. Reactor Control - The reactor is in manual control to maximize the pressure excursion (preclude runback).

5. Steam Release - No credit is taken for the operation of the steam dump system or steam generator PORVs. The steam generator pressure rises to the safety valve setpoint where steam release through the safety valves limits secondary steam pressure at the setpoint value.

The main steam safety valves are conservatively modeled to lift 6% above the nominal setpoint.

6. Pressurizer Spray and PORVs - No credit is taken for the j effect of pressurizer spray and PORVs in reducing or i limiting the coolant pressure.

7. Feedwater Flow - Main feedwater flow to the steam generators is terminated inutantaneously with the loss of load.

8. Pressurizer Safety Valves - The pressurizer safety valves are conservatively modeled to lift 5% above the nominal setpoint. Furthermore, one safety valve is assumed to fail closed, so that only two safety valves are available to limit the pressure excursion.

Reactor trip is actuated by the first RPS trip setpoint reached with no credit taken for the direct reactor trip on the turbine trip.

Results The transient responses for a total loss of load from full power operation are shown in Figures 6.2-7 through 6.2-12. Following 6-28

i i Non Proprietary 7-~g closure of the turbine stop valves, the primary pressure l j increases causing a reactor trip on high pressurizer pressure.

\s_ ,/ Control rod insertion terminates the transient. Actuation of the pressurizer safety valves ensures that the primary system pressure remains below the acceptance criterion. Likewise, the main steam safety valves ensure that the secondary system pressure does not exceed 110% of design. The DNBR generally increases throughout the transient. The sequence of events is shown in Table 6.2-2.

6.2.7.2 Conclusions Results of the analyses show that the plant design is such that a total loss of external electrical load without a direct or immediate reactor trip presents no hazard to the integrity of the reactor coolant system or the main steam system.

Pressure-relieving devices incorporated in the two systems are adequate to limit the maximum pressures to within the design limits.

The integrity of the core is maintained by operation of the RPS, ensuring that the DNBR will be maintained above the design value.

Thus, no core safety limit will be violated.

These results show that the SQN response with Mark-BW fuel to the loss of electric load event meets all applicable acceptance

[( .s\) criteria. Therefore, operation of SQN with Mark-BW fuel does not represent a reduction in the margin of safety.

w/

6.2.8 Loss of Normal Feedwater i A number of causes can result in a loss of normal feedwater to l the steam generators. These include pump failures, valve malfunctions, and a loss of offsite AC power. In all cases, the i loss of normal feedwater represents a decrease in heat removal l capability by the steam generators below the rate of heat I I

generation in the core. The core is protected by the low-low steam generator level trip which trips the reactor long before core thermal limits are approached. Ample steam generator inventory is available to provide, in conjunction with auxiliary feedwater, long term decay heat removal. For the limiting FSAR case, loss of feedwater with loss of offsite power upon reactor trip, the primary system response is the same as that of the loss of offsite power event.

Because of the continued capacity for heat transfer to the steam j generators, the loss of main feedwater event does not approach the more limiting DNBR conditions presented by other transients.

l Although the OTAT and OPAT trips are available to protect the l core, neither of these setpoints is encountered in the FSAR calculations. Post trip system statepoints are well removed from Ch

\x / 6-29

Non Proprietary the thermal limit conditions. The loss of feedwater and loss of AC power events, then, are analyzed primarily to demonstrate sufficient capability for removal of core decay heat by the steam generators without overpressurization of the primary coolant system or discharge of liquid coolant from the pressurizer.

The loss of normal feedwater and loss of nonemergency AC power events have been evaluated for the SON unit with respect to operation with Mark-BW reload fuel. The system conditions assumed in the reference safety analysis to determine the bounding sequence leading to reactor trip, and to startup and delivery of auxiliary feedwater to the steam generators, are not dependent upon the fuel design. The overriding core-related parameter used to bound the results is the decay heat level following reactor trip. The standard curve assumed for the reference FSAR will remain applicable to reload cycles with Mark-BW fuel. Other, less significant, core parameters that can affect these events are the limiting trip reactivities and delayed neutron fractions assumed in the analysis. These same parameters are applied to and confirmed for other transients in the reload safety evaluation. Therefore, no additional analysis or evaluation is required specifically for the loss of normal feedwater or loss of AC power events.

6.2.9 Loss of Non-Emergency AC Power to the Station Auxiliaries Loss of nonemergency AC power can cause loss of power to such plant auxiliaries as the reactor coolant pumps and condensate pumps--with attendant loss of main feedwater--among others.

Possible causes include loss of the offsite grid, with consequent turbine generator trip, and failure in the onsite AC distribution system. A detailed description of the loss of AC power sequence for the SON unit is provided in Section 15.2.9 of the reference FSAR. In the bounding loss of AC power case developed in the safety analysis, the loss of AC power initiates a loss of main feedwater event. The ensuing sequence, with reactor trip on low-low steam generator level, coolant pump coastdown following reactor trip, and long term heat removal via auxiliary feedwater is the same transient as that analyzed for the loss of main feedwater event with offsite power unavailable. That event is discussed in Section 6.2.8.

6.2.10 Excessive Heat Removal Due to Feedwater System Malfunctions Reductions in feedwater temperature or additions of excessive feedwater can overcool the primary system, leading to an increase in core power. The high neutron flux, OTAT, and OPAT reactor trips prevent any power increase which could lead to a DNBR less than the limit.

6-30

Non Proprietary e~w Excessive feedwater flow could be caused by a full opening of one f s

) or more feedwater regulator valves due to a feedwater control (s_ ,/ system malfunction or an operator error. At power this excess flow causes a greater load demand on the RCS due to increased subcooling in the steam generators. At no-load conditions, the addition of cold feedwater may cause a decrease in RCS temperature and thus a reactivity insertion due to the effects of the negative moderator coefficient of reactivity. Continuous addition of excessive feedwater is prevented by the steam generator high-high level trip, which closes all feedwater regulator isolation valves, trips main feedwater pumps and trips the turbine.

The reference safety analysis for the SON unit (FSAR Section 15.2.10) provides descriptions of the causes, effects, and consequences of the increased feedwater event. The primary system temperature response to increased feedwater is a function of the feedwater flow and temperature. Conservative reactivity coefficients are combined with the primary system cooling to obtain conservative reactivity insertion rates. The subsequent reactivity insertion rates are lower than those obtained for rod withdrawal events. Use of FCF fuel in the SON plant would not alter this relationship.

Evaluation of this event for operation with Mark-BW fuel 7- s indicates that the bounding values for core Doppler and moderator

/

) coefficients applied in the reference analysis are the relevant

, ( ,/ core-related parameters. The frequency and magnitude of the excess main feedwater transient will be unaffected by the fuel.

Therefore, confirmation that the reactivity coefficients for reload cores are within the bounding values applied in the reference safety analysis will assure that the consequences of the event will remain within the applicable design criteria and will be less limiting than those of other events in the excess heat demand category.

6.2.11 Excessive Increase in Steam Flow An increase in steam flow is defined as excessive if it is sufficient to cause a mismatch between reactor core power and the heat removal rate at the steam generators. More specifically, increases in steam flow beyond those accommodated by the design of the reactor control system--such as a 10% step load increase or a 5% per minute ramp load increase--may cause an increase in power that can challenge the reactor protection system. In the reference FSAR, Section 15.2.11, the bounding excess steam flow / load event is represented by a ramp increase in steam flow of 10%. Potential causes for this event are excessive loading by the operator, a malfunction in the ste am dump control system, or ,

a turbine speed control malfunction. j l

(sD) 6-31 l l

I l

Non Proprietary The reference safety analysis considers four cases of the 10%

step load increase: The event is analyzed for the reactor in both automatic and manual control modes, assuming both minimum and maximum moderator reactivity feedback cases for each control mode. While the reactor protection system is assumed operable for these analyses, no reactor trips are encountered. In each case, the plant is predicted to reach a new equilibrium operating point at a higher power value corresponding to the increased steam flow. DNBR results are presented, but in no case does the minimum ratio approach the design limit. Thus, the excessive load increase is not a limiting event from the DNB standpoint.

Furthermore, comparison of the primary coolant system conditions for the load increase transient with those predicted for other overcooling events, indicates that the steam generator relief or safety valve failure and the steam line failure events represent more limiting transients in the overcooling category.

The core-related parameters that could affect the consequences of the excessive load event are those that govern reactivity feedback during the overcooling. Use of a lower bound absolute value for Doppler feedback assures conservative prediction of the core power increase. Because both BOL and EOL cases are considered, both upper and lower bound values are used for the moderator density coefficient.

For operation of the SON unit with Mark-BW fuel, confirmation that the reactivity feedback coefficients for potential reload cores are within the limiting values assumed for the FSAR will assure that the consequences and conclusions of the reference analysis remain valid.

6.2.12 Accidental Depressurization of the Reactor Coolant System Spurious opening of a pressurizer relief or safety valve will result in reduction in reactor coolant system pressure. Should this occur during power operation, the decreasing reactor coolant pressure represents a corresponding reduction in the DNB ratio, potentially challenging the core thermal design limits. The reactor protection system provides two reactor trip signals, either of which will terminate this event before the core thermal limits are exceeded. These are the reactor trips on reactor coolant OTAT and on low pressurizer pressure. The transient can be characterized as one of decreasing pressure at a moderate rate while other system conditions--core power, loop flows, and average temperature--remain essentially constant. The OTAT setpoint is responsive to decreasing RCS pressure (whereas the low pressure reactor trip setpoint is static) and is the one encountered in the FSAR calculation. Because there is minimal lag between neutron and thermal power, the minimum DNBR for this event immediately follows the overtemperature trip and is well removed from the design limit.

6-32

l l

Non Proprietary gN The pressurizer safety or relief valve failure event is evaluated i i for SON by a DNBR statepoint analysis in LYNXT. The selection of

\s / the bounding statepoint is straightforward. Transient response of the reactor coolant system is such that the DNBR will decrease throughout the pre-trip period. It will, in fact, decrease slightly beyond the time at which the overtemperature trip is reached because of the trip time delay, the slight lag between thermal power response and neutron power, and the continued.

decrease in pressure during that short post-trip interval (typically, less than a second). The DNBR statepoint is selected by assuming that the OTAT trip is bypassed, and the reactor coolant system conditions proceed to and beyond the low l pressurizer pressure setpoint, 1845 psig.

The DNBR is calculated for a core inlet pressure of 1800 psi, which is arbitrarily conservative since core pressure will be l greater than pressurizer pressure during virtually all forced i circulation conditions. Reactor power, flow, and coolant l temperatures are taken at the nominal values. The LYNXT statepoint calculation shows that the minimum DNBR for this event using the BWCMV CHF correlation (Reference 6-6) will be greater than 1.9 and well above the design limit. Because this event is not limiting from the fuel reload standpoint, and is relatively insensitive to the core-related parameters that are checked for other, more limiting-transients, no cycle-specific analyses l

,s should be required for reload safety evaluations with

[ }

FCF-designed fuel.

6.2.13 Accidental Depressurization of the Main Steam System l l

The spurious opening or failure of a steam generator relief, I safety, or steam dump valve represents the most severe overcooling event of frequency associated with accidental depressurization of the secondary system. The increased steam flow resulting from the relief, safety, or dump valve failure can be sufficient to cause a reduction in coolant temperature and pressure. This cooldown could produce a positive reactivity insertion and power increase that could challenge the fuel thermal limits. More detailed descriptions of this transient and the systems that normally provide protection against accidental depressurization of the secondary system are presented in Section i 15.2.13 of the reference SQN FSAR.

The bounding analysis presented in the reference FSAR assumes initiation of the event from an end-of-life, shutdown condition and considers the most reactive RCCA to be stuck in its fully withdrawn position.

The core-related parameters that determine the potential return-to-power, the effective moderator reactivity, and power feedback coefficients, are taken at maximum and minimum values,

{A

\ 6-33

I i

l Non Proprietary 1

respectively. The assumed steam release represents the maximum capacity of any single steam dump, relief, or safety valve. l The results of the FSAR analysis show that the acceptance i criteria relative to spurious opening of a steam generator relief l or safety valve are met. In particular, the bounding case remains within the DNB design limits. Under the same set of 1 assumptions and initial conditions, this event is less limiting '

than the main steam line failure.

FCF analyzes the MSLB accident to meet Condition II acceptance I criteria. Because the plant response (e .g. reactivity insertion, i core power and minimum DNBR) would be more severe for a MSLB as l compared with this event, the consequences of an accidental depressurization of the main steam system are bounded by the MSLB analysis presented in Section 6.4.1. Because the Condition II acceptance criteria are met for a MSLB with FCF-supplied fuel, all applicable acceptance criteria are met for an accidental i depressurization of the main steam system with FCF-supplied fuel.

l 6.2.14 Spurious Operation of the Safety Injection System at 1 Power l An error by the operator or a false actuation signal could produce spurious operation of the emergency core cooling system during full power operation. Westinghouse-designed plants typically provide for reactor and turbine trip signals immediately following a safety injection actuation. These would normally produce immediate termination of this event.

In the absence of these trip signale, actuation of the safety  !

injection system will result in delivery of highly borated water to the reactor coolant system. While both high- and low-head safety injection trains are actuated, the injection flow is from  !

the high-head system only because the reactor coolant system is at normal operating pressure.

The effect of the injected boron is a negative reactivity l excursion and decrease in reactor power. The reactor coolant system temperature and pressure decrease in direct response to the decrease in neutron and thermal power. For the bounding i cases developed in the FSAR, the normal direct reactor and turbine trips are assumed bypassed, and the transient is terminated by reactor trip on low pressurizer pressure. Because the power and coolant temperature decrease throughout the l transient, and the reactor coolant conditions are following the l power reduction, DNBR increases throughout the event. I Evaluation of the spurious ECCS actuation event as represented in the reference SON FSAR indicates that operation with Mark-BW fuel will not affect the consequences of the transient. While the  !

6-34 Ol l

1 l

Non Proprietary i

gs\ results can be affected by variations in the reactivity coefficients assumed in the analyses, the DNBR behavior is far l \_/s removed from fuel thermal design limits.

The core-related parameters that define the bounding cases for this event, the Doppler and moderator reactivity coefficients at lower bound values, will be checked for other Condition II transients and need not be specifically confirmed for this event.

Table 6.2-1 Sequence of Events for an Uncontrolled RCCA Bank Withdrawal at Power l

Event Time (sec.)

Initiation of uncontrolled RCCA 0.0 withdrawal at a high reactivity insertion rate (75 pcm/sec).

Power range high neutron flux high 2.20 setpoint reached.

Rods begin to fall into core. 2.70 t

7-s Minimum DNBR occurs. 2.80 (J

R.

\

Table 6.2-2 Sequence of Events for Loss of Electrical Load Without Automatic Pressure Control i Event Time (sec.)

l Loss of electrical load. 0.0 j 1

High pressurizer pressure reactor 5.1 reactor trip set point reached.

Rods begin to drop. 7.1 Minimum DNBR occurs. 8.6 Peak pressurizer pressure reached. 9.8 Main steam safety valves lift. 10.4 l j Peak secondary pressure reached. 16.0 i

-~s f( 6-35

Non Proprietary Figure 6.2-175 PCM/s Rod Withdrawal 2.4 Neutron Power j i 2.0- -

4--

i i i i i_ _

+

2 }  ;

i I  !  !

1.6 - - - -

• I- -I-- -&

--+

e h 1

.2 -/

E -

2 0

0.8 - -+- -t-I i i~

e v  !

0.4 ' I

- - - -+i ---------4

~ ~ ~+ ---

i O.0 '

O 5 10 15 20 25 30 TIME. SECONDS Figure 6.2-2 75 PCM/s Rod Withdrawal 1.2 Thermal Power 1.0- - - - -

1 - - ~ ~ ---

z l

O.8 - - - - - -

~

s u.

0.6 - -- I 2

.4 0.4 - -

e 0.2 - + -+---- -

-+- ----

1 0.0 '

l O 5 10 15 20 25 30 TIME, SECONDS 6-36

Non Proprietary Figure 6.2-3 75 PCM/s Rod Withdrawal 2450 Pressurizer Pressure ,

i I i  !

i 2400-1 2350- ---* -

E 2300- ---

~4-9 --- ----

2250 +

2000- ---

i-1 2150 i 0 5 10 15 l 20 25 30 1 1

TIME. SECONDS i

Figure 6.2-4 75 PCM/s Rod Withdrawal Pressurizer Water Volume 1325 1300 +

E i 1275- j -- -

g 1250- -- '

L-

- - '- bh' I1225-1200-N +

I I

1175 0 5 10 15 20 25 30 TIME, SECONDS t

6-37 l

m 1-- -

• _ ___ --_ _m__-- __ __ _ _ __m_ s-

Non Proprietary Figure 6.2-5 75 PCM/s Rod Withdrawal Core Average Fluid Temperature 592 I

u. I '

588.

Y 584 * ---r-1

$ 580- 4 ~+

d \

=

576- - -4 1\ [--

1 r

4 l i

O 572- I- b- -- F-568 0 5 10 15 20 25 30 TIM 2, SECONDS Figure 6.2-6 75 PCM/s Rod Withdrawal Calculated DNBR 8

7- * *- s- I 6 4 1- 1 1 l S. ,

i 4- -

3-r -

r

!7 2

0 5 10 15 20 25 30 TIME, SECONDS 6-38

Non Proprietary Figure 6.2-7 Loss of Electric Load xiv Neutron Power 48 40' *--- *- --+ --

3 32 --

-+- - -

L 24 -- ---

I 16- r *----

l l

---~~t I {- '

i l 8- - - - - l r ,

i I

0 l s O 4 8 12 16 I l 20 24 i

BMES  !

i i 1 Figure 6.2-8 Loss of Electric Load 3250 Reactor Vessel Lower Plenum Pressure l 3000- --*

}-

2750- - =

2500- +- \

l 2250-

/ ,

p '

%+

2000-O 4 8 12 16 20 24 TIME, S

, l j 6-39 l

Non Proprietary Figure 6.2-9 Loss of Electric Load Steam Generator Downcomer Pressure 1200 , , , ,

! l i [ i  !

! 1120- -

I-1040- -

+- - -r-960< -

b t + -

= l E l l 880- - j -- +- e- -

1 i

800- F -

--+

720 1

i

! l O 4 8 12 16 20 24 TIME, S Fi ure 6.2-10 Loss of Electric Load kverage Core Fluid Temperature 600 595 * +- *-- I

, 590- +-- --+ +

N

\

• j- -

I585-580<

-}- j-

--l m

s 575- -

570 0 4 8 12 16 20 24 TIME S 6-40

l t

r Non Proprietary I

] ,

Figure 6.2-11 Loss of Electric Load t 1

1 Pressurizer Water Volume 1500

+

l I ,

i i

i 1

i

) 1400< -

i i

E i i

, l

) >

,s00 - . - .

L.- . L 3 . . - . .

4

,I 1

1 4

f 1200 1 0 4 8 12 16 20 4

24 1 TIME. S i

1 i Figure 6.2-12 Loss of Electric Load i DNBR h

4 5.0 I

i l

' 1 l

4.5- * -

'- + - - -

i l

4.0 4 . .+. - . . - - .

i i

3.5- - - - - - - 4--

p------+ " - - - - - - -

l; 3.0 - - -

-+ -- e t

1 1.s. ~

T~'~'~~~

3 2.0 N

{- 0 2 4 6 8 10 12 4

4 ,

TIME. S 4

i j

i.

6-41 k

i d

Non Proprietary 6.3 Condition III - Infrequent Faults By definition Condition III occurrences are faults which may occur very infrequently during the life of the plant. They will be accommodated with the failure of only a small fraction of the fuel rods. The release of radioactivity will not be sufficient to interrupt or restrict public use of those areas beyond the exclusion radius. A Condition III fault will not, by itself, generate a Condition IV fault or result in a consequential loss of function of the Reactor Coolant System or containment barriers. The Condition III events applicable to SQN are reported in Section 15.3 of the SON FSAR.

6.3.1 Inadvertent Loading of a Fuel Assembly into an Improper Position The arrangement of assemblies with different fuel enrichments in the core will determine the power distribution of the core during normal operation.

The loading of fuel assemblies into improper core positions or the incorrect preparation of the fuel assembly enrichment could alter the power distribution of the core, leading to potentially ,

increased power peaking and possible violation of fuel thermal '

limits.

The following fuel misloadings nave been considered in the SON FSAR:

1. Misloading a fuel pellet or pellets with an incorrect enrichment in a fuel rod.

2. Misloading a fuel rod with an incorrect enrichment in a fuel assembly. )

l

3. Misloading a fuel assembly with an incorrect enrichment or burnable poison rods into the core.

Evaluation Misloading the fuel pins in an assembly is prevented by loading controls and procedures. Each fuel rod is identified by an enrichment code. The manufacturing process relies on administrative procedures and quality control checks to ensure that fuel rods are placed in the proper assembly and proper assembly lattice location. These same loading controls and procedures apply to the placement of gadolinia-bearing fuel rods within the assembly lattice. As a further precaution, gadolinia-bearing fuel rods are normally placed symmetrically in the assembly lattice.

6-42

Non Proprietary fN Gross fuel assembly misplacement in the core is prevented by (N ) administrative core loading procedures and the prominent display

,/ of identification markings on the upper nozzle of each assembly.

Verifications utilized are as follows:

1. Prior to the commencement of loading operations a check is made to ensure that fuel assemblies have been stored in their proper storage locations in the new fuel vault and that any control components are inserted in the proper fuel assemblies.

2. The fuel unload / reload operations are performed in accordance with a predetermined and reviewed load sequence procedure.

3. After fuel loading has been completed, the core loading is verified by visually surveying the core and recording the fuel assembly numbers versus core location. This record is then compared to the core loading plan.

The procedures will specify how the fuel assemblies are to be oriented in the reactor and in the spent fuel pool with respect to fixed references.

,_ After refueling is completed, startup physics testing is

( ) performed at zero and low power levels prior to escalation to

( , / full power. The incore system of moveable flux detectors is used to verify that the core power distributions are consistent with the design. Results of representative analyses presented in the SON FSAR demonstrate that power peaking increases in excess of the uncertainty would be readily detected during startup testing by the incore moveable detector system.

This conclusion will remain valid for the introduction of Mark-BW fuel, which is neutronically similar to the Westinghouse standard fuel. Undetectable peaking perturbations smaller than the uncertainty value would be within the uncertainty allowance of the analyses.

The following conclusions were reached in the reference safety analyses:

o Fuel assembly enrichment errors would be prevented by administrative procedures implemented in fabrication; o In the event that a single pin or pallet has a higher enrichment than the nominal value, the consequences in terms of reduced DNBR and increased fuel and clad temperatures will be limited to the incorrectly lor.ded pin or pins, and perhaps the immediately adjacent pins;

(.

\ l

\s ' 6-43

l Non Proprietary o Fuel assembly loading errors are prevented by administrative procedures implemented during core loading. In the unlikely event that a loading error occurs, resulting power distribution effects will either be readily detected by the incore moveable detector system or will cause a sufficiently small perturbation as to be acceptable within the uncertainties allowed between !

nominal and design power shapes. '

These conclusions will continue to be valid with the insertion of Mark-BW fuel, and no radiological consequences will result from the inadvertent loading and operation of a fuel assembly in an 1 improper position.

6.3.2 Complete Loss of Forced Reactor Coolant Flow A complete loss of forced reactor coolant flow may result from a l simultaneous loss of electrical supplies to all reactor coolant pumps.

If the reactor is at power at the time of the accident, the immediate effect of loss of coolant flow is a rapid increase in l the coolant temperature. This increase could result in DNB with I subsequent fuel damage if the reactor were not tripped promptly. 1 Normal power for the reactor coolant pumps is supplied through buses from a transformer connected to the generator. Each pump is on a separate bus. When a generator trip occurs the buses continue to be supplied by automatic tre sfer to external power lines, and the pumps continue to supply coolant flow to the core.

The following signals provide the necessary protection against a i complete loss of flow accident: j i

1. Reactor coolant pump power supply undervoltage or  !

l underfrequency,

2. Low reactor coolant loop flow.

3. Pump circuit breaker opening.

The reactor trip on reactor coolant pump undervoltage is provided to protect against conditions that can cause a loss of voltage to all reactor coolant pumps, that is, station blackout. This function is blocked below approximately 10% power.

The reactor trip on reactor coolant pump underfrequency is provided to trip the reactor for an underfrequency condition, resulting from frequency disturbances on the power grid. The reactor trip on low primary coolant loop flow is provided to protect against loss of flow conditions that affect only one 6-44

Non Proprietary reactor coolant loop. This function is generated by

,7 's two-out-of-three low flow signals per reactor coolant loop.

! V) 6.3.2.1 Analysis of Effects and Consequences The loss of power to all four reactor coolant pumps at full power I operation is analyzed using two computer codes. The system response to the four-loop flow coastdown is analyzed using RELAP5/ MOD 2-B&W. This calculation determines such primary system parameters as the loop and core flows, the time of reactor trip based upon the calculated flows, the nuclear power transient, and the primary coolant system temperatures and pressures during the event. The LYNXT computer code is then used to calculate the heat flux transient based upon the neutron power and flows from RELAPS, determining the minimum core DNBR as a function of time.

Initial Conditions Initial reactor power, pressure, and RCS temperature in the DNBR analysis are assumed to be at their nominal values.

Uncertainties in initial conditions are included in the limit DNBR as described in Reference 6-3.

Reactivity Parameters I I The least negative Doppler power coefficient of -6.5 pcm/% is

[_s} used. A positive moderator temperature coefficient of +7 pcm/*F l

(' '/ is assumed. In addition, an end of life b ort of 0.0044 is used.

This combination of parameters maximizes the power response and provides a limiting value of minimum DNBR.

Results Figures 6.3-1 through 6.3-5 show the transient response for the loss of power to all reactor coolant pumps with four loops in operation. The reactor is assumed to be tripped on an l undervoltage signal. Figure 6.3-5 shows the DNBR to be always I

greater than the limit value. Key events in the loss of flow event are given in Table 6.3-1.

Since DNB does not occur, the ability of the primary coolant to remove heat from the fuel rod is not greatly reduced. Thus, the average fuel and clad temperatures do not increase significantly above their respective initial values. The reactor coolant pumps will continue to coast down, and natural circulation flow will eventually be established. With the reactor tripped, a stable i

plant condition will eventually be attained.

l

, At no time do reactor coolant system or secondary system pressures exceed the limits of General Design Criterion 15. The analysis conservatively includes margin for stuck rods, per n

(

\ 6-45 l

l

Non Proprietary General Design Criterion 26, while assuring that fuel design limits are not exceeded.

6.3.2.2 Conclusions The analysis performed has demonstrated that for the complete loss of forced reactor coolant flow, the DNBR does not decrease below the limit value at any time during the transient. Thus, no fuel or clad damage is predicted. Furthermore, primary and secondary system pressures remain well below the design values.

Consequently, all applicable acceptance criteria are met.

6.3.3 Single Rod Cluster Control Assembly Withdrawal at Full Power This event is evaluated in Section 6.2.3.

Table 6.3-1 Sequence of Events for Complete Loss of Reactor Coolant Flow Event Time Isec.)

All operating pumps lose power and O.

begin coasting down.

Reactor coolant pump undervoltage 0.

trip point reached.

Rods begin to drop. 1.5 Minimum DNBR occurs. 3.4 Peak pressurizer pressure occurs. 6.1 Peak secondary pressure occurs. 19.0 i

)

6-46 ,

i J

....- . .. .-.. - .-.-.-.-.. ~ -. - - -...-. . - - - - - - . . - - . - . - - - -- - ...-.-... .

Non Proprietary Figure 6.3-1 Four Pump Coastdown 1.2 Neutron Power  !

1.0 A i

+-- i i - +

0.8 - -- b A l -- b .

a 0.6- - -4 - - - - -

-- d-- --

0.4 - - - - -

0.2-

--} f-- -- -..y 4 f

J  !

0.0 0 5 10 15 20 25 30 TIME, SECONDS 1

Figure 6.3-2 Four Pump Coastdown Thermal P0wer 1.2 1.0 *----- ----

t 0.8 - - -

0.6 -

0.4 - ~~ --

0.2- --- - .

0.0 0 5 10 15 20 25 30 TIME, SECONDS 6-47

Non Proprietary Figure 6.3-3 Four Pump Coastdown i Core Flow Fraction I  !  ! I l  ! I 1.0 { f- *  ! 1 0.8- ----* * ,

I 0.6- i- - + ~

l l

E h 80.4- -+- r O.2- b p -- d 7

I 0.0 0 5 10 15 20 25 30 TIME, SECONDS l

3200 Figure 6.3-4 Four Pump Coastdown Pressurizer Pressure Oi 3000- + +

E 4- +- 1 g2800-S 2600- ' d I

m 2400-

-4 s +-

2200- + --

lN

+- 4- +

2000 0 5 10 15 20 25 30 TIME, SECONDS 6-48

4_ _ . _ - _ _ _ _ _ _ _ _ _

i ai Non Proprietary j' -

Figure 6.3-5 Four Pump Coastdown DNBR 12 i I 4

i i

10< * * +- 4-1 i

e d

8< +

- - +--

},

d 4

4

{ 6 -

/ .

4 i i l -,'

, 4 . ,1 1

i, 2<

t i

O 1

0 2 4 6 8 10 12 i

TIME, SECONDS i

1 i

b t,  !

i t

6 1 l 1 <

4 <

i i

d l

l, 1

I q

e i

i k

i 2

1 i

s j 6-49

i 1

i i .

i

+ l

Non Proprietary 6.4 Condition IV - Limiting Faults Condition IV occurrences are faults which are not expected to take place, but are postulated because their consequences would include the potential for the release of significant amounts of radioactive material. The Condition IV events applicable to SQN are presented in Section 15.4 of the SQN FSAR.

6.4.1 Rupture of a Main Steam Line The steam release arising from a rupture of a main steam line would result in an initial increase in steam flow that decreases during the accident as the steam pressure falls. The energy removal from the RCS causes a reduction in both coolant temperature and pressure.

In the presence of a negative moderator temperature coefficient, the cooldown results in an insertion of positive reactivity. If the most reactive RCCA is assumed stuck in its fully withdrawn position after reactor trip, there is an increased possibility that the core will become critical and return to power. A return to power following a steam line rupture could challenge fuel thermal limits because of the high power peaking factors consistent with assuming the most reactive RCCA to be stuck in its fully withdrawn position. The core is ultimately shut down by the boric acid injection delivered by the safety injection system.

The analysis of a main steam line rupture is performed to demonstrate that the following criteria are satisfied:

1. Assuming a stuck RCCA, and assuming a single failure in the engineered safety features, the core remains in place and intact. Radiation doses do not exceed the guidelines of 10CFR100.

2. Although DNB and possible clad perforation following a steam pipe rupture are not necessarily unacceptable, the following analysis, in fact, shows that no DNB occurs for any rupture assuming the most reactive assembly stuck in its fully withdrawn position. The DNBR design limit is based on the statistical core design methodology described in Reference 6-3.

The MSLB is the most limiting cooldown transient. It is analyzed at zero power and with no decay heat since decay heat would retard the cocidown, thereby reducing the return to power. The following functions provide the protection for a steam line rupture:

6-50

Non Proprietary

1. Safety injection system actuation from any of the following:

a. Two-out-of-three low pressurizer pressure signals.

b. Two-out-of-three low steam line pressure signals in any one loop.

d. Two-out-of-three high containment pressure signals.

2. The overpower reactor trips (neutron flux and AT) and the reactor trip occurring in conjunction with receipt of the safety injection signal.

3. Redundant isolation of the main feedwater lines.

Sustained high feedwater flow would cause additional cooldown. Therefore, the safety injection signal will generate a feedwater isolation signal that will rapidly close all main feedwater regulating valves, the regulating bypass valves and isolation valves.

4. Trip of the fast-acting main steam isolation valves (designed to close in less than 5 seconds) on:

a. Two-out-of-three low steam line pressure signals

(}

d in any one loop.

b. Two-out-of-four high-high containment pressure signals.

c. Two-out-of-three high steam line pressure rate signals in any one loop.

The main steam isolation valves will fully close within 10 seconds of a large break in the steam line. For breaks downstream of the isolation valves, closure of all valves would completely terminate the blowdown. For any break, in any location, no more than one steam generator would experience an uncontrolled blowdown even if one of the isolation valves fails i to close. i Steam flow is measured by monitoring dynamic head in nozzles located inside the steam pipes. The effective throat area of the nozzles is 1.4 f t*, which is considerably less than the main-steam pipe area. Thus, the nozzles also serve to limit the maximum steam flow for any break further downstream.

O

, 6-51

l l

Non Proprietary j l

6.4.1.1 Analysis of Effects and Consequences l Four separate cases were analyzed to identify the limiting combination of break location and reactor coolant pump status.

These were: ,

1

a. Double-ended rupture of a steam line downstream of the steam measurement device with offsite power available.

For this case the effective break area on the affected 2

steam generator is 1.4 ft,

b. Double-ended rupture of a steam line upstream of the steam measurement device with offsite power available.

For this case the effective break area on the affected steam generator is 5.326 ft 2,

c. Double-ended rupture of a steam line downstream of the steam measurement device with loss of offsite power (coolant pumps trip).

d. Double-ended rupture of a steam line upstream of the steam measurement device with loss of offsite power.  !

The limiting case was determined to be case b, a double-ended rupture upstream of the steam measurement device with offsite power available.

The analysis of the limiting steam pipe rupture, with offsite power available, was performed to determine:

1. The core heat flux and RCS temperature and pressure  !

resulting from the cooldown following the steam line I break. The RELAPS code is used to calculate system l response.

2. The thermal and hydraulic behavior of the core ,

following a steam line break. The LYNXT l thermal-hydraulic analysis code is used to determine whether DNB occurs for the core inlet conditions calculated by RELAPS in conjunction with the localized power peaking predicted by NEMO.

Based upon the bounding cases defined for the reference FSAR, the following conditions were assumed to exist at the time of the MSLB accident:

1. End-of-life shutdown margin at no-load, equilibrium xenon conditions, and the most reactive RCCA stuck in its fully withdrawn position. Operation of the control rod banks during core burnup is restricted in such a way that addition l 6-52

Non Proprietary f

~'s of positive reactivity in a steam line break accident will not lead to a more adverse condition than the case analyzed.

( }

i 2. The negative moderator coefficient corresponding to the

! end-of-life rodded core with the most reactive RCCA in the fully withdrawn position. The variation of the coefficient with. temperature and pressure has been included. The Krr l versus temperature corresponding to the negative moderator i coefficient used is shown in Figure 6.4-1. The effect of power generation in the core on overall reactivity is shown in Figure 6.4-2.

The core properties associated with the sector nearest the I

affected steam generator and those associated with the remaining sector are conservatively combined to obtain average core properties for reactivity feedback calculations. Further, it is conservatively assumed that the core power distribution is uniform.

These two conditions cause underprediction of the reactivity feedback in the high power region near the stuck rod. To verify the conservatism of this method, the reactivity and the power distribution are checked for the limiting statepoints for the cases analyzed.

The core analysis considers the Doppler reactivity from t.he l

[s high fuel temperature near the stuck RCCA, moderator feedback from the high water enthalpy near the stuck RCCA, power redistribution and nonuniform core inlet temperature effects. For cases in which steam generation occurs in the high flux regions of the core, the effect of void formation is also included. A 3-D statepoint analysis is used to confirm that the reactivity employed in the kinetics analysis was larger than the reactivity would be when

! calculated including the above local effects for the statepoints. These results verify conservatism; that is, underprediction of negative reactivity feedback from power generation. The 3-D statepoint is evaluated for each cycle to confirm the conservatism.

3. Minimum capability for injection of boric acid (assuming 1950 ppm from the RWST for this analysis) solution by the safety injection system. The emergency core cooling system consists of three systems: 1) the passive accumulators, 2) the residual heat removal system, and 3) the safety injection system. Only the safety injection system is l modeled for the steam line break accident analysis.

l High pressure safety injection is modeled to start 28 seconds after a safety injection signal is received. This delay accounts for pump start time, signal delays, and valve ks- 6-53 I

l l

i

Non Proprietary opening time. To minimize boron transport to the core, a single failure of one safety injection train is assumed.

Therefore, flow is provided by only one train (Figure 6.4-3). Furthermore, before boron could arrive in the reactor coolant system, the volume of water between the cold leg piping and the first check valve in the injection line (assumed to be at 0 ppm) had to be purged.

4. Although the SQN plant feeds the steam generators with auxiliary feedwater at zero power, main feedwater was introduced to the steam generators at the full power flow rate until the main feedwater isolation valves were completely closed. Auxiliary feedwater flow of 2350 gpm is started to the faulted steam generator at the time of main feedwater isolation.

5. Power peaking factors, corresponding to one stuck RCCA and nonuniform core inlet coolant temperatures, are determined for end-of-core-life from the 3-D statepoint analysis already described. The coldest core inlet temperatures are assumed to occur in the sector with the stuck rod. The power peaking factors account for the effect of the local void in the region of the stuck control assembly during the return to power phase following the steam line break. This void in conjunction with the large negative moderator coefficient partially offsets the effect of the stuck assembly. The power peaking factors depend upon the core power, temperature, pressure, and flow. Calculation of the l peaking factors is repeated for each cycle and MDNBR is 1 checked.

6. The analysis assumes initial hot standby conditions at time zero, since this represents the most pessimistic initial condition. Should the reactor be just critical or operating at power at the time of a steam line break, the reactor will be tripped by the normal overpower protection system when power level reaches a trip point. Following a trip at power, the RCS contains more stored energy than at no-load, '

the average coolant temperature is higher than at no-load, and there is appreciable energy stored in the fuel.

Thus, the additional stored energy is removed via the cooldown caused by the steam line break before the no-load conditions of RCS temperature and shutdown margin assumed in the analyses are reached. After the additional stored energy has been removed, the cooldown and reactivity insertions proceed in the same manner as in the analysis which assumes no-load condition at time zero.

6-54

. - - - - . . _ ~ ~ _ - . - . - . - . _ - _ . - . - -- _ . _ - . . . . - - -

Non Proprietary Results The results presented are a conservative indication of the events that would occur assuming a steam line rupture, since it is postulated that all of the conditions described above occur simultaneously. The sequence of key events for the limiting steam line rupture case is given in Table 6.4-1.

Figures 6.4-4 through 6.4-17 show the system and core response to the double-ended rupture of a main steam line, upstream of the i

flow measurement devices, from an initial no-load condition.

! Offsite power is assumed available so that full reactor coolant flow exists. Unrestricted blowdown of the affected steam generator (effective break area 5.326 f t*) is shown. Prior to main steam isolation, the other three steam generators blow down through a single 1.4 ft2 steam measurement orifice.

As shown in Figure 6.4-10, the core attains criticality with the f RCCAs inserted (with the design shutdown of 1.6% Ak/k, assuming l

one stuck'RCCA) in about 32 seconds. The effect of Doppler l

feedback results in a peak core power significantly lower (Figure l 6.4-4) than the nominal full power value.

)

l

'- It should be noted that following a steam line break only one steam generator blows down completely. Thus, the remaining steam generators are still available for dissipation of decay heat i after the initial transient is over.

(

A DNB analysis was performed using the W-3 CHF correlation for the limiting core condition during the event (Table 6.4-2). It was determined that the minimum DNBR remains greater than the limit value. The W-3 CHF correlation has been evaluated for the specific range of application and found to be acceptable.

Because there is no departure from nucleate boiling, the fuel and cladding temperatures cannot exceed the design limits. Thus, the 1 relevant requirements of General Design Criterion 10, as they relate to maintenance of fuel and cladding integrity, are met for all increased heat demand events.

6.4.1.2 Environmental Consequences

The postulated accidents involving release of steam from the secondary system do not result in a release of radioactivity unless there is leakage from the RCS to the secondary system in the steam generators. The FSAR bounding analysis of the potential offsite doses resulting from this accident is based upon parameters and bounding assumptions not affected by operation with Mark-BW fuel. l j

l l

6-55 I

Non Proprietary 6.4.2 Major Rupture of a Main Feedwater Pipe Major rupture of a main feedwater line represents an immediate decrease in the heat removal capability of the secondary system because it reduces the supply of feedwater to the steam generators. As considered in Section 15.4.2.2 of the SON safety analysis, main feedwater is assumed to be lost to all of the steam generators at the time of rupture. Reverse blowdown of the i affected steam generator results in a relatively rapid reactor trip signal on low-low level in that generator.

)

The post-trip transient, then, is marked by excess heat removal l from and cooldown of the primary coolant system as the affected steam generator continues to blow down through the break. When the low steamline pressure setpoint is reached, the main steam isolation valves are closed, and safety injection to the primary coolant system is initiated. The loss of steam generator inventory and rising steam pressure cause primary temperatures to rise. Successful termination of the transient is achieved when the auxiliary feedwater supplied to the steam generators is sufficient to remove core decay heat. The main feedwater line break is a design basis event, analyzed to demonstrate overpressure protection of the reactor coolant system and continued capability for core cooling.

Evaluation of the main feedline rupture event for the purposes of operation of the SON unit with B&W reload fuel need only consider those aspects of the transient that could be affected by differences in fuel and fuel cycle design. The significant assumptions, conditions, and system-related phenomena established for the reference safety analysis should not be affected. So long as the decay heat levels--and the core parameters related to decay heat during the transient--are within the bounding values represented in the reference analysis, the acceptance criteria applicable to the main feedwater line break transient will continue to be met for the reload cores.

6.4.3 Steam Generator Tube Rupture The steam generator tube rupture is a design basis accident that considers postulated failure of a single steam generator tube.

Under such conditions, the reactor coolant system depressurizes via flow through the ruptured tube to the affected steam generator.

The reactor is tripped, main feedwater flow is isolated, and the safety injection system is actuated on the low pressurizer pressure reactor protection signal. The primary system event is effectively terminated when makeup flow via the safety injection 6-56

l Non Proprietary f

e~x systen matches the rate of coolant loss through the failed steam

\ generator tube. The tube leakage is terminated when the operator

(

As _j/ depressurizes the primary system below the steam pressure of the affected steam generator.

As postulated for the SON FSAR, the steam generator tube rupture accident is analyzed to establish bounding environmental doses.

The bounding assumptions are those aimed at maximizing the concentrations of fission products in the reactor coolant and secondary inventories, and bounding the mass flow via the faulted steam generator. Evaluation of the steam generator tube rupture event indicates that the limiting consequences as developed for the reference SQN safety analysis are not affected by operation with Mark-BW reload fuel. Therefore, no cycle-specific evaluation is required for future reload cycles.

l 6.4.4 Single Reactor Coolant Pump Locked Rotor l 1

The accident postulated is an instantaneous seizure of a reactor coolant pump rotor. The subsequent loss of reactor coolant flow i results in a reactor trip and RCCA insertion. Loss of offsite '

power is assumed to occur coincident with reactor trip. Upon a l loss of offsite power, the three intact reactor coolant pumps begin to coastdown.

,~,x Heat stored in the fuel rods continues to be transferred to the h coolant causing the coolant to expand. At the same time, heat

[

t ) transfer to the shell side of the steam generators is reduced,

' first because the reduced flow results in a decreased tube side film coefficient and then because the reactor coolant in the tubes cools down while the shell side temperature increases (turbine steam flow is reduced to zero). The rapid expansion of the coolant in the reactor core, combined with reduced heat transfer in the steam generators, causes an insurge into the pressurizer and a pressure increase throughout the reactor coolant system.

The increase in reactor coolant system pressure would normally result in automatic actuation of the pressurizer spray system, opening of the power-operated relief valves, and, if necessary, opening of the pressurizer safety valves. However, the pressurizer spray and power-operated relief valves are not modeled in the analysis to provide a conservative pressure prediction. A conservative calculation of the hot pin minimum DNBR is made by holding the pressure constant at the initial value.

6.4.4.1 Analysis of Effects and Consequences The reactor coolant pump shaft seizure transient is analyzed for one reactor coolant pump rotor locked with four loops in p

t

\

\s_. 6-57

Non Proprietary operation. The RELAPS/ MOD 2-B&W computer code is used to calculate the resulting loop and core flow transients following the pump seizure, the time of reactor trip based on the loop flow transients, the nuclear power following reactor trip, and the peak pressure. The thermal behavior of the fuel located at the core hot spot is determined by applying the SCD methodology of Reference 6-3 using the RELAP5/ MOD 2-B&W predictions of core power l and flow as inputs.

A number of cases were analyzed assuming various combinations of reactivity coefficients. The limiting fuel thermal response was obtained for a moderator temperature coefficient of +7 pcm/ F and Doppler power coefficient of -6.5 pcm/%.

Results The transient results are shown in Figures 6.4-19 through 6.4-24.

The peak reactor coolant system pressure reached during the transient is less than the design value. Figure 6.4-24 shows that the minimum DNBR in the hot channel is 1.121, which is less than the limit of 1.5. Consequently, a clad temperature excursion of short duration is predicted (Figure 6.4-23). The peak cladding temperature is 1104 F. The peak fuel temperature is 3264 F. Less than 5% of the fuel pins experience DNB during l the accident. This result is bounded by the original licensing I analysis. The sequence of events is presented in Table 6.4-3.

6.4.4.2 Conclusions l 1

1. Since the peak reactor coolant system pressure reached during any of the transients is less than that which would cause stresses to exceed the faulted condition stress I limits, the integrity of the primary coolant system is assured.

2. Since the peak fuel temperature is well below the 5080 F fuel temperature limit and the peak cladding temperature is well below the 1800 F cladding temperature limit, the core will remain in place and intact with no consequential loss of core cooling capability.

6.4.5 Rupture of a Control Rod Drive Mechanism Housing (Rod Cluster Control Assembly Ejection)

This accident is defined as the mechanical failure of a control rod mechanism pressure housing resulting in the ejection of a RCCA and drive shaft. The consequence of this mechanical failure is a rapid positive reactivity insertion together with an adverse core power distribution, possibly leading to localized fuel rod damage.

6-58

Non Proprietary

<-s s 6.4.5.1 Analysis and Evaluation J

[

l \ , ,)\ The SQN safety analyses applicable to a current cycle contain l sufficient information to verify that the designed cycle is l within the applicable bounds for the ejected rod accident. The l fuel properties of Mark-BW fuel are very similar to Westinghouse standard and VANTAGE SH fuel. The system response and hot spot analyses are dependent upon the neutronics characteristics and thermal response of the fuel. Neutronics parameters for Mark-BW fuel are within the current bounding physics parameters in the FSAR. These values are calculated with NEMO.

The thermal response of Mark-BW fuel to an ejected rod power excursion is calculated using a representative core average nuclear power excursion shown in Figure 6.4-25. The transient power is analyzed by LYNXT to calculate the radial power distribution at which the DNBR limit is exceeded for BOC and EOC l at HFP and HZP. From these results the number of pins failed will be estimated for each cycle. Also, the thermal response of previously licensed fuel was compared to Mark-BW fuel, and no  :

discernible differences were found. These results'are shown in Figure 6.4-26.

The steady-state fuel temperatures calculated with TACO 3 (Reference 6-8) for the different fuels also show no discernible

,_s difference. Since the thermal responses of the different fuels

/ are shown to be similar at steady-state and during an ejected rod

( accident, the bounding parameters in the FSARs will be applicable as limits for the reload Mark-BW fuel.

6.4.5.2 Results The calculated ejected rod worths, Fq, delayed neutron fraction, and pin census for each cycle will be shown to be less restrictive than the values presented in the FSAR reference analyses. These values are shown in Table 6.4-4. These results will be demonstrated for full power and zero power at both BOC and EOC.

6.4.6 Steam Line Break at Power With Coincident Rod Withdrawal In September of 1979, IE-79-22 entitled " Qualification of Control Systems" was issued by the NRC identifying a potential unreviewed safety question resulting from Control and Protection Systems <

interactions. One of the postulated scenarios that was identified was the operation of the non-safety grade automatic rod control system following a steam line break inside or outside of containment. The automatic rod control system derives signals i

l l

! \s / 6-59 I l

1

Non Proprietary from the Nuclear Instrumentation System (specifically the excore power range neutron detectors) and the turbine impulse pressure, among other inputs, to determine if control rod motion is required. Since a steam line break may occur inside containment in the vicinity of the excore detectors, or outside containment in the vicinity of the turbine impulse pressure transmitters, the automatic rod control system may be exposed to an adverse environment. This equipment is not qualified to preclude the steam line break from causing a control rod withdrawal due to an adverse environment. In addition to the potential rod withdrawal, the Power Range High Neutron Flux and OTAT reactor protection trip functions may not be available as a result of the harsh environmental conditions which could exist.

6.4.6.1 Analysis of Effects and Consequences Method of Analysis l I

The RELAP5/ MOD 2-B&W computer code was used to analyze the system '

response to a MSLB at power with a coincident withdrawal of control bank D. A spectrum of steam line breaks was analyzed to find the limiting break size with respect to minimum DNBR. The i rod withdrawal was modeled as a 15 pcm per second reactivity )

insertion, which is based on the maximum speed of the rod speed controller (45 inches per minute) and the maximum differential l rod worth of control bank D at hot full power conditions (20 pcm i per inch).

Two reactor trip functions are modeled to mitigate this event.

For large breaks, the low steam line pressure safety injection signal will trip the reactor, trip the turbine and close the main steam isolation valves. For small breaks, the OPAT function will l trip the reactor. l The sensitivity analyses of rod withdrawal at power showed that the limiting DNBR condition was obtained using maximum reactivity feedback coefficients. Consequently, this analysis was performed with the same feedback coefficients, namely, a moderator temperature coef ficient of -45 pcm/ F and a Doppler power i coefficient of -12.5 pcm/%.

Results The spectrum study showed that the minimum DNBR occurs for a break of 0.583 ft . The sequence of events for this case is shown in Table 6.4-5. The calculated plant response is shown in Figures 6.4-27 through 6.4-32. The hot channel DNBR remains above the limit for the duration of the event.

6-60

i Non Proprietary N 6.4.6.2 Conclusions The analysis of steam line break at power with coincident rod withdrawal showed that the OPAT and low steam line pressure safety injection trip functions successfully terminate the event.

No fuel pins experience DNB. Consequently, no fuel damage would occur.

Table 6.4-1 Sequence of Events for Main Steam Line Break Upstream of the Steam Flow Measurement Device With Offsite Power Available Event Time (sec)

Steam line ruptures. 0 Low steamline pressure signal. <2 l 1

( MSIVs closed. 10 l l

l Main feedwater isolated. 22 Safety injection flow begins. 30

/ \ Pressurizer empties. 30

( Criticality attained. 35 Boron reaches the core. 97 i

l l

r l

l l

i I

l l s_s 6-61

Non Proprietary Table 6.4-2 Limiting Core Parameters Used in Steam Line Break DNB Analysis Parameter Value Reactor Vessel Inlet 399.4 F (Faulted Loop)

Temperature RCS Pressure 742.3 psia RCS Flow 106% (of nominal HZP)

Heat Flux 16.4% (of nominal)

Time After Rupture 80.0 seconds Table 6.4-3 Sequence of Events for Reactor Coolant Pump Shaft Seizure (Locked Rotor)

Event Time (sec)

Rotor on one pump locks. O.

Reactor trip on low flow. 0.1 Rods begin to fall. 1.1 Loss of offsite power. Reactor 1.1 coolant pumps trip.

Minimum DNBR occurs. 3.0 l l

Maximum RCS pressure occurs. 4.1 i 1

i 6-62

._ __ _ . _ . . _-.___m_ ___-_-._____.._..__-._-m.-. - _ _ - _ _ _ _ _ . _ -_

Non Proprietary l $(_,\ / Table 6.4-4 Parameters Used in the Evaluation of the Rod l Cluster Control Assembly Ejection Accident

---

Time in Cycle -------

Parameter Begin Begin End End Power level, % 102 0 102 0 Ejected rod worth, %Ak/k 0.20 0.75 0.21 1.01 Delayed neutron fraction, % 0.55 0.55 0.44 0.45 j Fq before rod ejection 2.62 --

2.62 --

L F, af ter rod ejection 7.11 14.05 7.08 22.2 l Fuel failure, % <10 <10 <10. <10 l

l i

Table 6.4-5. Sequence of Events for a 0.583 Ft2 Steam Line s Break With Coincident Control Rod Withdrawal h Event Time (sec)

Steam line break occurs. O.

Control bank D begins to withdraw 0.

causing 15 pcm/sec insertion.

l Reactor trip on overpressure AT. 10.20 Rods begin to fall. 18.20 i

Minimum DNBR occurs. 18.40 i

g 6-63 1

Non Proprietary Figure 6.4-1 Main Steam Line Break With Offsite Power Available K Effective Versus Average Core Fluid Temperature 1.Os -

3 ,

i

' i i 1.04- 4 - -- '

T 1.02 1 --+

L w

1.00 + +- + +

0.9 8 < - - - -

t- t l j t

! I i

l

'8$s0 240 320 400 40 560 640 AVERAGE CORE FLUID TEMPERATURE, F Figure 6.4-2 Main Steam Line Break With Offsite Power Available 0

Doppler Power Feedback Ol

-400' i

800- t- ~ +

i2.1200' x '

1600<

N O 1 2 3 4 5 POWER FRACTION 6-64 O

l Non Proprietary _

Figure 6.4-3 Main Steam Line Break With Offsite Power Available Single Train High Pressure Safety injection Flow

! l l i

1600<

h i

[1200- N \ -

\

800- --

--+ +- - -+- +

l 400- , -

l J- - y '

O O 100 200 300 400 500 600 SAFETY INJECTION FLOW, GPM

)

Figure 6.4-4 Main Steam Line Break With Offsite Power Available Neutron Power O.24 0.20- 4

I l '

l 0.12 -

y V '

) 0.08- - -

i i

0.04-i

)

' 0.00 '

O 20 40 60 80 t 100

% TIME, S

{ 6-65

Non Proprietary Figure 6.4-5 l Main Steam Line Break With Offsite Power Available 2400 Pressurizer Pressure I

i i

2000< --- --t--

t 4 +-

1600- &

-+

1200 -

i 3 800- -

-j

~7 i

r-

! I 400- -- -A -$ -- 4 -

o 0 20 40 60 80 1@

TIME, s Figurc 6.4-6 Main Steam Line Break With Offsite Power Available Pressurizer Water Level 16 b

u.

12< -- -+ - -

d Q

o 8< j +-- -

s - --<

I m

4 --

--+ - -- -

0 -

0 20 40 60 80 100 TIME, S 6-66 O

Non Proprietary Figure 6.4-7 Main Steam Line Break With Offsite Power Available Single Loop Reactor VesselInlet Temperature w HO } ' i I I

' l i

I560- I 480- 1 -

---

+-

N 400- 4 f

5 ,  !

$ 320-in 1

--- - k ~4 4 4 l l 240 '

O 20 40 60 80 100 TIME. S Figure 6.4-8 Main Steam Line Break With Offsite Power Available O]

\

720 Triple Loop Reactor Vessel inlet Temperature

• 640 l

--- 4- --

p I560- N

$ 480- - -

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D 400- 4 320- +

240 0 20 40 60 80 100 TIME. S

\ 6-67

Non Proprietary Figure 6.4-9 Main Steam Line Break With Offsite Power Available 720 Core Average Temperature i +

i 640- - - - - - ~

u.

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480- ---- 1

, ~ -

< 400-  !

, f 320- --

- h-

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9

-' 500-i

- ' /

1000 . . 1 -

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O 20 40 60 80 100 TIME, S 6-68 O

1 1

Non Proprietary Figure 6.4-11 Main Steam Line Break With Offsite Power Available x Core Boron Concentration 120 100-l 80- i I 60-

. BORON ENTERS;THE CORE AFTER 97 SECONDS 40- --

l l l

g f +

l i

I 20- - -

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+

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( Figure 6.4-12 Main Steam Line Break With Offsite Power Available

(-

Single RSG Feedwater Flow 1400 1200- "- -------+ - -+- ~~

l 1000- 6- + - - - ---

800 '

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1 -

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l l

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l V 6-69 l

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l' Non Proprietary Figure 6.4-13 Main Steam Line Break With Offsite Power Available Triple RSG Feedwater Flow 4800 ,

i i 1

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800 ---

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l Figure 6.4-14 Main Steam Lint ? eak With Offsite Power Available xiO' Single RSC Naam Flow '

24 -

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Non Proprietary  !

Figure 6.4-15 Main Steam Line Break With Offsite Power Available  !

Triple RSG Steam Flow Q I l

4000-i 3200, ..

d so 2400' -

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Non Proprietary Figure 6.4-17 Main Steam Line Break With Offsite Power Available Triple RSG Steam Pressure 1200 ,

I i ,

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0 0 20 40 60 30 100 TIME, S Figure 6.4-18 Main Steam Line Break With Offsite Power Available Reactor Coolant Flow 1.2 4 _v h 1.0 - ---F i O.8 -

t

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a l

i Non Proprietary ,

j Figure 6.4-19 Locked Rotor i

jk Neutron Power 1.2 j- l I

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) Thermal Power l 1 1.2 '

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Non Proprietary 1 Figure 6.4-21 Locked Rotor i Core Flow Fraction 1.2 I

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3000- --

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1400- H- - - - - ' * - - - - ' + - - - - ' - - L - --

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Non Proprietary Figure 6.4-27 SLB with RWAP Neutron Power i  !'

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1 TIME, SECONDS Figure 6.4-28 SLB with RWAP Thermal Power ,

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TIME, SECONDS

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Non Proprietary Figure 6.4-29 SLB with RWAP Pressurizer Pressure 2450 l i i 2400' - 't-

1 i

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Figure 6.4-30 SLB with RWAP T average O\

588 584 -

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o i N 576- -j l- '

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, . . - ~ . . - . - . . - . . - - . ~ . . . . . - . . . . . - . . . . . . - . . . . . . . - . _ . . . . . . . -. ---. -.

Non Proprietary Figure 6.4-31 SLB with RWAP 1

Steam Pressures 1200 LEGEND Unbroken loop 1120- -- - + -

- Btoken loop .

1040- --4 I ----4 t-- -- ---

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6-79 1

% \

I l

\ l

Non Proprietary 6.5 Conclusions This section has presented a set of evaluations and analyses for non-LOCA transients applicable to the SON unit as reference for future reload safety evaluations for operation with fuel of the Mark-BW design. The scope included all events delineated in Sections 15.1 through 15.6 of Regulatory Guide 1.70 and as presented in the FSAR for SQN. Also included in the scope is the analysis of the steam line break at power with coincident rod withdrawal. Although not a part of the original SQN licensing basis, this analysis was pereformed to address concerns expressed in IE-79-22.

The analyses and evaluations performed for this section confirm that operation of the SQN unit for reload cycles with Mark-BW fuel will continue to be within the previously reviewed and licensed safety limits. Reanalysis of the transients affected by the fuel reloads demonstrates that the acceptance and design criteria specified in Regulatory Guide 1.70 continue to be met.

The safety evaluations and analyses presented in this section complement other BAW topical reports that describe the Mark-BW fuel assembly design; the mechanical, nuclear, and thermal-hydraulics methods supporting the design; and ECCS codes, methods, and applications. The scope of non-LOCA transients addressed in this report will provide, in conjunction with these oth>r topical reports, a reference for future reload safety evaluations of the SQN unit for operation with FCF-supplied fuel assemblies.

6.6 References i 6-1 BAW-101649, Revision 3, RELAP5/ MOD 2-B&W, An Advanced Computer Program for Light Water Reactor LOCA and Non-LOCA ,

Transient Analysis, October 1992. l l

6-2 BAW-10169P-A, RSG Plant Safety Analysis, October 1989. l 6-3 BAW-10170P-A, Statistical Core Design for Mixing Vane Cores, December 1988. l l

6-4 BAW-10173P, Revision 1, MARK-BW Reload Safety Analysis for Catawba and McGuire, March 1989.

6-5 BAW-10156-A, LYNXT Core Transient Thermal-Hydraulic Program, August 1993.

6-6 BAW-10159P, BWCMV Correlation of Critical Heat Flux in  !

Mixing Vane Grid Fuel Assemblies, May 1986. l 1

6-80 J

i l

i 1

Non Proprietary  ;

6-7 BAW-10180-A, Revision 1, NEMO - Nodal Expansion Method l l- Optimized, March 1993. j l~ l l 6-8 BAW-10162P-A, TACO 3 - Fuel Pin Thermal Analysis Computer 1- Code, October 1989, t

i 4

I t

i j'

k

]~

{.

i ,

i i

i I

4 s

5

+

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i 6-81 I i

Non Proprietary f- s 7.0 Thermal-Hvdraulic Evaluation

~

( )

G' This section provides a summary of the thermal-hydraulic analysis methods and models used by Framatorae Cogema Fuels (FCF) to support the licensing of the Mark-BW fuel design for operation at the Tennessee Valley Authority's (TVA) Sequoyah Nuclear Plant (SON). The purpose of the thermal-hydraulic analysis is to establish safety and operating limits that insure fuel and clad  :

integrity are maintained during normal operation and transients of moderate frequency. The design criteria that have been established to meet this goal are as follows:

i 1) During Condition I and II events, there must be at l least a 95 percent probability with a 95 percent I confidence level that the hot pin will not experience a departure from nucleate boiling (DNB); or a 99.9%

probability that DNB will not occur core wide. j

2) During Condition I and II events, there must be at least a 95 percent probability with a 95 percent confidence level that no fuel rod will experience centerline melting.

This section describes the thermal-hydraulic analyses that are performed to show that these criteria are met.

I_h

( / 7.1 Thermal-hydraulic core Models To perform the various thermal-hydraulic analyses needed to license the Mark-BW design, FCF uses the LYNXT thermal-hydraulic analysis code. LYNXT, a single-pass code, employs crossflow methodologies to evaluate subchannel thermal-hydraulic conditions for both steady-state and transient conditions. A more complete description of LYNXT is provided in Appendix A.2 and in the NRC-approved topical report BAW-10156A (Reference 7-1).

7.1.1 LYNXT Modeling As part of the Mark-BW thermal-hydraulic analysis task, LYNXT models of the Mark-BW assembly and the SON core have been developed. The methods that are used to define the core and assembly geometries in the LYNXT code are documented in Reference 7-1. The following paragraphs provide descriptions of the typical models being used in the Mark-BW thermal-hydraulic analysis. The models described include the 31-channel bundle-by-bundle hydraulic model and the 12-channel DNB model. Figures 7.1-1 and 7.1-2 depict the layout of each of these models.

Figure 7.1-1 shows the channel numbering scheme and the eighth core symmetry used in defining the bundle-by-bundle model of the 193 fuel assembly core. Figure 7.1-2 shows the eighth core 7-1

Non Proprietary symmetry and the channel layouts used in the 12-channel DNB model. As shown on these figures, LYNXT models the core with a group of channels of varying sizes. These channels increase in size from individual subchannels, to a group of subchannels, to a group of bundles. By using this variable-scaling method it is possible to model the entire core, while having a detailed subchannel model of the area around the hot subchannel.

For the DNB analysis of the Mark-BW fuel assembly, the applicable CHF correlation is BWCMV as documented in the NRC-approved topical reports BAW-10159P and BAW-10189P (References 7-7 and 7-13, respectively). In its original development (BAW-10159P),

BWCMV was based on an extensive set of inconel mixing grid CHF test data. However, in August 1993, FCF submitted additional test data (BAW-10189P) that showed that the Mark-BW zircaloy mixing grid performed at a level that was superier to the original data base. This enhanced CHF performance of the Mark-BW mixing grid is incorporated into the thermal-hydraulic analysis through the use of design-specific equivalent grid spacing. When using the BWCMV correlation in this manner, referenced specifically to BAW-10189P, it is referred to as BWCMV-A.

7.1.2 Form Loss Coefficients In addition to modeling the assembly and core geometry, it is necessary to model the hydraulic characteristics of the assemblies and subchannels using form loss coefficients. The Mark-BW grid form loss coefficients were originally developed from a series of flow tests performed in the Control Rod Drive Line Facility (CRDL) at the B&W Alliance Research Center (ARC).

The CRDL is a closed loop flow test facility that can produce coolant conditions representative of those occurring during reactor operation. Form loss coefficients for the assembly

]

nozzles and spacer grids were developed by measuring the pressure i drop across the various components. Additional confirmatory flow testing has also been performed with the FCF's Transportable Flow Test Rig (TFTR). The combined results from these tests form the basis for the current component form loss coefficient set.

Subchannel form loss coefficients were determined analytically from the total spacer grid form loss coefficients. These grid and subchannel form loss coefficients are used in LYNXT to model the fuel assembly flow characteristics.

7.1.3 Core Power Distribution The final step in the model development is to define a reference core power distribution that will conservatively bound the actual power distributions occurring during normal operation. The peaking conditions listed below define the limiting peaking factors associated with the reference power distribution being used for the SON core thermal-h'/d raulic design analyses:

7-2

__ _m_ _ _ .. _ _ _ . - - _ _ . - - _ _ _ _ _ _ _ _ _ . _ __

I l

Non Proprietary l

Radial Peak (Fj) = 1.64 ks _s/ Axial Peak (F,)

= 1.55 Axial Peak Location = 0.5 x/L (steady-state)

= 0.7 x/L (transient) l Bundle Average Peak = 1.557 where x/L = normalized axial location along the heated length

( It should be noted that the radial peak of 1.64 corresponds to a l

maximum allowable radial peak of 1.70 when a 4% total rod power

uncertainty factor is included.

l l 7.1.4 Core Conditions A summary of general core conditions used in the SON thermal-hydraulic analyses is provided on Tables 7.1-1 and 7.1-2.

7.1.5 Engineering Hot Channel Factors

! Engineering hot channel factors (HCF's) are penalty factors that l are used to account for the effects of manufacturing variations l on the maximum linear heat generation rate and enthalpy rise.

r 7.1.5.1 Local Heat Flux Engineering Hot Channel Factor

( _j / The local heat flux engineering hot channel factor, F$, is used in the evaluation of the maximum linear heat generation' rate, i This factor is determined by statistically combining manufacturing variances for pellet enrichment and weight and has a value of 1.03 at the 95% probability level with 95% confidence.

As discussed in References 7-2 and 7-3, relatively small heat flux spikes such as those represented by F$ have no effect on DNB, therefore this factor is not used in DNBR calculations.

7.1.5.2 Average Pin Power Engineering Hot Channel Factor l The average pin power factor, Fj, accounts for the effects of

! variations in fuel stack weight, enrichment, fuel rod diameter, and pin pitch on hot pin average power. This factor, which has a value of 1.03, is combined statistically with other uncertainties

to establish the statistical design limit (SDL) DNBR used with t

the statistical core design method (discussed in Section 7.2).

Since Fj is incorporated into the statistical design limit (SDL), this factor is not included in the LYNXT model used for

. SCD analyses. For non-SCD analyses, Fj is incorporated into the

LYNXT model as a multiplier on the hot pin average power.

i l M 7-3 1

i i

l'

1 1

Non Proprietary i 7.1.6 Fuel Rod Bowing The bowing of fuel rods during reactor operation has the  ;

potential to affect both local power peaking and the margin to l DNB. As discussed in Section 4.1.1.7 of BAW-10172P (Reference 7- '

4), the Mark-BW fuel design has several features that make its l fuel rod bow performance similar to that of other FCF fuel  ;

designs. In BAW-10186P (Reference 7-17), FCF presented new data that extended the rod bow data base for FCF fuel to 58,300 i mwd /mtU. The topical report concluded that the rod bow l correlations from BAW-10147PA-R1 (Reference 7-5) are applicable j at extended burnups and apply to the Mark-BW. Using that i prediction, no DNBR penalty due to rod bowing is applied to the l Mark-BW, [ ]. l l

7.1.7 Active Fuel Stack Heicht .

The active fuel stack height varies during reactor operation due to the combined effects of fuel densification, swelling, and thermal expansion. Densification, which acts to shrink the stack, occurs predominantly at low fuel burnup values, while swelling, which increases stack height, predominates at higher burnups. For high density fuel, such as the Mark-BW fuel design ,

to be used in the SON core, stack shrinkage due to densification l is less than the increase caused by thermal expansion of the fuel pellets upon initial heatup; therefore stack shrinkage is not considered in thermal-hydraulic analysis models. The active fuel ,

height in these models is conservatively assumed to be the nominal initial stack height.

7.1.8 Reactor Coolant Flow Rate The statistical core design (SCD) method, discussed in Section 7.2, incorporates uncertainties associated with the reactor core coolant flow into the overall DNBR uncertainty, as represented by the SDL. Calculations performed with the SCD method therefore use a core coolant flow rate that is equal to the nominal thermal design flow rate, less the core bypass flow fraction. Non-SCD calculations, which do not account for the flow measurement uncertainty, use the minimum thermal design flow rate and the l maximum core bypass flow fraction. These calculations were i performed using a 7.0% bypass flow, which was verified as being conservative through SQN-specific plant analyses. The minimum i thermal design flow is equal to the nominal thermal design flow l less the measurement uncertainty allowance.

1 7-4 O

l Non Proprietary A Table 7.1-1 Thermal-Hydraulic Analysis Design Parameters Design Parameter Value Core Configuration:

Number of Fuel Assemblies 193 l \

Fuel Assembly Type 17x17 l Number of Fuel Rods Per Assembly 264 l Number of Control Clusters 53 1

Number of Absorber Rods per Control Cluster 24 l l

l i l l Reactor Coolant System:

l Rated Thermal Power, MWt 3411 Heat Generated In Fuel, % 97.4 V Nominal-System Pressure, psia 2280 Nominal Thermal Design Flow, gpm 360,100 l Flow Fraction Effective for Heat Transfer 0.93 l (7.0% Bypass) l Minimum Thermal Design Flow, gpm 348,000 Mechanical Design Flow, gpm 410,000 Hot Channel Core Inlet Flow Factor 0.95 Average Vessel Coolant Temperature (nominal) at 100%RTP, F 578.2 t

I 4

I j 7-5 4

J 1

i . . . . - _ _ , . _ -

Non Proprietary Table 7.1-1 (cont'd)

Thermal-Hydraulic Analysis Design Parameters Design Parameter Value Power Distribution:

Reference Relative Assembly Power at 100%RTP 1.557 Reference Relative Pin Power at 100%RTP 1. 64 t u Axial Flux Shape 1.55 Chopped Cosine x/L = 0.5 steady-state x/L = fraction of heated length x/L = 0.7 transient DNBR Calculations:

CHF Correlation BWCMV-A for MK-BW BWCMV for VANTAGE 5H Statistical Design Limit 1.345 Thermal Design Limit 1.50 cu This corresponds to a maximum value of 1.70 when the 4%

measurement uncertainty of Table 7.2-1 is considered.

7 -6 Oi

. - . . . . . - . - - . . . . . - . - . - ~ . . . - . . . - - - . . . - . - ..-.- . . . . . . - . - - . . -- - . - - .

Non Proprietary i

l O Table 7.1-2 Core Thermal-Hydraulic Conditions i

t 100% Power 360,100 gpm 7.0% Bypass Heat Balance Summary i

l Rated Thermal Power, MWt 3411 ,

Ave Heat Flux, BTU /hr-sq-ft 194,100 i Stack Height, in 144 l'

Fuel Rod Outer Diameter, in 0.374 Fuel Pins per Assembly 264.

Assembly Flow Area, sq-in 38.7 Nominal Thermal Design RCS Flow, gpm 360,100 Flow Fraction Effective for Heat Transfer 0.93

[\ Core Inlet Velocity, ft/sec 14.4 Inlet Mass Flux, Mlb/hr-sq-ft 2.44 Vessel Mass Flow Rate, Mlb/hr 136.19 Pressure, psia 2280 Vessel Ave Temp, F 578.2 Inlet Temperature, F 546.2 l Vessel Outlet Temp, F 610.2 I

( Vessel Delta Temp, F 64.1

! Core Outlet Temp, F 614.5 l Core-Ave Temp, F 580.4 i Core T out - Vessel Tout, F 4.3 i -

3 7-7 1

d I

i 3-....__ _ _._, . . _ _ . . _ _ . . . _ . - - . . _ - - _ _ _ _ _

Non Proprietary Figure 7.1-1 l

LYNXT 31-Channel Bundle-By-Bundle Model g 8 9 10 11 12 13 14 15 l

K '1 . -2 -- 3 4- 5- -7 -8

'9 L 10 11 12 13 14 15 M '16 17 18 19 20 21

)

N '22 23 24 25 26 1/8 Core Symmetry l o . 27 28 29 ,

1 P '30 31 1

I i

i 7-8 i

2

Non Proprietary Figure 7.1-2 LYNXT 12-Channel Model Hot Bundle 1/8th Core Pin Number '

g1 _2 .. e q )* o g .. o .. o o I

tt g.ge; 4 s

6 e

7 7

o o'o o' -

'. , ,o OOOO noi suse /haaa* '-"

' .8g b' -

%g .O@OO u x "c "" "2

's 0000

',2 Instrument sheath Guide Thimble

'@000 i i

'O. 0 O

., Pin 13 = Remainder of Hot Bunde Pin 14 = Remainder of1/8th Core

., i Channel 11 =. Remainder of Hot Bunde Channel 12 = Remainder of 1/8th Core Bundle Pitch Line /

7-9 O O O

Non Proprietary 7.2 Application of Statistical Core Design To support the licensing of the Mark-BW fuel assembly for SON,  !

FCF is using its SCD thermal-hydraulic analysis method. The  !

following section provides a brief overview of the SCD approach and analysis method. A more in-depth discussion of the SCD l method can be found in BAW-10170P-A (Reference 7-6).

7.2.1 Statistical Design Limit -- SDL Whether traditional analysis methods that compound analysis uncertainties, or SCD analysis methods that treat analysis uncertainties statistically are being used, the purpose of the core DNB analysis is to insure that a 95 percent probability exists, with a 95 percent confidence level, that the hot pin will not experience a departure from nucleate boiling (DNB) during normal operation or during transients of moderate frequency. The DNB criterion is met when the minimum calculated departure from nucleate boiling ratio (DNBR) is greater than the design limit DNBR that has been established for the critical heat flux (CHF) correlation being used. Traditionally, the value of the design limit DNBR has been based on an assumption that the worst case level of all input uncertainties occurs simultaneously. In the SCD method, described in BAW-10170P-A, the uncertainties on a group of input variables are subjected to a statistical analysis and an overall DNBR uncertainty is established. This uncertainty i is then used to establish a new DNBR design limit known as the Statistical Design Limit (SDL). All variables treated in the l

development of the SDL are then input to the thermal-hydraulic l analysis computer codes at their nominal level. The SCD method applies to any fuel design and any analysis code with any CHF correlation. However, the application documented in BAW-10170P-A was performed for a 17 x 17 fuel design, with uncertainties and core models selected to be generically representative of a 193-fuel assembly core, which is specifically the SON plant and fuel assembly types. The SDL development documented in that reference utilized the LYNXT code and the BWCMV CHF correlation (BAW-l 10159P, Reference 7-7). The input variables, uncertainties, and distributions treated in that development are as follows:

Variable Name Uncertainty Distribution Q Core Power 2% Normal W Core Flow 2.2% Uniform W Core Bypass Flow 1.5% Uniform P Core Pressure 30 psi Uniform T Core Inlet Temperature 4F Uniform 7-10

Non Proprietary f'~N R Measured F5 5% Normal

! )

\s_ / R Hot Channel Factor 3% Normal R Bundle Spacing 1.5% Normal A Axial Peaking Factor 2.5% Normal Z Axial Peak Location 8 inches Normal D BWCMV Uncertainty 16.87% Normal D LYNXT Code Uncertainty 5% Normal D RSM to LYNXT Fit 2.,5% Normal The resulting SDL for these conditions was calculated in BAW- ,

20170P-A to be 1.345 BWCMV. For plant-specific applications the l SDL must either be verified as applicable or recalculated using actual uncertainty values. To verify that the SDL is applicable, the uncertainties and distributions listed above must be verified as being bounding for the real core or the plant-specific uncertainties and distributions must be incorporated. A listing of plant-specific uncertainties and distributions for the SQN core is given on Table 7.2-1. As seen on that table, for the SQN analyses the core flow uncertainty is 1.3 percent higher than the

\

[sx

"')

generic value, while the measured F5 uncertainty is 1 percent lower. All other uncertainties are the same. A recalculation of the BWCMV SDL based on the SON-specific uncertainties showed that the 1.345 value documented in BAW-10170P-A was bounding.

7.2.2 Thermal Design Limit -- TEL For application of the SCD method to SON, margin is added t0 the SDL to define an analysis limit known as the Thermal Design Limit (TDL). The retained thermal margin made available by using the TDL is calculated using the following formula.

Retained Thermal Margin (%) = TDL - SDL X 100 TDL This retained thermal margin is then used to provide flexibility in the fuel cycle design. Examples of offsets that might be i assessed against the retained margin include transition core effects and penalties for input uncertainties greater than those considered in the SDL development. For the SON application, a thermal design limit of 1.50 BWCMV has been selected. This i results in 10% retained thermal margin. A summary of plant specific penalties to be assessed against the retained margin will be included in the cycle-specific Reload Safety Evaluation l Report. I

(

r~'3 a

\

7-11 1

Non Proprietary Table 7.2-1 Statistical Core Design Application Summary Sequoyah Plant Specific Uncertainties Plant Uncertainties Variable Name Uncertainty Distribution Q Core Power [ ] i W Core Flow [ ]

P Core Pressure [ ] '

T Core Inlet Temperature [ ]

R Measured Fjl [ ]

FCF Ana'.ysis Uncertainties Variable Name Uncertainty Distribution W Core Bypass Flow [ ]

R Hot Channel Factor [ ]

R Bundle Spacing [ ]

A Axial Peaking Factor [ ]

Z Axial Peak Location [ ]

D BWCMV Uncertainty [ ]

D LYNXT Code Uncertainty [ ]

D RSM to LYNXT Fit [ ]

• Also applies to the Westinghouse VANTAGE SH 7-12

Non Proprietary l gN 7.3 Calculation of Safety Limits x_- To assure that the DNBR and fuel melt criteria are met, limits are placed on core operating parameters. The overpower AT trip function provides protection against centerline fuel melt, while the overtemperature AT trip function provides protection against DNB. The range of conditions over which the overtemperature AT trip must provide protection is limited to full flow operation within the low pressure and high pressure trips and to transients l that are slow relative to fluid transport time from the core to the temperature sensors. The overtemperature AT trip limit is based on the core safety limits. These limits, which are calculated with the design power distribution discussed in Section 7.1.3, represent the various combinations of pressure, temperature, and power that preclude hot leg boiling and result in a calculated DNBR greater than or equal to the thermal design limit (TDL). Typically, these safety limits are generated at four pressure levels corresponding to the high and low pressure j trip setpoints plus two intermediate pressures. When the safety limits are plotted in AT versus T., coordinates they form a family of lines which provide the basis for setting the overtemperature AT trip limit. A penalty term, f(AI), is applied to the overtemperature AT trip limits to account for skewed axial power shapes. The f(AI) function is based on maximum allowable peaking limits discussed in Section 7.5. More information on the methods used to determine the f(AI) function is included in BAW-10163P-A (Reference 7-8).

- [V,-}

The following section discusses the methods used to determine the core safety limits.

7.3.1 Hot Leg Boiling Limits The first step in establishing reactor core safety limits is to establish the hot leg boiling limit lines. These lines, which are plotted in average temperature versus fraction of rated power l coordinates, define the locus of points at which the vessel outlet temperature is equal to the saturation temperature for a given pressure at various power levels. The hot leg boiling limits are calculated at pressurizer pressure values ranging from the low pressure limit to the high pressure limit, at power levels ranging from zero to 120% of rated power. Through an iterative process using a heat balance across the vessel, the core inlet temperature that results in the saturation temperature in the hot leg is computed. To ensure that the vessel outlet enthalpy is less than the saturated liquid enthalpy, a target outlet temperature, slightly below the actual saturation temperature, is chosen. From the heat balance calculations, the corresponding average temperature and AT are determined. The

segments of the core safety limits that are set by the hot leg boiling limit are indicated on Figure 7.3-1.

/"'N

\~ / 7-13

l Non Proprietary 7.3.2 DNB Limits The DNB limit lines define the locus of pressure, temperature, and power conditions for which the DNBR value calculated with  !

LYNXT is equal to the thermal design limit (TDL). For this l calculation, the LYNXT code with the design peaking distribution discussed in Section 7.1.3 is utilized. The DNB limit lines are determined by calculating the minimum DNBR as a function of l coolant inlet temperature and power at core exit pressures  ;

corresponding to the hot leg boiling limit lines. The reference  ;

radial peak (F5) is 1.64, as discussed in Section 7.1.3. For l the calculation of safety limits at power levels less than 100%

power, the radial peak is increased according to the equation: )

F2n sju t.a = Flnnomin1 [1 + 0.3(1 - P)]

where P = the fraction of rated power A plot of ty'ical reactor core safety limits is provided on l Figure 7.3-1 Those segments of the core safety limits which are j set by the DNB limit have been indicated on the figure.

7.3.3 Conversion to AT Versus T.,, Coordinates The reactor core safety limits establish the bounds of the overtemperature AT trip function. To calculate the i overtemperature AT trip limits, the reactor core safety limits l are first converted into AT versus T.,y coordinates. The heat '

balance data that is generated when the hot leg boiling and DNB limit lines are calculated provides the inlet, outlet and average ter.iperatures at the statepoints which define the limit lines.

From that data the AT and T.,, coordinates for the statepoints are i determined.

A plot of typical overtemperature limit lines in AT versus T.,,

coordinates is provided on Figure 7.3-2.

l 7-14

6 Non Proprietary O Figure 7.3-1 Typical Reactor Core Safety Limits Four Loop Operation 670 =

660 -- Unacceptable Operation l

l 2400 PSM 640 --

2250 PSM 630 --

DNB Umit C Unos 620 -- 20@ PSM 8

a: Hot Leg Boiling Umit Unos 610 --

1775 PSM 800 --

Acceptable Operation 590 --

I 580 --

l 570  :  :  :  :

0 0.2 0.4 0.6 0.8 1 1.2 v) Fraction of Rated Thermal Power 7-15

Non Proprietary l Figure 7.3-2 l Typical Ove temperature Limit Lines Vessel Delta Temperature Versus Vessel Average Temperature 80 Unacceptable Operation 70 -- Unos 60 --

2400 PSM 50 --

u. 2000 PSM 1775 PSM 40 .

2250 PSM g

30 --

Acceptable Operation 20 --

Hot Log Bolling Umit Unes 10 --

0 '

570 580 590 600 610 620 630 640 650 660 670 Vessel Average Temperature, F 7-16

Non Proprietary i\s_]i/

7 7.4 Transient Analysis Methods The thermal-hydraulic performance of the Mark-BW fuel assembly during accident conditions is demonstrated by examining the DNB response of the fuel for the SON design basis transients. The transients that are evaluated are divided into two categories; '

those evaluated using the SCD methods described in Section 7.2, and those evaluated using traditional, non-SCD methods.

Generally, a transient is analyzed using SCD methods if the plant conditions during the transient are bounded by the range of conditions assumed in the statistical design limit development and the CHF correlation used is applicable. If either of these criteria is not met, then non-SCD methods must be used. Results from the transient analyses performed to support the SQN reload safety analysis are contained in Section 6.0 of this report.

7.4.1 SCD Transients Transients that use SCD methods are analyzed with the 12-channel LYNXT model, described in Section 7.1.1, with the addition of a fuel rod model. Several fuel rod modeling options are available in LYNXT including both variable and constant gap conductance models. Generally, the constant gap conductance model is more conservative than the variable gap conductance model. In either case, the fuel rod model is initialized using data from a s bounding TACO 3 fuel thermal analysis.

k'~') In addition to adding the fuel rod model, the LYNXT input deck is also modified for transient applications to include the actual transient input data. The time-dependent variation of pressure, temperature, flow, and power are determined using the RELAPS/ MOD 2 system safety analysis code. The evaluation criterion for DNB-limited transients is that the minimum DNBR must be greater than or equal to the thermal design limit DNBR. SCD methods were used in the thermal-hydraulic evaluation of four of the transients documented in Section 6.0. The four transients were as follows: f the complete loss of forced reactor coolant flow, the loss of external electrical load (turbine trip), the uncontrolled rod cluster control assembly bank withdrawal at power, and the single reactor coolant pump locked rotor.

7.4.2 Non-SCD Transients The analysis of non-SCD transients uses essentially the same LYNXT model as the SCD transients. However, non-SCD methods require that the uncertainties identified in Table 7.2-1 be accounted for deterministically. Accounting for these I uncertainties requires the addition of engineering hot channel ,

factors to the model and requires that uncertainties on radial peak, core power, system flow, bypass flow, core pressure and l 6

inlet temperature be taken into account. In addition, for

/ 1

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N-- 7-17 f

l 1

Non Proprietary transients analyzed with the BWCMV CHF correlation, instead of comparing the calculated DNBR to the SCD thermal design limit (TDL), the calculated DNBR's are compared to the 1.21 BWCMV CHF correlation design limit. When non-SCD methods are used, NRC-approved CHF correlations other than BWCMV may also be used, with design limits appropriate for the conditions being analyzed.

Non-SCD methods were used in the thermal-hydraulic evaluation of the steam line break transient documented in Section 6.

7.5 Maximum Allowable Peaking Limits Both transient analyses and steady-state analyses performed to define core safety limits use a reference design radial and axial power distribution, called the " design peaking", that is assumed to bound, in terms of DNBR performance, real power distributions occurring during plant operation. To provide assurance that this assumption is valid, maximum allowable peaking, or MAP, limits are developed. These limits are a family of curves, typically plotted as maximum allowable total peak versus axial location of peak, with the axial peaking factor as a parameter. The family of curves is the locus of points for which either the minimum DNBR is equal to a target value, typically the thermal design limit (TDL), or the coolant quality at the minimum DNBR location is equal to the CHF correlation quality limit. The MAP limits provide linkage between the DNBR analyses, using a single design peaking distribution, and the core operating or safety power distribution limits. BAW-10163P-A (Reference 7-8) defined four Technical Specification Limiting Conditions for Operation (LCO's) that are related to the DNBR analyses through the MAP limits.

They are as follows:

T.S. 3.2.1 Axial Flux Difference (AFD)

T.S. 3.2.2 Heat Flux Hot Channel Factor - Fo(X,Y, Z)

T.S. 3.2.3 Nuclear Enthalpy Rise Hot Channel Factor - Fa(X,Y)

T.S. 3.2.4 Quadrant Power Tilt Ratio For typical applications, two sets of MAP limits are developed.

The first set, known as safety limit MAPS, is based on various statepoints at the reactor core safety limits. The second set, known as operating limit MAPS, is based on conditions at the point of minimum DNBR from the most limiting loss of coolant flow transient. MAP limits are generated using the steady-state 12-channel LYNXT model discussed in Section 7.1.1. For both sets, allowable peaking is determined for a range of axial peaking values,[ ). This produces a family of lines that define the maximum allowable radial (or total) peak as a function of axial peak location for different axial peak magnitudes.

MAP .'.imits are used in the plant maneuvering analysis as documented in BAW-10163P-A (Reference 7-8) to determine the 1

1 7-18 0

i Non Proprietary i 1

7

'g margin that exists for each predicted power distribution relative :

j to the DNBR limit. Core safety and operating limits are then r( /

,_ defined such that only positive margin cases are permitted.

Safety limit MAPS are used in defining the f(AI) modifier to the overtemperature AT trip function. The operating limit MAPS are used in setting initial condition peaking limits that preserve the minimum DNB ratio during the most limiting loss-of-flow accident. Additional discussion on setting core safety and operating limits is included in Section 7.5.3.

7.5.1 Safety Limit Maximum Allowable Peaking Limits

[ ]

A plot of typical safety limit maximum allowable peaking limits is included on Figure 7.5-1.

7.5.1.1 Safety Limit MAPS Variation With Power MAP limits are used in the plant maneuvering analysis to determine allowable core power as a function of axial flux difference (AFD). The power versus AFD limit is then used to define the f(AI) portion of the overtemperature AT trip function.

To define the allowable variation of MAP limits with power, [

]

o u

\s_- 7-19

than Proprietary

[ ]

7.5.2 Operating Limit Maximum Allowable Peaking Limits Transient DNBR calculations are performed with a reference design power distribution, typically the same radial distribution as used for safety limit analyses. For non-overtemperature AT transients, a second set of MAPS, the operating limit maximum allowable peaking limits, are developed. These MAP's represent a minimum DNBR equivalent to that calculated for the limiting non-overtemperature AT transient. These operating limit MAPS are developed in a manner that is similar to that used for the safety limit MAPS [ ).

A plot of typical operating limit maximum allowable peaking limits is included on Figure 7.5-2.

7.5.2.1 Operating Limit MAPS Variation with Power The power peaking adjustment factor for reduced power for the l operating limit MAPS represents the additional DNB margin that is I available for transients starting from a reduced power condition. i The factor is calculated in a manner similar to that used to i define the reduced power safety limit MAPS. [ ]

7.5.3 Core Safety and Operating Limits The purpose of the safety limit and operating limit maximum  !

allowable peaking limits is to define combinations of radial and I axial peaking that provide DNBR performance equivalent to the i design power distribution. These allowable peaking limits are then incorporated into the core maneuvering analysis that establishes core safety and operating limits as described in the 7-20

Non Proprietary i

f- Core Operating Limit Methodology topical report, BAW-10163P-A l

( (Reference 7-8). As outlined in that document, the safety limit l MAPS provide the basis for the f(AI) term in the Limiting Safety System Settings (LSSS) overtemperature AT trip function. The overtemperature AT function provides DNB protection for steady-state operation and slow transients. The f(AI) term is a l function of axial flux difference ( ?.?D) and protects the core j from highly skewed power shapes that may approach the fuel  !

thermal limits. The operating limit MAPS provide the bounds that set initial Condition AFD limits to preserve the DNBR thermal design limit during the most limiting loss-of-flow accident. To determine the core safety and operating limits, [ ] ,

7.5.3.1 Physics check Cases When generating the safety and operating limit MAP curves, the peaking limits are based on a series of smooth, mathematically-derived axial power shapes. [ ]

I i

i

..s s

i

)

I l

l l

l l

ks 7-21

Non Proprietary Figure 7.5-1 Typical 118% Power Safety Limit MAP Curves Allowable Total Peak Versus Axial Peak Location M

l 1

l l

O 7-22

1

Non Proprietary a'

i f

j Figure 7.5-2 i Typical 100% Power Operating Limit MAP Curves i

! Allowable Total Peak Versus Axial Peak Location i

I l ,

4 i

I-J l

i b

! i i

t l

l l

)

l 1

l 1

i l 7-23 I

}

i

Non Proprietary 7.6 Mixed Core Analysis For any new fuel design that is being introduced on a reload basis, hydraulic compatibility must be established with the existing, or resident, fuel in the core. Therefore, when a new or modified fuel design, having different hydraulic characteristics from the resident fuel is introduced, a transition core analysis is performed. For each mixed core configuration during the transition, performance of each fuel type is evaluated relative to a reference analysis. This reference analysis is typically based on a full core of the new fuel. To determine the performance of each fuel type relative to the reference analysis, each fuel type is modeled in a conservatively bounding mixed core configuration. These mixed core analyses are used to determine the transition penalty that is applicable to each fuel type. A typical transition core analysis determines the effect of the mixed core on parameters such as minimum DNBR, core pressure drop, fuel assembly lift, and lateral flow velocities.

The Mark-BW fuel assembly, described in BAW-10172P (Reference 7-4), was designed to be hydraulically compatible with the full range of 12-foot Westinghouse 17x17 fuel designs, including the Westinghouse Standard and the Westinghouse VANTAGE SH. During the transition to Mark-BW fuel at the SON, the resident fuel (i.e. the fuel being displaced by Mark-BW fuel assemblies) will primarily be the Westinghouse VANTAGE SH fuel design with some Westinghouse Standard fuel reinserts. Table 7.6-1 provides a comparison of the key design differences between the three assembly types. Compatibility of the Mark-BW fuel design with the Westinghouse Standard design was demonstrated as part of the effort to license the use of the Mark-BW at Portland General Electric's Trojan Plant (References 7-9, 7-12, and 7-14). The results from those studies indicated that a [ ]. However, for the SON application, the limited use of Westinghouse Standard reinserts will result in a transition core where the influence of the VANTAGE SH predominates. Therefore, no transition penalty will be taken because of the presence of the Westinghouse Standard assemblies, and, as a result, the following discussion will focus solely on the transition from the VANTAGE SH fuel design to the Mark-BW.

7.6.1 Methods and Models for Resident Fuel The Mark-BW is similar in design to the Westinghouse VANTAGE 5H 17x17 fuel design, with the only significant differences being that the first Zircaloy grid on the Mark-BW is a non-mixing grid and the VANTAGE SH upper guide thimble OD is 0.474 inch as opposed to 0.482 inch on the Mark-BW. From the thermal-7-24

Non Proprietary g 's hydraulics perspective the two designs are identical with the exception of the hydraulic form loss coefficients for the spacer

\s - grids and nozzles. Therefore the thermal-hydraulic core models described in Section 7.1.1 apply also to the VANTAGE SH fuel, when the appropriate form loss coefficients are used.

For the evaluation of DNB effects in the transition core, the

, BWCMV CHF correlation (BAW-10159P, Reference 7-7) is used for evaluation of the VANTAGE SH fuel. FCF bases the extension of BWCMV to the VANTAGE SH on the fact that the BWCMV data base includes CHF data representative of the Westinghouse Standard fuel design, and the fact that BWCMV has been approved for application to that fuel and to the Westinghouse OFA design (Reference 7-15). In addition, since the VANTAGE SH grids are similar to, and bracketed by, the geometries already contained in the BWCMV data base, the correlation applies without modification.

The statistical core design thermal-hydraulic analysis method, documented in BAW-10170P-A and summarized in Section 7.2, was developed independently of fuel type, therefore the SCD method applies to the VANTAGE SH fuel when the appropriate uncertainties e

are incorporated. The treatment of uncertainties in the development of a statistical design limit is discussed in Section

[ 7.2 of this report. From that discussion it can be seen that the J

,- s only uncertainty parameter dependent on the fuel design is the

[ engineering hot channel factor. A suitably conservative value i (1.03) has been selected for this parameter, such that the SDL applies to the VANTAGE SH fuel.

7.6.2 Mixed Core DNBR Analysis For transition cycles in which the resident VANTAGE SH fuel is being displaced by Mark-BW fuel, core DNBR safety and operating limits and DNBR margin during transients are based on analysis of the final full-core Mark-BW configuration. Similarly, the reference reload safety analysis as documented in Section 6.0, is performed for the Mark-BW fuel design. To ensure that these full-core analyses remain applicable to the transition cycles a mixed core analysis is performed. This mixed core analysis quantifies the transition cycle penalty that must be applied to either the resident or the new fuel designs. To maintain the applicability of the full-core analyses, the transition core penalty may be applied either as an assessment against retained thermal margin that is incorporated in the DNB analysis through the use of the thermal design limit (TDL, per BAW-10170P-A, Reference 7-6), or by the identification of an offsetting conservatism.

The transition core penalty is determined generically by modeling a bounding core configuration. For applications where it is

\

/

-\s / 7-25

Non Proprietary desirable to reduce the penalty to a value less than the generic value, a core-specific calculation is performed, using a model that represents the actual transition cycle core geometry.

Following is the process used for the evaluation of the transition from VANTAGE SH to Mark-BN:

[ ]

l O

1 l

l 7-26

Non Proprietary i

[ ]

7.6.3 Mixed Core Pressure Drop. Lift and Crossflow With the transition from one. fuel design to another, the mixed core pressure drop is bracketed by the homogeneous core pressure drop values for each of the fuel designs. Therefore, mixed-core pressure-drop calculations are not required.

For fuel assembly hydraulic lift force determination, the limiting core configuration is the homogeneous core with the highest pressure drop. For the Westinghouse VANTAGE SH to Mark-BW transition, this would be the full VANTAGE 5H core. As part of the mixed core methodology, an evaluation is performed to determine the limiting configuration for lift forces on each assembly type. This evaluation is based on the relative pressure drops of the two fuel designs. Knowing the relative pressure drop of the two assembly types, the flow diversion characteristics of the core can be determined, and in turn, the core configurations that produce the highest lift forces in the two fuel types can be determined. Mixed core pressure drop and hydraulic lift force calculations are determined by modeling the mixed core with the LYNXT bundle-by-bundle model. LYNXT is also used to calculate inter-assembly crossflow velocities, which are used to evaluate the potential for flow-induced vibration.

Typically the limiting configuration for lift in a transition core is a single lower pressure drop assembly in a full core of the higher pressure drop assembly. [ _]

[ 7-27

Non Proprietary Although the total unrecoverable pressure drop of the Mark-BW is very similar to that of the VANTAGE SH, there are distributional differences that cause some flow redistribution within the core.

However, lateral crossflow velocities for the mixed core configuration have been calculated and found to be acceptable.

O O

7-28

. . - - . - . - _ - . - - . - . - . - . - - . . -. - ..~ - .- - -. - - - . . - - . -- - - -- - ..

Non Proprietary Table 7.6-1 Comparison of Mark-BW, W VANTAGE 5H and W Standard Fuel Nominal Designs

-i 17x17 FCF 17x17 W 17x17 W, Mark-BW VANTAGE 5H Standard Fuel Assembly Fuel Assembly Fuel Assembly Parameter Design Design Design I ]

7-29

Non Proprietary l

7.7 Fuel Thermal Performance Analysis Fuel thermal performance analyses that predict fuel rod temperature and internal pressure conditions during core operation are performed for the UO 2fuel rods with TACO 3 (BAW-10162P-A, Reference 7-10) and for the Gadolinia fuel rods with GDTACO (BAW-10184P-A, Reference 7-11). These analyses are used to determine the centerline melt limit and the maximum fuel rod burnup limit, which is based on the fuel rod internal pressure.

l In addition, these analyses provide initial fuel temperature and I

pressure conditions for LOCA and non-LOCA safety analyses.

The TACO 3 code and internal gas pressure analysis methodology have been extended to address fuel rod operation with pressure greater than reactor coolant system pressure in the NRC-approved topical report BAW-10183P-A (Reference 7-16), which addresses five primary issues:

• Determining the rod internal gas pressure where the clad creep out rate exceeds the fuel diametral swelling rate.

• Verifying that this pressure is an acceptable upper limit criterion for fuel rod internal gas pressure.

• Verifying that hydride reorientation effects are less limiting than fuel rod internal gas pressure.

• Verifying that DNB will not propagate with rods operating above system pressure.

• Verifying that LOCA limits will not change with rods operating above system pressure.

l l Together these codes and the fuel rod gas pressure criteria for TACO 3 form the basis for FCF's fuel performance licensing methods.

7.8 Thermal-hydraulic Conclusions The purpose of this section was to define the thermal-hydraulic licensing methods and models that are used to justify the transition from the Westinghouse VANTAGE SH fuel design to FCF's Mark-BW fuel design at the SQN. As this section has shown, during the transition to Mark-BW fuel, thermal-hydraulic safety and operating limits are typically defined using a full core Mark-BW analysis, with transition core effects evaluated with appropriately bounding mixed core models. This section has also shown that, based on the geometric and hydraulic similarities of the Mark-BW, Westinghouse VANTAGE SH, and the Westinghouse Standard, the transition effects will be small. Therefore based on the thermal-hydraulic methods, models, and assessment 7-30

a Non Proprietary 7sg discussed in this section, it is shown that the Mark-BW is compatible with the SON core and able to meet all operational ks _j*l design requirements during transition cycles and during full core ,

operation.

7.9 Thermal-hydraulic References 7-1 BAW-10156-A Revision 1, LYNXT: Core Transient Thermal- 1 Hydraulic Program, August 1993. l l

7-2 Letter, K.E. Suhrke (B&W) to Mr. S.A. Varga (NRC),

December 6, 1976.

7-3 Letter, S.A. Varga to J.H. Taylor, Update of BAW-10055,

" Fuel Densification Report," December 5, 1977.

7-4 BAW-10172P, Mark-BW Mechanical Design Report, July 1988.

7-5 BAW-10147PA-R1, Fuel Rod Bowing in Babcock & Wilcox Fuel Designs - Revision 1, May 1983.

l 7-6 BAW-10170P-A, Statistical Core Design For Mixing Vane Cores, December'1988.

7-7 BAW-10159P, BWCMV Correlation of Critical Heat Flux in

['~

Mixing Vane Grid. Fuel Assemblies, May 1986.

7-8 BAW-10163P-A, Core Operating Limit Methodology for ,

Westinghouse-Designed PWR's, June 1989. i 7-9 BAW-10176P, Mark-BW Reload Safety Analysis for Trojan, January 1990.

7-10 BAW-10162P-A, TACO 3 -' Fuel Pin Thermal Analysis Computer Code, October 1989.

7-11 BAW-10184P-A, GDTACO - Urania-Gadolinia Thermal Analysis Code, February 1995.

7-12 BAW-10178P, Mark-BW Thermal-Hydraulic Applications for the Trojan Nuclear Plant, March 1990.

4 7-13 BAW-10189-P, CHF Testing and Analysis of the Mark-BW Fuel Assembly Design, August 1993.

7-14 BAW-10177, Mark-BW Reload LOCA Analysis for the Trojan Nuclear Plant, October 1990.

\s_s 7-31

Non Proprietary 7-15 Letter dated May 22, 1989, Ashok C. Thadani (NRC) to Mr.

J. H. Taylor (B&W), " Acceptance for Referencing of augmented Topical Report BAW-10159P, BWCMV Correlation of Critical Heat Flux in Mixing Vane Grid Fuel assemblies, May 1986.

1 7-16 BAW-10183P-A, Fuel Rod Gas Pressure Criterion (FRGPC),

July 1995.

1 7-17 BAW-10186P, Extended Burnup Evaluation, November 1992. l l

7-18 Sequoyah Nuclear Plant - Final Safety Analysis Report, i Amendment 10, April 14, 1994. l O

l 7-32 1

1 1 1 I

Non Proprietary 8.0 Mechanical Analysis and Evaluations

/'"s (s i To ensure safe and reliable operation, the structural integrity of the Mark-BW fuel assembly was evaluated for the loadings I associated with the normal operation, seismic and loss-of-coolant-accident (LOCA) events, and shipping and handling.

Section 8.1 discusses the normal operation and evaluation of mechanical compatibility of the Mark-BW fuel assembly  ;

configuration with the Westinghouse resident fuel and  ;

Westinghouse-designed reactor internals. Section 8.2 presents )

the results from the seismic and LOCA evaluations. Section 8.3 1 discusses the shipping and handling analysis.

Margins reported are calculated by the following:  ;

1 Margin % = ((Allowable-Predicted)/ Predicted]

• 100% l l

The methods of analysis for each criteria and the inputs used are conservative. A margin of greater than () is sufficient. l References for the analysis methods are provided where l appropriate in the text of each section.  !

Design bases and the methodology for the fuel assembly structural evaluations are essentially the same as described in topical s report BAW-10172P, Reference 8-1. This topical report has received the NRC approval for referencing in licensing

(. applications. References are made to this topical report, where the design bases and evaluations are directly applicable.

8.1 Normal Operations )

8.1.1 Growth Allowance Evaluations The axial gaps between the top nozzle and reactor upper core plate, and between the top nozzle grillage and fuel rods were conservatively analyzed to show that these gaps allow sufficient margin to accommodate the fuel assembly and fuel rod growth to maximize design burnups. These target burnups for the Sequoyah Nuclear Plant (SON) Mark-BW design are [ ] mwd /mtU for the peak assembly and [ ] mwd /mtU for the peak rod. The analysis was conducted using the latest irradiation growth data for Zircaloy guide thimbles and fuel rods collected during several post-irradiation examinations (PIE) for each of the Framatome Cogema Fuels (FCF) designs.

The minimum fuel rod shoulder gap at end of life (EOL) for a target burnup of ( ) mwd /mtU is predicted to be [ ] inch at j worst-case hot conditions. For the fuel rod growth evaluations, i worst case is considered to be maximum fuel rod growth and j minimum guide thimble growth. Differential thermal expansion is  !

l O 8-1 1

1

i Non Proprietary also considered at operating temperatures. The FCF growth data are used to predict fuel assembly growth due to irradiation and

thermal expansion. These data demonstrate that the Mark-BW l assembly is expected to reach burnups up to [ ] mwd /mtU at l nominal conditions, and [ ] mwd /mtU at worst-case conditions.

l

[ ]

8.1.2 Fuel Assembly Comoonent Stress Analysis This section discusses the stress analysis of assembly components under the normal operating loads.

The structural design requirements for the Mark-BW fuel assembly are derived from past experience with the McGuire and Catawba Plants, as well as experience with other FCF designs. The design bases and design limits are set based on this experience. These values are then verified to ensure conformity with SQN fuel plant-specific design requirements.

The evaluations performed to verify the structural integrity of the Mark-BW fuel assembly components are presented in the following sections. Table 8-1 shows the design limits and margins for each of the major Mark-BW structural components under normal operating loads.

8.1.2.1 Fuel Assembly Holddown Sprinos The design bases for the Mark-BW fuel holddown springs require that the springs be capable of maintaining fuel assembly contact with the lower support plate during normal operating conditions.

During pump overspeed condition, the fuel assembly should not cause the springs to deflect to the solid state nor produce any permanent set.

The Mark-BW holddown springs were analyzed to show that the holddown springs can accommodate irradiation growth of the fuel assembly, and the differential thermal expansion between the fuel assembly and the core internals. The fuel assembly lift evaluation was performed by comparing the holddown force provided by the leaf springs with the SON hydraulic forces at both normal operating conditions and at the pump overspeed condition.

The analysis indicated that the Mark-BW fuel assembly will not liftoff under any normal operating condition. The minimum margin to fuel assembly liftoff occurs at beginning of life (BOL),

[ ]. The margin at this condition is []. At 120% pump 8-2

Non Proprietary l

f'~'}f

(

overspeed condition, the fuel assembly will experience some liftoff. The liftoff will be minimal and the holddown spring

\__/ deflection will be less than the worst-case normal operating cold

, shutdown condition. The margins for the conditions above can be seen in Table 8-1. [ ]  ;

1 The Mark-BW fuel assembly holddown spring stress calculations show that the springs are structurally adequate under all static and fatigue loading conditions. Furthermore, in the unlikely ,

l event of spring failure, the top nozzle provides positive l retention of the holddown springs. The integrity of the holddown l springs has been proven in the McGuire and Catawba power plants.

l The clamp screws which mount the holddown springs on the top plate of the top nozzle were also analyzed for normal operation I and fatigue loading to determine their structural adequacy. The 4

reactor coolant system design transients used for analyzing 2 fatigue failure were taken from the SQN Final Safety Analysis Report (FSAR - Reference 8-2). A listing of these transients may be seen in Table 8-2. The analysis indicated that the clamp screws are structurally sound for all loading conditions. The margins for the clamp screw at operating temperature for thread shear stress, head shear stress, and bearing stress are [ ],

respectively.

f- 8.1.2.2 Guide Thimble ks _,) The design bases for the guide thimbles state that no buckling of the thimbles shall occur during normal operation or any transient condition under which control rod insertion is required.

As can be seen in Table 8-1, guide thimble buckling was analyzed for all loading conditions. The guide thimbles were analyzed for both normal operating loads and pump overspeed loading. Margins were calculated for the following load cases:

1) 100% full power rechanical design flow rate

2) 120% full power mechanical design flow rate (Pump overspeed condition)

3) 100% full power mechanical design flow rate with an upper bound scram load of [ ]

For case three, the rod cluster control assembly spring in the hub is assumed to be fully compressed during a scram. The margins for this case were determined assuming the load needed to produce the maximum allowable compressive yield stress. For the O\

\N- j 8-3

Non Proprietary first two cases, the margins were based on the load required to produce a midspan deflection of [ Jinch using the secant formula. The midspan lateral deflection of [ J inch is based on not affecting control rod insertion or trip performance.

In calculating the maximum compressive load in the guide thimbles, the worst-case ferrule-to-grid gap assumptions were used. The stiffness of the dimple interface is assumed to be linear. The linear representation is conservative, resulting in higher loads. In the actual assembly, a yielding ferrule will simply slide, distributing its load to other ferrules, resulting in well distributed loads.

Allowable loads for the two forms of the secant formulas in the Reference 8-1 report were calculated based on using one conservative design eccentrity value for all the guide thimble spans. To obtain more realistic limits, these loads were revised based on a span-based value for guide thimble eccentricity.

At 100% full power plus scram load, the design margin is []. For 120% pump overspeed, the design margin is []. These margins were limiting and all other margins were found to be acceptable.

8.1.2.3 Spacer Grids The design bases of the Zircaloy intermediate and Inconel end grids require that no crushing deformations occur due to normal operation. The grids must also provide adequate support to maintain the fuel rods in a coolable configuration under all conditions.

The structural integrity of the spacer grid assemblies was confirmed by testing for Duke Power as discussed in Reference 8-

1. Testing of the grids included spacer grid impact tests of the Zircaloy intermediate grids to determine the dynamic characteristics, static crush tests of the Zircaloy grids to determine the static stiffness and elastic load limit, and slip tests of both the Inconel and Zircaloy spacer grids to determine the required soft stop forces needed to prevent slipping of the grids relative to the fuel rods. All testing indicates that the grids provide adequate design margins.

8.1.2.4 Too and Bottom Nozzles The top and bottom nozzle design bases follow those outlined in Section III of the ASME Boiler and Pressure Vessel Code (Reference 8-3). These values, along with the margins, are listed in Table 8-1.

8-4

J Non Proprietary )

, l 4

fN g Finite-element analyses of the top nozzle grillage and the bottom nozzle grillage using ANSYS (Reference 8-4) have been performed

)

(d N to show that the designs are more than adequate to withstand the normal operating loads. The loads used for these analyses were from EOL shutdown condition at [], since this is the condition j at which the holddown force is maximum. At the operating <

i condition temperature of { ], a conservative scram load was 4 applied to the grillage in addition to the holddown force. The

' calculated stresses are given in Table 8-1.

For the bottom nozzle, in addition to the scram load applied to

. the top nozzle, the weight of the fuel assembly was considered l

when analyzing the structural integrity of the grillage. Finite- ,

element analysis of the grillage using these loads shows the l

adequacy of the debris-resistant bottom nozzle design for all j normal loads. The calculated stresses using ANSYS are given in Table 8-1.

4 8.1.3 Mechanical Comoatibility l Mechanical compatibility of the Mark-BW fuel assembly with 4

reactor internals, handling and storage equipment,~and resident ,

' fuel assemblies has been confirmed by direct comparisons with i measurements taken from Westinghouse VANTAGE SH fuel and standard

! fuel assemblies. Additional interface information was obtained 3

from drawings. Since the insertion of a Westinghouse Standard

I fuel assembly adjacent to a Mark-BW is a possibility, 4 ( '

compatibility between these two assemblies was also verified.

Interface information for the Westinghouse Standard was obtained

previously from measurements taken from four such assemblies (Reference 8-5).

i A dimensional comparison to show compatibility between the Mark-

) BW, Westinghouse VANTAGE SH and Standard fuel assemblies is shown

! in Table 8-3. The comparison in Table 8-3 shows that all i l

! critical interface dimensions on all three fuel assemblies are

. similar. l l

l 8.2 Faulted Condition Loads l These analyses address the adequacy of the Mark-BW fuel assembly

design under accident conditions. The accident conditions

analyzed were (1) Safe Shutdown Earthquake (SSE), (2) LOCA and (3) the combination of an SSE plus a LOCA. The loads for the j worst-case LOCA break were combined with those of the SSE to j determine maximum fuel assembly loads. Fuel assembly responses

, resulting from seismic excitation and LOCA were analyzed using  :

j the general procedure outlined in topical report BAW-10172P (Reference 8-1).

4

)I 8-5 i

4

Non Proprietary In the accident analyses, the lateral effect ('OCA L and seismic) ,

and the vertical effect (LOCA) are analyzed separately. All the I methodology and techniques for the analysis were identical to I those used in previous Mark-BW fuel assembly licensing analysis (References 8-1 and 8-5).

This section discusses the results of the fuel assembly horizontal seismic and LOCA and the vertical LOCA analyses.

l 8.2.1 Horizontal Seismic and LOCA l The SSE and LOCA time history motions of the upper and lower core plates and the upper elevation of the core baffle plate were l applied to the reactor core model as shown in Figure 8-1. The model was analyzed using the computer program STARS Version 16.6RS described in Reference 8-6. The maximum grid impact forces obtained from the SSE analysis were within the allowable elastic load limits as determined by the spacer grid impact tests described in the Reference 8-1 report.

The design basis LOCA time histories uses " leak-before-break" methodology. The displacements provided are those associated '

with a worst-case attached pipe break. For the cold legs, data for an accumulator line break are provided. For the hot legs, data for a pressurizer surge line break are provided. The fuel assembly response resulting from the design pipe breaks was analyzed using the core model shown in Figure 8-1 to determine the grid impact forces and deflections. The loads for LOCA plus SSE were combined by the square-root-of-sum-of-squares (SRSS) method as discussed and accepted by the NRC in Reference 8-7, "NUREG-0800, Standard Review Plan 4.2, Appendix A." The spacer grid impact loads were found to be well within the allowable limits as determined by the spacer grid impact tests. No permanent deformation of the grid is expected in any of those cases; the loads are acceptable. l A mixed core bounding analysis of seismic and LOCA events was performed for Mark-BW and Westinghouse VANTAGE SH fuel assemblies representative of transition cores. Properties of the Westinghouse assemblies were based on the test results as described in Section 3.2 and taking into consideration the differences in assembly masses, stiffness, and spacer grid construction. The results of the analysis were compared with the faulted conditions analysis results of the Mark-BW full core configuration. The resulting changes in spacer grid impact loads are minor ([ ]). The spacer grid impact loads are well within the elastic load limit, and hence the requirement of a core coolable geometry is met.

A mixed core bounding analysis of seismic and LOCA events was performed for Mark-BW and Westinghouse Standard fuel assemblies.

8-6

1 Non Proprietary

The resulting changes in spacer grid impact loads are minor ([])

Hence, the

('~'N)

V and well within the spacer grid elastic load limit.

requirement of a core coolable geometry is met for all combinations of Westinghouse (VANTAGE 5H and Standard) and Mark-BW fuel assembl.ies in mixed cores.

8.2.2 Vertica. LOCA Analysis The SQN reactor coolant piping analysis uses " leak-before-break" i methodology. The forces provided are those associated with a worst-case attached piping break. The vertical force time i histories for the cold legs are for an accumulator line break

! (cold leg break - CLB) and those for the hot legs are for a pressurizer line break (hot leg break - HLB).

A schematic of the fuel assembly vertical model is given in Figure 8-2. The fuel assembly vertical model was analyzed using the general purpose finite-element code ANSYS (Reference 8-4) because of its capability to distribute a given hydraulic force time history over a large number of mass nodes. The methods outlined in the Reference 8-1 report were applied to the model.

Only the LOCA cases were evaluated in the vertical direction, as  ;

the upper and lower grid plates move in phase for the seismic 1 case and cause no fuel assembly loading, l

The maximum component forces obtained from the analysis are l

[g) summarized in Table 8-4. The results of the analysis show that V because of the holddown spring applied preload and stiffness, the fuel assembly does not impact the upper core plate during the LOCA. These forces are well below conservatively calculated allowable loads for the guide thimbles and fuel rods.

8.2.3 Fuel Assembly ComDonent Stresses Under Faulted Condition Loads The Mark-BW fuel assembly. component stress analysis under faulted conditions was performed using loads generated by a separate seismic and LOCA analysis. For the structural analysis, two load cases were addressed. These were SSE and combined SSE plus LOCA loadings. The loads for the worst-case LOCA break are combined with those of the SSE to determine maximum fuel assembly loads.

The design criteria per NUREG-0800 (Reference 8-7) for each of these loadings are as follows:

1) For the SSE, the fuel assembly is designed to allow control rod insertion and to maintain a coolable geometry.

2) For LOCA or combined LOCA plus SSE, the fuel assembly is designed for the safe shutdown of the reactor system.

8-7

1 I

l 1 l

Non Proprietary ,

1 l t l For conservatism, the design criteria used for all of the above l loadings are that the fuel assembly must maintain structural l integrity, a coolable core geometry, and a path for control rod I insertion. All components in the path of the control rods must not yield or buckle. The stress intensity limits for the components in the analysis are taken from Section III of the ASME Code. Because of the need to maintain a path for control rods, the guide thimble assembly stresses are compared to a more conservative elastic criteria.

The axial and lateral loads used for the faulted analysis were  :

obtained from seismic and LOCA loading analysis. The values given in section III of the ASME code were used for the fuel I assembly general stress criteria. The analysis of the components  !

and determination of failure loads were conducted utilizing both i standard engineering techniques and testing. j The guide thimble buckling analysis was performed using the l column secant formula, and the limit used in the analysis is 1 considered conservative. When guide tube buckling is analyzed in  !

this manner, the question is not how long the column can remain straight or stable under an increasing load, but rather how long the column can be permitted to bend under the increasing load, if the allowable deflection or stress is not exceeded.

In addition to the component stress analysic, the fuel assembly maximum possible deflection, and the resulting guide thimble and fuel rod stresses were calculated. The maximum deflection would be possible in a fuel assembly on the perimeter. Deflection of these assemblies ie limited only by the reactor core baffle plates. The ,maxinwr. fuel assembly probable deflection was determined using the accumulated fuel assembly gaps across one row of the core. The total stress intensity in the fuel rods was calculated by considering: dynamic bending loads, dynamic axial

' loads, and steady-state hoop stress caused by the pressure differential between the reactor system pressure and the fuel rod internal pressure. The fuel rod internal pressure considered was I the BOL minimum fill gas pressure at room temperature, which is a l conservative approach.

The maximum stresses for various fuel assembly components during faulted conditions are summarized in Table 8-5. The design l margins indicate that all major components of the Mark-BW fuel assembly meet the design requirements for the SSE and SSE plus LOCA loading events. The SSE requirement of control rod insertion is also met with an adequate margin.

8.3 Fuel Assembly Shippino and Handlina The Mark-BW fuel assembly is designed to maintain its dimensional and structural integrity when subjected to normal shipping and 1

8-8

Non Proprietary handling loads. To preclude spacer grid crushing failures O resulting from excessive clamping loads imposed by the shipping container, FCF has developed a clamping bracket that mechanically limits the imposed load. Crush tests as described in the Reference 8-1 report have been performed to ensure that the spacer grid dimensional stability is maintained during normal shipping and handling conditions. This configuration has been used to successfully ship over [] Mark-BW fuel assemblies.

The design bases and evaluation for shipping and handling loads in detail are provided in Section 4.2.1 of Reference 8-1.

In addition, handling loads which could be applied to the bottom nozzle are the result of vertical impact of the fuel assembly l' during [] f t/ min setdown. A finite-element model of the lower nozzle was used to calculate the stresses for this loading condition. The results of the finite-element analysis are then compared to the ASME code allowable stress. The minimum margin of safety for handling is [ ] for membrane plus bending stress during assembly setdown.

8.4 Mechanical conclusions The mechanical design of the Mark-BW fuel assembly has been shown to meet the design goals provided. The basic design parameters of the Mark-BW are compared to those of the VANTAGE SH and

[ \ Standard Westinghouse 17x17 designs. This comparison shows that

( )

all critical interface dimensions on all three assemblies are similar and assures compatibility between the Mark-BW and the Westinghouse VANTAGE SH and Standard designs.

The mechanical design analyses have been performed using NRC approved FCF standard design methodologies and all acceptance criteria were met.

Based on the results of these evaluacions and the performance of the Mark-BW fuel assemblies in the'AcGuire and Catawba plants, the Mark-BW fuel assembly design can safely operate in the SQN Units 1 and 2 core up to the lice.ised fuel assembly burnup of

[ ] mwd /mtU.

8.5 References 8-1 BAW-10172P, Mark-BW Mechanical Design Report,, July, 1988.

8-2 Final Safety Analysis Report, TVA Sequoyah Nuclear Units 1 and 2, Section 5.2.

8-3 ASME Code,Section III, Nuclear Power Plant Components, 1992 i Edition.

l l 8-9 l

Non Proprietary 8-4 BWNT-TM-83, Rev. Orig, June 1992, ANSYS 4.4a Engineering Analysis System User's Manual, Volumes 1 and 2.

8-5 BAW-10176, Mark-BW Reload Safety Analysis for the Trojan Nuclear Unit, January, 1990.

B-6 2A-4, STARS 16.6RS, Certification File, STARS-Structural Analysis of Reactor System, May, 1995.

I 8-7 Standard Review Plan, Section 4.2, NUREG-0800, Rev 2, U.S. l Nuclear Regulatory Commission, July, 1981.

I 1

L O .

8-10 O!

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V n Proprietary (,,/

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Table 8-1 Limiting Load Conditions for Fuel Assembly Components Design Design Actual Percent Component Load Condition Limit Basis Limit Load Marcin Holddown [ ]

Spring Margin to Liftoff at Normal Operation and Mechanical Design Flow Rate

1) BOL, 100*F E s lbs

2) BOL, 140*F E s lbs

3) BOL, 500*F I gg

4) BOL, 100% Power I g

5) BOL, 118% Power E g

Margin to Set at 120% Pump Overspeed And Mechanical Design Flow Rate [ ]

at BOL, 100% Power and 118% Power in in Margin to Solid Deflection at 120% Pump [ }

Overspeed and Mechanical Design Flow Rate in in 8-11

Non Proprietary Table 8-1 Limiting Load Conditions for Fuel Assembly Components, (Cont'd)

Design Design Actual Percent Component Load Condition Limit Basis Limit Load Marcin Guide Thimble Holddown Spring and Differential Thermal Expansion Loads at

1) 100% Full Power Mechanical Control Rod [ ]

Design Flow Rate (MDFR) Insertion lbs. lbs.

Drag Load

2) 120% Full Power MFDR Control Rod [ ]

Insertion lbs. lbs.

Drag Load

3) 100% Full Power plus Yield [ ]

SCRAM load Strength lbs. Lbs.

Upper Nozzle Holddown Spring and Differential Thermal Expansion Loads at

1) EOL Cold Shutdown, 100*F ASME Code [ ]

Section III ksi ksi

2) EOL Operating, 650*F ASME Code [ ]

plus SCRAM Load Section III ksi ksi Bottom Nozzle Holddown Spring and Differential Thermal Expansion Loads at

1) EOL Cold Shutdown, 100*F ASME Code [ ]

Section III ksi ksi

2) EOL Operating, 650*F ASME Code [ ]

plus SCRAM Load Section III ksi kai 8-12 O O O

, _-. . ~ _ . . - - . . - . ...-.-----..-. --- . . - - - . - - . -.

4 i

Non Proprietary J

4 l Table 8-2 i

Summary of Reactor Coolant System Design Transients 8

t i Event Description Anticipated I Life-time I Normal Conditions Occurrences

1. Heatup and cooldown.at 100oF/hr 200 (each) l 2. Unit loading and unloading at 18,300 (each) 5% of full power / min i 3. Step load increase and decrease 2,000 (each) j of 10% of full power i 4. Large step load decrease 200 l 5. Steady state fluctuations infinite i

f Moset Conditions

1. Loss of load, without immediate 80 turbine or reactor trip A

/

^ '

2. Loss of power (blackout with natural 40 circulation in the reactor coolant system)

3. Loss of flow (partial loss of flow, 80 one pump only)

4. Reactor trip from full power 400

5. Spray actuation with a differential 12 temperature > 320*F s 560*F

6. 1/2 safe shutdown earthquake Reactor Vessel 200 cycles Steam generator and pressurizer 50 cycles

\

8-13

I l Non Proprietary s

Table 8-3 Comparison of Measured Mark-BW and VANTAGE SH Critical Interface Dimensions, in.

Mark VANTAGE 4

.i Dimension Descrintion M _5_H Standard

[ ]

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8-14 .4 4

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, Table 8-4 Component Vertical LOCA Forces Component LOCA Case Maximum i Load (Lbs) l

[ ]

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Table 8-5 Mark-BW Fuel Assembly Stress Analysis Results for SSE and Combined SSE plus LOCA Conditions Applied LoadM Desien Basis Allowable Load  % Margin Component SSE SSE + LOCA SS_E SSE + LOCA SS_E SSE + LOCA SSE SSE + LOCA

[ ]

t L

i i

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8-16 O O O

j Non Proprietary i

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j Figure 8-1 i

! Horizontal Core Seismic and LOCA Model l

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3 Figure 8-2 i Fuel Assembly Vertical Model J

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,-s 9.0 Long Term Containment Integrity Analysis Evaluation

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9.1 Evaluation Results The Sequoyah Nuclear Plant (SQN)long term containment integrity analysis documented in Reference 9-1 was performed using the LOTIC-I version of the LOTIC Computer Code (Reference 9-2).

The important aspects of the fuel change that potentially impact the analysis are changes in the flow characteristics past the fuel, the reactor coolant system (RCS) operating T , the core stored energy and fuel heat capacity, and the decay heat. Table 9-1 lists these features for the present analysis and for those revised due to the change in fuel design.

Flow testing on both the Westinghouse VANTAGE 5H and the Mark-BW fuel has concluded that the results of the comparison were within the established 15% acceptance criteria for thermal-hydraulic compatibility. Therefore, there are very small deviations in flow characteristics past the fuel. For an ice condenser design, I the peak pressure occurs late in the transient well after the ice bed has melted out. The single effect of small deviations in flow is insignificant relative to analysis results. Total energy content, or total energy available for release to containment, is significant. However, since To remains at 578.2 F, the energy

/'"'x content of the RCS remains the same. I ks A detailed compatibility inspection of the Westinghouse VANTAGE SH and the Mark-BW fuel assemblies was also performed. The inspection found that the mass, material, and dimensions of the fuel assembly components are very similar. It was concluded that the mechanical heat capacity is also very similar. Therefore, there is no reportable difference in the mechanical heat capacity of the fuel. The core stored energy has, however, increased from 3.58 full power seconds (FPS) to [] FPS. This has been determined to add 1.32X10' BTUs to the RCS for release to containment. However, there are margins in the current calculations that offset this small increase. For example, the current basis utilizes the specific SQN Decay Heat Curve as defined in Table 2-2 of Reference 9-1 until the time of steam generator equilibration. The LOTIC code then conservatively determines the decay heat based upon Table 6.2.1-8 of Reference 9-1 after equilibration. If the SQN specific data are also used after steam generator equilibration, it is found that 2.11E+06 BTUs can be removed from the calculation up to the time of ice bed meltout, and 6.00E+06 BTUs can be removed up to the time of peak pressure. This conservatism more than offsets the increased l core stored energy effect. l i

\

b 9-1 1

3

- i

Non Proprietary The metric tons of uranium, the enrichment, and the fuel reload cycle utilized in the latest SON analysis remain bounding for the Mark-BW. Therefore, the SON-specific decay heat curve remains bounding.

In the summary, the effect of including Mark-BW fuel on the latest SON loss-of-coolant-accident mass and energy and the containment integrity analysis has been evaluated. It has been concluded that the latest analysis results remain bounding.

9.2 References 9-1 WCAP-12455, Containment Pressure Calculations with an Extended Containment Spray Pump Diesel Generator Loading Delay for the Sequoyah Nuclear Plant.

9-2 WCAP-8354-P-A, Long Term Ice Condenser Containment Code -

LOTIC Code.

O.

l l

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l l

9-2 O l 1

Non Proprietary

,~s Table 9-1

)

('

' / Comparison of Key Inputs Affected by the Fuel Change Important to the Containment Integrity Evaluation for Sequoyah Current Mark-BW Direction of Parameter Fuel Fuel Conservatism Flow Behavior Through Core {1} {1} O RCS T.,, F 578.2 578.2 +

Core Mechanical Heat Capacity (2} {2} +

Core Stored Energy, FPS 3.58 [] +

Metric Tons of Uranium 72.25 87.255 -

Enrichment, % 3 3 -

Fuel Cycle Length, months 18 18 +

where: 0 = No significant effect

+ = An increased value is conservative

- = A decreased value is conservative

\s/ {1} Flow testing of the VANTAGE SH and Mark-BW fuel assemblies confirmed the compatibility of the assemblies. The result was within the 5% acceptance criteria for thermal-hydraulic compatibility.

{2} The inspection of the VANTAGE 5H and Mark-BW fuel assemblies was conducted relative to the mass, material, and dimensions. No differences were noted that affect the mechanical heat capacity.

(h Ed 9-3

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