ML20101U170
| ML20101U170 | |
| Person / Time | |
|---|---|
| Site: | Browns Ferry  |
| Issue date: | 10/12/1984 |
| From: | Rochelle J TENNESSEE VALLEY AUTHORITY |
| To: | |
| Shared Package | |
| ML18029A361 | List: |
| References | |
| RTR-NUREG-0661, RTR-NUREG-661 CEB-83-34, NUDOCS 8502070120 | |
| Download: ML20101U170 (140) | |
Text
lc CEB REPORT g
REPORT No.
TV A 10752 (EN DES-2-8 3)
CEB-83-34 TITLE I
Browns Ferry Muclear Plant P NT/ UNIT Torus Integrity Long-Term Program 33, 3gcy,ong3, Plant Unique Analysis Report J UNID Y 4.M L )
KEY NOVNS CONTRACT No.
VENDOR Systems Torus Integrity MEDS ACCESSION NUMBER N/A N/A (FOR MEDS USE)
APPLICABLE DESIGN REV U!,J 09D0018ff) C E B '8312 21 008 BFN-50-D706,
-:M 0712 004 no I
840725E0015gp>IB BFN-50-D711 R1
_CEB '841210 00 8 NUREG-0661, CEB-76-23 g
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TENNESSEEVALLEY AUTHORITY DIVISION OF ENGIN'EERING DESIGN CIVIL ENGINEERING SUPPORT BRANCH S
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R1 R2 Revision 0 i
December 21, 1983 7-12-84 12-10-84 Date b*
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Prepared ep[(
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- Engineer,
- Prepared by CEB, NEB, and BWP representatives, J. K. Rochelle, Lead 8502070120 850125 05Mg See Acknowledgment.
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COORDINATION LOG CEB-83-34
Title:
hO Ikhh.hRI PROGRAM PLANT UNIQUE ANALYSIS REPORT Revision:
2 R-Oenotes review A-Denotes approval DIVISION OF ENGINEERING AND TECHNICAL SERVICES (DETS)
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- I BFN-PUAR TABLE OF CONTENTS Page i
ABSTRACT.............................................
Ii TABLE OF CONTENTS....................................
xix LIST OF ILLUSTRATIONS................................
xxiii LIST OF TABLES.......................................
xxvi
' LIST OF ABBREVIATIONS................................
1-1
1.0 INTRODUCTION
1-1 1.1 Objective..................................
1.2 O_riginal Design of Browns Ferry 1-1 Containment Systems......................
I 1.3 Formation of the Mark I Owners Group 1-2 1-2 1.4 Short-Term Program Activitles 1.5 Generic Long-Term Program Activities 1-3 I
1.6 The Browns Ferry Torus Integrity 1-4 Long-Term Program........................
1-4 1.6.1 Hydrodynamic Load Mitigation 1-5R2 1.6.2 Composition of the BFN-PUAR 1.6.2.1 Contents'...............................
1-5R2 I
1-6R2 1.6.2.2 Arrangement............................
2.0 GENERAL DESCRIPTION OF STRUCTURES 2-1 AFTER MODIFICATION............................
2-1 2.1 Drywell....................................
2-2 2.2 Wetwell....................................
_g-2-2 2.3 Vent System................................
g Internal Structures and Piping 2-2 2.4 Torus 2-2 2.4.1 Submerged Structures 2.4.1.1 ECCS Suction Nozzle Strainers 2-2 2.4.1.2 HPCI, RCIC, and RHR Supports 2-3
_I 2.4.1.3 Quenchers and Quencher Support 2-3 Structures...........................
I I
ii PUAR.TC
. I
BFN-PUAR TABLE OF CONTENTS (Continued)
Page
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2.4.2 Partially Submerged Structures 2-3 2.4.2.1 S/RV Discharge Lines 2-3 2.4.2.2 Turbine Exhaust and Return Lines 2-3 2.4.3 Above-Pool Structur_e_s_....................
2-4 2.4.3.1 Catwalk and Vacuum Breaker Platforms 2-4 2.4.3.2 RHR Spray Header 2-4 l
2.4.3.3 Monorail 2-4 5
2.4.3.4 Drywell/Wetwell Vacuum Breaker Valves 2-4 2.5 S/RV Discharge Lines in the Drywel1 2-4 2.6 Torus Attached Piping 2-5 2.7 Active Components 2-5 2.8 Torus Penetrations 2-5 3.0 LOADS FOR STRUCTURAL ANALYSES 3-1 3.1 Original Design Loads 3-1 3.2 Newly De fined IDCA-Induced Loads 3-1 3.2.1 Design Basis Accident 3-1 3.2.2 Intermediate Break Accident 3-4 3.2.3 Small Break Accident 3-5 3.3 Sa fety/ Relief Valve Discharge-Induced Loads 3-5 4.0-GENERAL DESIGN CRITERIA.........................
4-1 4.1 Introduction 4-1 4.2 Load Definitions 4-1 4-1 4.2.l General... ~..
4.2.2 S/RV Hydrodyr.cmic Loads 4-1 4.2.2.1 Interpretation 4-1 4.2.2.2 Jus t i fi ca t ion 4-2 4.2.3 DBA Condensation Oscilletin Hydrodynamic Loads 4-3 l
l 4.2.3.1 Interpretation 4-3R2 4.2.3.2 Justification 4-4 5
I iii PUAR.'IU I
1 I
I BFN-PUAR TABLE OF CONTENTS (Continued)
Page 4-4 4.2.4 Post-Chug Hydrodynamic Loads 4-4
- I 4.2.4.1 Interpretation.........................
4-5 4.2.4.2 Justification..........................
4-5 4.2.5 DBA Pool Swell Ilydrodynamic Loads 4-SR2 l
4.2.5.1 Interpretation.........................
4-6 4.2.5.2 Justification..........................
4-7 4.3 Structural Acceptance Criteria 4-7 4.3.1 General..................................
I 4.3.2 Torus, Drywell, and Vent System 4-7 Pressure Boundary Components 4.3.3 Piping System Components -
4-7RI I
Excluding Supports.....................
4-8 4.3.4 Linear Supports and Snubbers
)
4-8 4.3.4.1 Allowab!c Stress Criteria 4-12 4.3.4.2 Justification..........................
4-12 4.3.5 Variable Spring Supports 4-12RI 4.3.5.1 Objective..............................
4-13 4.3.5.2 Minimum Requirements...................
4-15 4.3.5.3 Justification..........................
4.3.6 Operability and Functionality of 4-15 Components.............................
I 4.3.7 Nonsa fety-Related Internal 4-16 Structures.............................
4-16 4.3.8 Reinforced Concrete Structures 4-16 4.3.9 Fatigue Evaluation.......................
4-17 4.4 Analysis Procedures........................
4-17 4.4.1 General..................................
4-17 4.4.2 Load Combination Techniques 4-17 4.4.2.1 Torus and Vent System..................
I 4.4.2.2 Torus At tached Piping System, S/RV Piping Systems inside the Torus and Other Nonsafety-Related 4-17 I
Internal Structures..................
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PUAR.'IU iy I
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I BFN-PUAR I
TABLE OF CONTENTS (Con t inued)
Page 4.4.2.3 S/RV Piping Systems Inside the l
Drywell and Vent System................
4-17 5
4.4.2.4 Justification 4-18 4.4.3 S/RV Load Reduction Factors 4-18 4.4.4 Torus Analysis Procedure 4-18 4.4.5 Vent System Analysis Procedure 4-19 4.4.6 Torus Attached Piping Systems l
Analysis Procedure 4-20 5
4.4.7 S/RV Piping Systems Analysis Procedure 4-22 4.4.8 Component Operability Procedure 4-23 4.4.9 Other Internal Structures Analysis Procedure 4-24R1 4.5 Construction Code for Modifications 4-25 g
4.6 S/RV Confirmatory Test 4-25 3
4.6.1 Test Objective 4-25 4.6.2 Basic Test Requirements 4-25 4.6.3 Test Report 4-26 4.6.4 Correlation of Test Data With Analysis 4-26 4.7 Permanent Analysis and Design Documentation 4-26 4.7.1 Design Criteria 4-26 4.7.2 Analysis Calculations 4-26 4.7.3 Design Requirements 4-26 4.7.4 References 4-27 g
4.7.5 Design Calculations and Drawings 5
for Modifications 4-27 4.7.6 Sunmary Report 4-27 5.O TORUS CONTAI!IMENT STRUCTURE - ANALYSIS AND MODIFICATIONS.................................
5-1 5.1 General Description........................
5-1 5.2 Torus Modifications 5-1 5.2.1 Dynamic Ring Girder Restraints (Snubbers) 5-2 5.2.2 External Ring Girder Reinforcement 5-3 5.2.3 Tiedowns 5-3 5.2.4 Local Stiffening 5-4 5
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PUAR.TC l-
I BFN-PUAR TABLE OF CONTENTS (Continued)
Page 6.4 Vacuum Breaker / Main Vent End Cap 6-7 Intersection.............................
6-7 6.4.1 Analytical Procedure.....................
6-7 6.4.1.1 Analytical Model 6-7 6.4.1.2 Static and Dynamic Loads 6.4.2 Controlling Load Combinations 6-7 I
6-7 6.4.3 AShE Code Allowables.....................
6-8 6.4.4 Results and Comparisons 6.5 Vent Header /Downcomer Intersection 6-8 6-8 6.5.1 Analytical Procedure I
6-8 6.5.1.1 Analytical Models......................
6.5.1.2 Static and Dynamic Loads 6-8 6.5.2 Controlling Load Combinations 6-8 l
6-8R2 6.5.3 ASsm Code Allowables.....................
6-9 6.5.4 Results And Comparisons l
6-9R2 6.5.4.1 Stress Evaluation......................
6-9 6.5.4.2 Fatigue Evaluation 6-10 6.6 Vent Pipe Drain............................
6-10 6.6.1 Analytical Procedure
-g 6.6.2 Controlling Load Combinations 6-10 6-10 m
6.6.3 Allowable Stress.........................
6.6.4 Stress Results And Comparisons 6-10 6-10 6.6.5 Description o~ Modifications 6.7 Downcomer/Tiebar Intersection 6-11 6-11 6.7.1 Analytical Procedure 6-11 6.7.1.1 Analytical Models 6-12 6.7.1.2 Static and Dynamic Loads I.
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yIii PUAR.TC 4-
I BFN-PUAR TABLE OF CONTENTS (Continued)
Page 6.7.2 Design Load Combinations 6-12 6.7.3 Allowable Stresses 6-13 l'
6.7.4 Results and Comparisons 6-13 5
6.7.4.1 Stress Evaluation 6-13R2 l
6.7.4.2 Fatigue Evaluation 6-13 6.7.5 Descr ipt ion o f Modi fica t ions 6-13 6.8 Vent Column Support Columns 6-14 6.8.1 Analytical Procedure 6-14 g
6.8.2 Controlling Load Combinations 6-14 g
6.8.3 Allowable Stresses 6-14 6.8.4 Stress Results and Comparisons 6-15 6.9 Vent System Miter Bends 6-15 6.9.1 Analytical Procedure.....................
6-ISR2 l
3 6.9.2 Controlling Load Combinations 6-15 g
6.9.3 AShE Code Allowables 6-15 6.9.4 Stress Results and Comparisons 6-16 6.10 Torus Bellows 6-16 6.10.1 Analytical Procedure 6-16 6.10.1.1 Analytical Model 6-16 6.10.1.2 Static and Dynamic Loads 6-16 6.10.2 Design Loading Conditions 6-17 6.10.3 ASME Code Allowables 6-17 6.10.4 Results and Comparisons 6-17 6.11 Vent IIeader Shell 6-17 6.11.1 Analytical Procedure 6-17 6.11.1.1 Analytical Model 6-17 6.11.1.2 Design Loading 6-18 6.11.2 Results and Comparisons 6-18 6.11.3 Description of Modifications 6-18 I
ix PUAR.TC
I BFN-PUAR TABLE OF CONTENTS (Continued)
Page 10.4 Results and Conclusions 10-4 10.5 Description of Temperature Monitoring I
System..................................
10-5 Il-1 11.0
SUMMARY
AND CONCLUSIONS.........................
11-1 11.1 General..................................
11.2 Browns Ferry Design Criteria l
11.3 Structural Analyses and Design of l
11-1 5
Requi red Modi fica t ions 11-2 11.4 S/RV Confirmatory Test 11.5 Ins talla t ion of Modi fica t ions and Il-2R2 l
I Final Conclusions.......................
12.0 REFERENCES
12-IRI APPENDIX A.0 TORUS ATTACHED PIPING ANALYSIS PROCEDURES AND CRITERIA...............
A-1 A-1 A.1 Introduction...............................
A-1 A.2 Scope......................................
A-2 A.3 Definitions................................
.g A-2 5
A.3.1 Essential Piping.........................
A-2 A.3.2 Nonessential Piping......................
A-2 A.3.3 Active Component.........................
A-2 A.4 Analytical Models..........................
A-3 A.4.1 Piping Model Boundaries..................
A.4.2 Torus Interface..........................
A-3 A.4.2.1 Coordinate System......................
A-3R1 I
A-3 A.4.2.2 Flexibility............................
A-4 A.4.3 Process Piping...........................
A-4 A.4.4 Branch Lines.............................
A-5 A.4.5 Valves...................................
A-5 A.4.6 Flanges..................................
A-5 A.4.7 Reducers..................................
I A-5 A.4.8 Supports.................................
A-6 A.4.9 Special Considerations A.4.10 Component Nozzle Attachments A-6 I
I xiy PUAR.TC I
I I
BFN-PUAR TABLE OF CONTENTS (Continued)
I Pa ge A.5 Load Sources A-6 A.5.1 General A-6 A.5.2 Seismic Loads A-7 A.5.2.1 Operative Basis Earthquake A-7 A.5.2.2 Safe Shutdown Earthquake A-7 A.5.3 The rna l Loads A-7R1 A.S.4 Torus Motion and Drag Loads A-8 A.6 Analysis Procedures A-8 A.6.1 Introduction A-8 l
A.6.2 Modeling Assumptions A-8 A.6.3 Deadweight Analysis A-9
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A. 6.' 4 Thernml Load Case Analysis A-9 A.6.5 Scismic Analysis A-9 A.6.6 LOCA and S/RV Analysis A-10 A.6.6.1 Dynamic Inertial Efforts A-10 A.6.6.2 ^ Dynamic Displacements A-Il A.6.6.3 Thernal and Pressure Displacements A-ll
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A.6.6.4 Fluid Motion A-Il A.6.7 Analysis Results A-ll A.7 Load Case Combinations A-ll A.7.1 Introduction.............................
A-Il A.7.2 Combinations Used for Piping, Supports, and Active Component Evaluations A-12 g
A.7.3 Combinations Used for Evaluation of g
Piping System Reactions on Torus Penetrations A-12 A.7.4 Combinations Used for Evaluation of Valve Accelerations A-12 A8 Process Line Evaluations A-12 A.8.1 Code Jurisdiction A-12 A.8.2 Piping Evaluation Procedure A-12 I
xv PUAR.TC
I BFN-PUAR
'S I
TABLE OF CONTENTS (Continued)
Page D.I.l.1 Bubble-Induced Pool Swell Drag Loads D-1 I
D.I.l.2 Downcomer Water Jet Loads D-1 D.I.l.3 Bulk Pool Motion Loads D-2 D.l.2 Condensation Oscillation and Chugging I
D-2 Drag Loads..............................
D.I.2.1 Pseudo Response Spectrum Method D-2 I
D.I.2.2 Equivalent Static Load Method D-16 D.I.2.3 DBA Condensacion Oscillation Drag D-17 Loads.................................
D-17 D.I.2.4 Chugging Drar Loads D-19 D.1.3 S/RV Drag Loads...........................
D.I.3.1 T-Quencher Bubble-Induced Drag D-19 Loads.................................
D-21 D.I.3.2 T-Quencher Water Jet Loads
.E W
D.2 Above-Pool Fluid-Induced Loads D-21 D.2.1 Pool Swell Impact and Drag Loads D-21 I
D-21 D.2.2_
Froth Impingement Loads...................
D-22 D.2.3 Fallback Loads............................
-l APPliNDIX E.0 IMPLEMENTATION OF CONSISTENT FLUID W
MASS MATRIX FOR TORUS MODEL............
E-1 E-1 E.1 Background..................................
I E-1 E.2 Implementation..............................
F-1 APPENDIX F.0 MAJOR COMPUTER CODES G-1 APPENDIX G.0 PIIOTOGRAPIIS..............................
APPENDIX H.0 FINAL TVA RESPONSES 'IO NRC AND BROOKIIAVEN I
NATIONAL LABORATORY QUESTIONS H-GR-1R2
' APPENDIX I.0 FINAL TVA RESPONSES TO NRC AND FRANKLIN RESEARCII CENTER QUESTIONS..............
I-GR-IR2
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xvill PUAR.TC
BFN-PUAR I
LIST OF ILLUSTRATIONS (Continued)
Figure Title 10-6 Local Pool Temperature Response, Case 2AX I
80 Percent Water Fluid Mesh 10-7 Thermowell Orientation Inside Torus 10-8 Schema t ic of BFN Torus Showing Bulk Tempera ture I-Suppression Pool Temperature Monitoring System Sensor Locations 10-9 Block Diagram A-1 Global Coordinate Definition A-2 Local Coordinate Definition for Torus Nozzles A-3 Local Coordinate Definition Detail A-4 Directions for Torus Penetration Spring Constants I
Di f ferent ial Fluid Tempera ture Be tween Torus A-5 and ECCS Ring Header During DBA, IBA, and SBA B-1 Typical Concrete Anchors C-1 Strain Gage Rosette Locations - Torus Shell C-2 Pressure Transducer Locations Acceleration and Displacement Transducer C-3 Locations - Outer Shell C-4 Saddle Gage. Locations I
C-5 RCIC Turbine Exhaust Gage Locations
'C-6 RCIC Vacuum Relief and Test Line Transducer Locations C-7 ECCS Header and Associa ted Gage Locat ions I
C-8 RHR Test and Restraint Gcge Locations C-9 Downcomer Gage Locations C-10 Quencher Support and S/RV Gage Locations E-1 80 Percent Water Fluid Mesh BNL-16-1 Typical Pressure Traces for BFN Torus Shell BNL-22-1 Typical Mid-Vent Bay CrossSection I
xx,
, OAR.Iv g
I BFN-PUAR LIST OF ILLUSTRATIONS (Continued)
Plate Title 1
Torus Snubber and Base Plate; EOCS IIeader Support 2
External Ring Girder Rein forcement 3
Torus Tiedown Sys tem 4
Vent Header Colunm Support Extension g
5 Cradle Reinforcement g
6 EOCS IIeader Penet ra t ion Shell Re in forcemen t 7
Reinforcement o f Vent IIeader and Downcomer 8
Rein forcemen t of Vent Header 9
Downcomer Tlebar and V-Bracing 10 Typical Overhead View of Ramshead/ Quencher /Boxbeam Arrangement 11 Typica l Quencher /Boxbeam Insta lla t ion Showing Quencher Midspan Clamp and End Point Lateral Lug 12 Typical Boxbeam to Ramshead Support Connection 13 Typical Perspec t ive o f S/RV Pipe / Quencher /Boxbeam in Torus Showing Overhead Reroute of S/RV Lines E and M 14 Typical S/RV Long Line Showing Vertical and g
Horizontal Stru ts with Ring Girder 5
Re in forcemen t 15 hbin Vent Reaction Box for Typical Two S/RV Line Penetration 16 Main Vent Reaction Box for Typical Three S/RV Line Penetration 17 Torus Spray IIeader Support 18 RHR Vertical and Horizontal Bracing 19 RHR Support Collar 20 IIPCI Support Collar 21 HPCI. Vertical and Horizontal Bracing l
22 X-204 A Thru D Penetration Support for EOCS IIeader 5
23 Process Suc t ion Line " TEE" Support and ECCS IIcader 24 RCIC Turbine Exhaust First Axial Restraint 25 Catwalk From Below 26 Catwalk Knee Brace and Longitudinal Pipe Connection 27 Ca twa lk, Handra ils, Vacuum Breaker, Va lve Pla t f orm, and Monorail 28 Typical S/RV Pipe Support Inside Main Vent 29 Triple S/RV Line Routing Inside Shin Vent; Replacement o f 700 M!ters with Elbows 30 General View of Work Including V-Brace, RIIR Re t u r n Line, Quencher and Quencher Support, and S/RV S t ru t Mod i fi ca t i on xxii PUAR.IV
I BFN-PUAR LIST OF TABLES Table Title 3
2-1 Principal Design Parameters and Characteristics
!g of Primary Containment l
l 2-2 Torus Attached Piping Penetrations l
2-3 Active Conponents for Operability Evaluation 3-1 Event and Load Combinations j
5-1 Allowable Stresses - Torus Shell and Ring Girder 5-2 Maximum Torus Reactions 6-1 Load Combinations and Service Levels for Drywell/Nbin Vent Intersection 1
6-2 Maximum Stress Intensities on Drywell/ Main Vent Intersection l
6-3 Vent System Analysis Temperature and Pressure (Wetwell) 6-4 Vent System Analysis Temperature and Pressure (Drywell) 6-5 Controlling Load Combinations and Service Levels of hkin Vent / Vent Header Intersection 6-6 Maximum Stress Intensities on Main Vent / Vent Header Intersection 3g 6-7 Controlling Load Combinations and Service Level of Vacuum Breaker / Main Vent End Cap Intersection 6-8 Maximum Stress Intensities on Vacuum Breaker /
Main Vent Intersection 6-9 Controlling Load Conbinations and Service Levels of Downcomer/ Vent Header Intersection 6-10 Maximum Stress Intensities on Downcomer/ Vent Header Intersection 6-11 Fatigue Usage Factors for Containment Vent System 6-12 Controlling Load Combinations and Service Levels of Downcomer/Tiebar Intersection I
Maximum Stress Intensities on Downcomer/Tiebar 6-13 Intersection 6-14 Stress Evaluation of Tiebar I
6-15 Controlling Load Combinations and Service Levels of the Vent Column Supports Stress and Buckling Evaluation on Vent Column Supports 6-16 6-17 Maximum Stress Intensities on Main Vent Miter Bends I
6-18 Maximum Stress Intensities on Vent Header Miter Bends Maximum Stress Intensities on Downcomer Miter Bends 6-19 6-20 Stress Evaluation at Key Locations for Zero AP 6-21 St ress In tensi fica t ion Factors for Vent System Miter Bends I
xxill PUAR.TB
BFN-PUAR l
LIST OF TABLES (Continued)
Table Title 7-1 Drywell Load Combinations 7-2 NOC Service Levels B and C Load Combinations l
7-3 SBA/IBA - Service Levels C and D Load Combina t ions 7-4 DBA - Service Level D Load Conbinations 8-1 Torus Attached Piping Systems 8-2 Limiting Case Event Combinations and Service Levels for Torus Attached Piping, Piping Supports, and Equipment Nozzle Loads 8-3 Limiting Case Event Conbinations and Service Levels for Piping Loads on Torus Shell and Nozzles 3
8-4 Number of Supports Removed, Modified, and Added g
to Unit 1 RIIR Sys tem Torus At tached Piping 10-1 Sumnary of Results - Browns Ferry Suppression Pool Temperature Response 10-2 Summa ry of Resul t s - RHR Hea t Exchanger Flow Rate 10-3 Suppression Pool Temperature Monitoring System Environmental Requirements 11-1 Summary of BFN LTP Modi fica t ions A-1 Spring Constant Sunne ry A-2 Load Cases for Torus Attached Piping Systems A-3 Expected Piping Segment Temperatures for BFN g
A-4 Analysis Criteria for Torus Attached Piping -
E LOCA Effects A-5 Limiting Case Event Combinations for Valve Accelerations C-1 Torus Shell Stress Comparison C-2 Torus Shell Pressure Comparison C-3 Torus Acceleration and Displacement Comparison BNL-7-1 Comparison of FSTF and Browns Ferry DBA CO Responses BNL-16-1 Maximum and Minimum Pressures - Single Valve Actuation Tests BNL-16-2 Maximum and Minimum Pressures - Multiple Valve Actuation Tests BNL-16-3 Initial Gas Volume at Nornal Operating Condition Blowdown I
I xxiv PUAR.TB
BFN-PUAR I
LIST OF TABLES (Continued)
Table Title FRC-7-1 Comparison of FSTF and Browns Ferry Chugging Responses j
I-FRC-20-1 Drywell Load Combinations - Maximum Stress i
FRC-20-2 NOC - Service Levels B and C Load Combinations l
Maximum Stresses - Wetwell Evaluation i
FRC-20-3 SBA/IBA - Service Levels C and D Load Combinations I
Maximum Stresses - Wetwell Evaluation FRC-20-4 DBA - Service Level D Load Combination l
Maximum Stresses - Wetwell Evaluation I-FRC-24-1 Typical Branch Line Analysis Results I
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BFN-PUAR I
LIST OF ABBREVIATIONS ABBREVIATION ABS Absolute Summation I
ACI American Concrete Institute ADS Automatic Depressurization System AE Architect / Engineer AISC American Institute of Steel Construction I
ANSI American National Standards Institute ASME American Society of Mechanical Engineers ASTM American Society of Testing Materials I
BDC Bottom Dead Center BFN Browns Ferry Nuclear Plant BWR Boiling Water Reactor
- C FR Code of Federal Regulations CII Chugging C3N Consistent Mass Matrix 00 Condensation Oscillation I
DBA Design Basis Accident DLF Dynamic Load Factor ECCS Emergency Core Cooling System EN DES Ergineering Design I
EP Engineering Procedure EPRI Electric Power Research Institute EQ Earthquake I
F Fahrenheit FSAR Final Safety Analysis Report FSI Fluid Structure Iteration FSTF Full Scale Test Facility 2
g Acceleration due to gravity (32.2 ft/sec )
GE General Electric Company gpm Gallons per minute I
IIPCI High Pressure Coolant Injection liz lie r t z IBA Intermediate Break Accident ksi Kips per square-inch Ii LDR Load De fin it ion Repor t LOCA Loss-of-Coolant Accident LTP Long-Term Program I
MCR hhin Control Room MS Main Steam MSIV hhin Steam Isolation Valve NOC Nornal Operating Condition I
NRC Nuclear Regulatory Commission I
I PUAR.BV
I DFN-PUAR ABBREVIATION (Continued)
OBE Operating Basis Earthquake Pb Primary bending stress PL Local primary membrane stress Pm General primary membrane stress PIN Pittsburgh Des Moines psi Pounds per square inch psia Pounds per square inch absolute psid Pounds per square inch differential psig Pounds per square inch gage PUAAG Plant Unique Analysis Application Guide PUAR Plant Unique Analysis Report PULD Plant Unique Load De fini t ion Q
Secondary stress due to primary plus bending QSTF Quarter Scale Test Facility RCIC Reactor Core Isolation Cooling RIIR Residua 1 Ileat Remova1 RPS Reactor Primary System l
RTD Resistance Temperature Detectors S
Yield Stress y
S Alternating Stress Intensity a
S Spectral Acceleration an SAC St ructura l Accep tance Cri teria SBA Small Break Accident SER Sa fety Evaluation Report SORY.
Stuck Open Relie f Valve SRSS Square Root of the Sum of the Squares S/RV Sa fety/Relie f Valve SSE Sa fe Shu tdown Ear thquake STP Short-Term Program TES Toledyne Engineering Services
'DN Tr ibu tary Mass Ma tr ix TVA Tennessee Valley Authority WRC Welding Research Council g
ZPA Zero Period Acceleration g
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g m R.ov
I BFN-PUAR I
Truncation of the downcomers reduces pool swell loads.
Addition of S/RV quenchers and RHR return elbow discharge devices ensures stable steam condensation during S/RV
-g blowdowns for all postula ted accident conditions.
The g
quenchers also mitigate S/RV discharge loads on the tori for all normal and pos tula ted LOCA even t s.
- Finally, addition of the 10-inch S/RV vacuum breakers reduces water I
clearing loads on the S/RV piping systems for rapid second actuation conditions.
The basic functional requirements for these modifications were defined from generic and plant unique information provided by the Mark I Owners Group and GE and approved by NRC.
.The other generically approved load mitigation methods were inappropria te or unnecessary for DFN.
1.6.2 Compos i t ion o f t he BFN-PUAR 1.6.2.1 Contents BFN containment systems are described in Section 2.
The new hydrodynamic loads for structural analysis of those systems are summarized in Section 3.
Structural analysis of the BFN containment systems and structural design of the necessary
'E Piaat modi fications were performed according to the BFN LTP 3
general design criteria described in Section 4.
The five basic categories of structural analysis and design activity are described in Sections 5 through 9.
I An evaluation of the bulk and local pool temperatures for various postulated accident conditions was conducted as required by NUREG 0661.
Section 10 summarizes the results I
of that evaluation and describes the new pool temperature monitoring system for each BFN unit.
Section il gives a general summary and status of BFN LTP and rela ted modi fica t ion act ivi t les upon submi t tal of this report for NRC review (on approximately December 31, 1983).
It also draws basic conclusions regarding completion of BFN LTP activities for all three BFN units.
Additional information on structural analysis and design methods, as well as confirmatory postmodification S/RV test results, are given in Appendices A through F.
Appendix G contains construction photographs of some major BFN LTP modi fica t ions.
Appendices II and I contain questions and
'I responses resulting from review by NRC's consultants, Brookhaven National Laboratory and Franklin Research Center.
I I
1-SR2 PUAR.1 I
I l
BFN-PUAR 1.6.2.2 Arrangement Th e Ta ble o f Con t en t s, beginning on page ti, lists the I
headings and subheadings of each section and appendix.
It E
locates the List of Tables, List of Illustrations, and List of Abbrevia tions, as well as References.
The. title page of each section and appendix is brown, to provide easy access.
The text of each section and appendix is numbered separately.
For example, page 2-9 is the ninth page in Section 2.
Illustrations include both figures and pla tes (photographs).
All figures and tables are located at the end of each section and appendix, with the exception of Appendices 11 and I.
For example, Figure 2-1 is the first figure in Section 2 and Table A-2 is the second table in Appendix A.
Plates are l
located in Appendix G.
5 Appendices H and I are arranged in accordance with TVA's g
October 11, 1984, response to NRC's request for additional g
information on the BFN PUAR.
For each item the arrangement is:
1)
Questlon/ request from NRC's consultant 2)
TVA's response including tables and figures Page, table, and figure numbers correspond to the item being addressed.
For example, page 1-FRC 2-1 is the first page for Franklin Research Center Item 2, and Table BNL 18-3 is j
the third table for Brookhaven Na tional Laboratory Ite n 16.
3 I
I I
I I
I 1-6R2 PUAR.I r
I BFN-PUAR The QBUBS02 code predicts shell pressures which envelop both rigid-wall and flexible-wall test data, it also predicts conservative attenuation rates with time which produce
' I conservative dynamic amplification of torus motion inputs to piping systems and other components attached to the torus.
The first mode frequencies of the vent system downcomers in I
the longitudinal and transverse directions were within both the frequency range of S/RV discharge air bubbles.
Therefore, use of the conservative load attenuation with I
t ime according to computer code 'IQFORBF would resul t in
. excessively conservative predictions of downcomer The more realistic attenuation rates of TQFOR03 responses.
I provided a reasonable drag load definition for combination with other downcomer loads and design of downcomer bracing modi fica t ions.
SRSS of multiple valve effects in combination with the conservative aspects of this load definition produced a reasonable analysis and design approach.
Absolu te sunma t ion I
of multiple valve ef fects would be excessively conservative because there are 13 S/RV lines discharging into 16 bays of each BFN torus (see Figure 7-3).
Both single and multiple velve tests were run in the S/RV; confirmatory test, thereby verifying the overall load interpretation and analytical approach.
Load reduction factors were conservatively defined based upon correla tion
'I of both single and multiple valve test results (Appendix C).
4.2.3 DBA Condensation Oscillation Hydrodynamic Loads 4.2.3.1 Interpretation The torus was analyzed for shell pressure harmonic forcing functions at 1-Hz intervals from 1 to 30 Hz.
Forcing functions above 30 Hz were neglected.
Referring to Table 4.4.1-2 and Figure 4.4.1-1 of the LDR
- (Reference 14), the largest input pressure coef ficient for each 1-Hz interval was selected from the three alternatives.
The response ' for each interval was determined on the basis I
o f maximum response for any f requency within the 1-Hz band.
Then the responses were combined by the following procedure:
[
1)
The responses for four forcing functions (at 4-5 Hz, 5-6 Hz, 10-11 Hz, and 15-16 Hz) were added absolutely.
I.
4-3R2 PUAR.4 I
I BFN-PUAR 2)
The responses for the other 26 forcing functions were combined by SRSS.
3)
The results of 1) and 2) were added absolutely.
DBA condensation oscillation (CO) drag load responses for g
each structural mode were determined by the same procedure.
E-4.2.3.2 Justification This interpre ta t ion wa s developed early in 1980 on the basis of Full Scale Test Facility (FSTF) da ta analysis by GE and Mark I LTP consultants.
That data analysis indicated that input above 30 liz is of such low energy content as to be negligible in determining torus response.
Further, the forcing functions were found to have little or no phase relationship to each other.
Very loose phase relationships were seen by one study for forcing functions at 5-6, 10-11, and 15-16 IIz, whereas a more de finit ive study, Re ference 19, showed essentfally random phasing of all forcing funetions.
The procedure ou tlined above recognized the remote possibility of cons tant phase relationships between forcing f une tions a t 5-6, 10-11, and 15-16 Hz.
It aIso recognized the random phasing between all other forcing functions and assured that the desired 84 percent nonexceedance probability was achieved.
Additional conservatisms which were inherent to the BFN DBA CO analysis methods are described in Section 5.4.2.9 and Appendix D.
This i n te rp re ta t ion reduced the total calculated response by a fac tor o f 2 or more rela t ive to absolu te summa t ion o f maximum responses for all 50 forcing func tions defined in the LDR.
Therefore it eliminated excessive conservatism but ensured a sa tis f actory nonexceedance probability of the predicted dynamic responses.
4.2.4 Po s t -Ch u sr livdrodynamic Loads 4.2.4.1 Interpretation The torus wa s ana lyzed for shell pressure harmonic functions at 1 -IIz intervals from 1 to 30 liz.
Forcing functions above 30 IIz were neglec ted.
The dynamic response for each of the 30 : forcing funct ions was calcula ted separa tely on the basis of maximum-response for any f requency within the 1-IIz band.
Then the responses were combined by absolute summatlon.
I 4-4 PUAR.4
1 I
I BFN-PUAR i
Post-chug drag loads were defined and analyzed for harmonic forcing functions at 1-Hz intervals from 1 to 50 Hz.
The dynamic responses for each interval were determined on the I
basis of maximum response for any frequency within the 1-Hz l
i band.
Then the combined response for each structural mode was determined by absolute sumnation of the response for the five largest input coefficients plus SRSS of the other 45 I
responses.
4.2.4.2 Justification This interpretation was justified by analysis of FSTF data as documented by Reference 20.
The procedure for torus analysis.was established and the analysis was performed I
before completion of Reference 20.
By relating Reference 20 results to those obtained by this procedure it was clear that the desired 84 percent nonexceedance probability g
E
'**poa** "a" attaia'd-(S** Sect'oa 5 4.2.11 for a more J
detailed discussion of this topic.)
The procedure for post-chug drag load on submerged structures I
.is in compliance with the reconinenda tion o f Re f erence 20 for 84 percent nonexceedance probability loading.
Appendix D gives a detailed discussion of the BFN fluid drag load analytical method and identified conservatisms Inherent to that method.
This interpretation reduced the analytically predicted responses by a factor of 2 or more relative to the absolute sumna t ion o f responses for all 50 inputs.
Therefore, it
-I-eliminated significant excess conservatism from the. load definition, but preserved the desired nonexceedance probability.
.E 5
'4.2.5 DBA Pool SwelI Hydrodynamic Loads 4.2.5.1 Interpretation The torus was analyzed for average hydrodynamic pressure loads as de fined by the PULD and LDR Section 4.3.2.
A I
6.5 percent margin was added to predicted responses to account for uncertainties in the test data for both operating and zero AP cases.
1 The vent system and S/RV piping systems in each torus were analyzed for pool swell impact and drag loads at operating and zero AP conditions, as defined by RO of the PULD and the l
LDR.
Zero AP velocity, displacement, and circumferential time I
4-SR2 PUAR.4 I
I
I BFN-PUAR I
delay ;urves were defined from the 1/4 scale BFN test results.
Resulting vent support column reaction time g
histories for each condition were applied to the torus model 5
in combination with the corresponding pool swell average pressure loads, prior to addition of the 6.5 percent uncertainty margin described above.
Pool swell impact and drag loads for other internal structures were analyzed for one enveloping load case in accordance with LDR Section 4.3.4 and the PULD.
4.2.5.2 Justification This interpretation was based upon the fact that the 1/12 scale Electric Power Research Institute (EPRI) 3-dimensional test model was a prototypical model of BFN in every g
significant detail and it was consistent with the BFN 1/4 g
scale model.
This fact eliminated the majority of NRC's concerns expressed in NUREG 0661, which led to specification of an additional 15 percent upload margin and definition of g
an enveloping longitudinal time delay and velocity 3
distribution for the vent system and other above-pool structures.
The BFN torus was analyzed with a constant effective added fluid mass equal to 80 percent of the total contained fluid mass.
The 6.5 percent margin exceeded one standard E ',
deviation of the BFN 1/4 scale test data.
These 3
considerations ensured an upper bound prediction of torus response, particularly during the upload phase.
(See g
Section 5.4.2.7 for a more detailed discussion in this I
regard.)
The vent system and S/RV piping systems pool swell impact E
analysis in the unmodified and modified conditions was 5
completed well before release of NUREG 0661 and subsequent revision of the BFN PULD.
The interpretation defined above
.a was more accurate for BFN than that identified by NUREG 0661 E
Appendix A and it predicted higher impact velocities for the critical regions o'f the vent header and S/RV piping.
Therefore, reanalysis of the BFN vent system and S/RV piping 3
for revised longitudinal variations was not necessary or 5
appropriate.
(See Sections 6 and 7 for more discussion of the vent system and S/RV piping analyses.)
Other above-pool structures were conservatively analyzed for one enveloping pool swell impact and drag load case in accordance with NUREG 0661.
(Appendix D describes the g
specific analytical method which was applied.)
3 I
4-6 PUAR.4
I BFN-PUAR 6.4 Vacuum Breaker / Main Vent End Cap Intersection located on the end cap of the main The vacuum breaker valves vent pipes are evaluated in the following paragraphs.
- I Figure 6-6 shows the vacuum breaker / main vent intersection.
6 6.4.1 Analytical Procedure I.
6.4.1.1 Analytical Model intersection was modeled using the TPIPE The vacuum breaker I
computer program (see Appendix F) for consideration of reac tions induced by pool swell vent response and coincident loads.
A set of shell pipe intersection spring rates was I
calculated using Bijlaard procedures from Reference 64.
reactions from the TPIPE model, the shell Using the outputthe vacuum breaker penetration were determined stresses at using the WERCO computer program.
6.4.1.2 Static and Dynamic Loads the load experienced by the vacuum breaker Due to location, from the loading imposed on structures valve varies The vacuum breaker elevation is such previously discussed.
level inside the torus.
This that it is above the water I
significantly reducesS/RV loads, CO loads, and chugging loads the number of phenomena that will act upon the valves.
are insignificant at this location, leaving only the effects from 9001 swell, deadweight, and seismie.
Time history data
- 3 for the pool 5
generated by Bechtel Corporation (Reference 17)into the TPIPE swell impact loading analysis was input model.
l 6.4.2 Controlling Load Combinations The number of controlling load combinations required to reduced to one.
Table 6-7 evaluate this component was identifies that combination.
6.4.3 ASME Code Allowables The allowable stress intensities for SA-516 GR 70 carbon steel are shown in Table 6-8.
This is consistent with the I.
material composition at the temperature provided in Table 6-3.
I I
I PUAR.6 6-7 I
I' I
BFN-PUAR I
6.4.4 Results and Comparisons The stress experienced by the vacuum breaker / main vent intersection was in large part due to the pool swell impact load.
Thus, only combination event 18 from Table 3-1 was required for analysis as seen in Tables 6-7 and 6-8.
The calculated stress intensity of 16.7 ksi was well below the Service Level B allowable of 28.95 ksi.
6.5 Vent Header /Downcomer Intersection The vent header and downcomer pairs typically intersect as shown in the finite element representation which simulated this intersection (Figure 6.6).
6.5.1 Analytical Procedure 6.5.1.1 Analytical Models The. vent header /downcomer intersection was modeled into 450 and 1800 beam models for the purpose of evaluating loads described in Section 6.5.2.
Flexibility constants were input at adjacent nodes which, when connected, formed a g
short beam portraying the spring rate of the intersection.
5
- Analysis output forces and moments from the beam models were input to a STARDYNE computer code finite element plate and shcIl model as shown in Figure 6-6.
The fine mesh of
-elements extending around the intersection served to obtain
. accurate stress output.
6.5.1.2 Statle and Dynamic Loads The loading conditions to which the vent header /downcomer intersection was subjected are identical to those outlined in Seetion 6.2.1.2.
6.5.2 Controlling Load Combinations The 27 load combinations were reduced to the controlling combinations shown in Table 6-9.
6.5.3 ASNE Code Allowables The material-composition of the vent header /downcomer 3
intersection is SA-516 GR 70 carbon steel.
Stress intensity 5
allowable values are listed in Table 6-10.
l I
I 6-8R2 PUAR.6
I BFN-PUAR 6.5.4 Results and Comparisons The results and comparisons presented in this section ate I.
products of stress and fa tigue evaluations.
The fa t igue analysis provides in forma t ion for DBA, SBA, and IBA conditions.
-6.5.4.1 Stress Evaluation The downcomer/ vent header stress evaluation was completed I
using the controlling combinations in Table 6-9.
The intersection was most affected by the PL + Q stress category during the DBA condensation oscilla tion event from Case 21.
The calculated stress intensity of 57.6 ksi, compared to an I
allowable of 57.9 ksi, could be further reduced by removing thermal expansion since it is a one-time occurrence.
The most critical pool swell event combination resulted in a l
stress intensity of 44.4 ksi as compared to a 45.2 ksi allowable.
The SBA chugging combination 15, realizing a PL + Q stress intensity of 53.3 ksi, could also be reduced by removing the thermal loads as previously mentioned.
6.5.4.2 Fatigue Evaluation The ASME Code for Class MC requires that a component or structure be. evaluated to demonstrate adequate margin against fa tigue damage in a cyclic load environment.
The approach for this evaluation is to compare maximum stress I
cycle histogram components with conservative strain cycling fa tigue da ta.
The strain cycling data is defined by the fa tigue curve in Figure I-9 of the Appendices to the ASME
~I Code.
This figure plots the alternating stress intensity (Sa) against the number of allowable cycles which may occur for that particular stress intensity.
An analysis for cyclic service is not required for a vessel, component, or I
structure, provided that Paragraph NE-3221.5d is satisfied for all conditions.
The only components of the vent system requiring further evaluation were the downcomer/ vent header I-and downcomer/tiebar intersections and the torus bellows /
main vent connection.
These three portions of the vent system were critical because of their discontinuity and stress concentration characteristics which resulted in high
.I.
localized stresses.
Since the occurrence of one accident condition (DBA, SBA, or
'I IBA) and cumula t ive nornal load occurrences was postula ted in'the fa t igue life of the vent system, all three LOCA I
'I 6-9R2 PUAR.6 m
I BFN-PUAR events were examined.
Table 6-11 presents the usage factors compared with the allowable fatigue usage.
Note that thermal transient through-wall stresses were not included in E
the fatigue evaluation since that stress profile would occur 3
only one time in the design life.
6.6 Vent Pipe Drain The vent drain is located at the lowest elevation of the head at the end of the main vent inside the wetwell.
The g
drain extends into the water and must be able to withstand 3
the hydrodynamic and accident related loads resulting from a LOCA.
The following sections confirm the fact that the drain and modified support configuration shown in Figure 6-8 and described in Section 6.6.5 are qualified.
6.6.1 Analytleal Procedure The vent drain and support were modeled into STARDYNE and a modal analysis was performed.
It was determined that the dominant frequency (35.1 Hz) was in a key range for chugging.
The resulting loads were evaluated by simple hand calculations.
6.6.2 Controlling Load Combinations Since the frequency of the vent drain and support is out of the range of the condensation oscillation event, it was determined by inspection that either combinations 11, 16, 18,.or 25 would control.
Because the structure would be in resonance with key post-chug frequencies, combination 11 g
(SBA + S/RV + CH) actually controls.
3 6.6.3 Allowable Stress The ellowable stress for SA-333 GR B carbon steel is 0.66 times the yield stress, or 23.1 ksi, in accordance with Section 4.3.4.
6.6.4 Stress Results and Comparisons The calculated stress in the new support structure is 21.1 ksi.
This is less than the allowable value of 23.1 ksi.
- 6. 6. 5 ' Description of Modifications The vent pipe drains were truncated to the same elevation as the vent header downcomers, i.e., three feet below minimum pool level.
The existing support for each drain was removed I
PUAR.6 I
6-10
BFN-PUAR 5
6.7.3 Allowable Stresses S
The downcomer/31ebar intersection is made of SA-516 Grade 70 g
=
carbon steel.
Table 6-13 compares actual stresses derived I
i from the analyses with the ASME and AISC Code allowable Z=
stresses per Section 4.3.2 and Section 4.3.4.
y 6.7.4 -Results and Comparisons
=
6.7.4.1 Stress Evaluation g
.a I
The downcomer/tlebar intersection was analyzed to Service i"
Level B for Cases 15, 21, 25, and 27.
From Tables 6-12 l
y and 6-13, the greatest primary plus secondary stress occurs
!gg I
during the SBA event.
The maximum calculated stress W
intensity was 53.8 ksi, compared to the strass allowable of 57.9 ksi.
The largest primary local membrane stress M
I intensity of 27.9 ksi also occurred during the SBA event JM for Case 15, as compared to a 28.95 ksi allowable stress.
$46 The tiebar itself was further analyzed as a linear support for the loads described in Tables 6-12 and 6-13.
As seen I
from Table 6-14, the most severe stress occurred during the CO event combination at a level of 20.7 ksi.
This is below l
C the allowable of 0.66 times the yield stress, or 23 kai.
7 I
6.7.4.2 Fatigue Evaluation y
I The fa t igue evaluat ton of the downcomer/ t lebar intersection is comparable to the evaluation discussed in Section
_g 6.5.4.2.
It can be seen from Table 6-11 that the usage factors are well below 1.0, as required.
=y a
6.7.5 Deseript ton of Modi fica tions
=am M
A new tiebar with V-bracing members was required between cach downcomer pair to minimize downcomer lateral response induced by condensation oscillation ef fects.
As a result, 3
I bending stresses in the vent header /downcomer intersection
==ig reduced and no further reinforcement of that area was 4
are e
required, except as described in Section 6.11.
A$
I The tiebar was installed at elegation 534'-0" and the V-bracing members intersect the tiebar at midspan.
s Fabrication consisted of 3-and 4-inch schedule 40 pipe for the tiebars with short sections of 3-1/2-inch schedule g
I 40 pipe for the bracing at the downcomer end to facilitate s
field adjustability.
All pipe material i s AS1N A 53 Grade B or A 106 Grade B.
I a!!E ST I
S 3
6-13R2 PUAR.6 E
I
I BFN-PUAR I
The connections to the downcomer shell were reinforced using pad plates rolled to match contour and wrapped 1200 The plates were 1/2-inch thick by 7 inches wide for the tiebar connection and S/5-inch thick by 7 inches wide for the bracing connection.
Additional 3/8-inch gussets and small pad plates were provided at the tiebar ends to g
distribute the loads for adequate structural integrity.
A 3
3/8-inch thick saddle plate was provided at the bracing to tiebar intersection to distribute stresses.
All plate material, is ASME SA-516 Grade 70.
For the configuration, see Figure 6-9 and Plate 9.
6.8 Vent Header Support Columns 6.8.1 Analytical Procedure The vent header support columns are 8-inch diameter double l
W extra strong pipes attached to the vent header miter bends vla pipe collars.
Figure 6-2 shows the support columns as they are represented in the computer model.
The beams 3
extending from the straight portion of the columns to the 5
-vent header are rigid, representing the minimal flexibility of the collars.
6.8.2 Controlling Load Combinations The 27 design load combinations were reduced to three controlling cases given in Tables 6-15 and 6-16.
6.8.3 Allowable Stresses The vent column supports are constructed of 8-inch double extra strong A 53 Grade B piping.
Table 6-16 compares actual stresses derived from the analysis with AISC stress allowables per Section 4.3.4.
6.8.4 Stress Results and Comparisons The calculated stresses indicated by Table 6-16 are less than the-allowable stress of 16.6 ksi.
The combination of events in Table 6-15-are composed of Service Level C loads 3
and' compared against Service Level B allowables.
The most g
. severe case is due to pool swell impact which imposes a stress of 10.4 ksi.
The buckling check evaluates the maximum axial load plus bending moment and shows the highest combined effect occurring for pool swell as expected.
The maximum buckling factor is 0.63 compared to a Service Level B allowable of 1.0.
I 6-14 PUAR.6
BFN-PUAR
,l e
6.9 Vent System Miter Bends There are three structural areas in the vent system at which I
miter bends are located.
The main vent, vent hea6er, and downcomer miter bends are all Class MC components.
- However, for analysis purposes these items lend themselves more to
'I treatment as piping components.
P..lgraph NB-3690 was i
introduced for stress evaluation wt. ale retaining the Class i
MC allowables.
The following subsections summarize the t
analysis of these three types of miter bends.
l 6.9.1 Analytical Procedure
'I As mentioned above, the ASME Code provides a guide for the evaluation of miter bends.
The modeling of the bends is indicated in Figure 6-2.
Stress intensification factors are presented in Table 6-21.
These stress in tensi fica t ion l
I factors were calculated using the vent system beam model in conjunction with results from the Bechtel analysis, which provided detailed modeling of the miters in question.
.I Maximum primary plus secondary stresses were ratioed to the nominal section stresses generated from the beam model thereby defining the stress intensi fication factors.
6.9.2 Controlling Losd Combinations The three controlling load combinations are shown in Tables 6-17, 6-18, and 6-19.
6.9.3 ASME Code Allowables The ASME Code allowable stresses are presented in Tables 6-17, 6-18, and 6-19.
6.9.4 Stress Results and Comparisons The main vent miter bend was evaluated for three load cases using Service Level C loads and Service Level B allowables.
I As seen from Table 6-17, these stresses are considerably below the allowables, showing a maximum value of 3.9 ksi, compared with a 28.95 kai allowable for the condensation I
oscillation event combination.
In the same manner, the vent header miter bend and the downcomer miter bend were evaluated for three cases, each I
involving CO, chugging, and pool swell (Tables 6-18 and 6-19).
The maximum local membranc stresses for the Service I
6-15R2 PUAR.6 I
I 1
BFN-PUAR Level C loading on the miter bends were found to be 18.6 ksi and 20.7 ksi, respectively.
When compared with the Service Level B allowable of 28.95 ksi, these areas are qualified.
6.10 Torus Bellows The torus bellows are flexible expansion joints allowing movement of the main vent pipes through the torus wall while maintaining the required pressure boundary.
The analysis performed on this structure was done in accordance with Standards of the Expansion Joint Manufacturer's Association, Inc., (Reference 24).
Fatigue life is the dominant concern.
6.10.1 inalytical Procedure 6.10.1.1 Analytical Model The flexibility o' the bellows was the concern in 0
accurately representing the bellows in the 450 and 180 beam models.
In Figure 6-2, local springs were inserted.
Output from the 450 and 1800 models in the form of displacements was extracted from the various loading events.
The combination of these cases as described in Section 6.10.2 was then used to calculate stresses in the bellows.
6.10.1.2 Static and Dynamic Loads The loading events to which the torus bellows are subjected are explained in Section 6.2.1.2.
6.10.2 Design Loading Conditions The controlling loading conditions were provided by Reference 21 for SBA, IBA, and DBA events.
6.10.3 ASME Code Allowables The ASME Code makes reference to bellows in Paragraph NE-3365.
Standards of the Expansion Joint Manufacturer's Association, Inc., offers a more straightforward and acceptable approach for fatigue evaluation of bellows.
6.10.4 Results and Comparisons liesul ts of the fatigue evaluation are shown in Table 6-12 of Section 6.5.4.
No significant usage factor is observed.
Therefore, the fatigue failure of these components does not present a significant concern.
I I
PUAR.6 I
6-16
I I
TABLE 6-9 CONTROLLING LORD COMBINATIONS AND SERVICE LEVELS OF I
DOHNCOMER/ VENT HEADER INTERSECTION I
EVENT COMBINATION SERVICE LEVEL l
I 27 DBA + SSE EQ + S/RV + CO B
I 15 SBA + SSE EQ + S/RV + CH B
I I
I I
TABLE 6-10 MAXIMUM STRESS INTENSITIES ON DOWNCOMER/ VENT HEADER INTERSECTION g
EVENT _
STRESS CATEGORY _
STRESS ALLONABLE P+Q 57.60 KSI 57.9 KSI l
21 L
14.4 37.6 27 Pt I
44.4 45.2 25 PL 31.4 37.6 18 PL P+0 53.3 57.9 5
15 L
21.8 37.6 15 PL I
I I
I I
I I
TABLE 6-11 FATIGUE USAGE FACTORS FOR CONTAINHENT VENT SYSTEM I
I I
I COMPONENT E
U GE G
1.0
.559
.610 E DE TERSE TION SN 1.0
.107
.353 E
ION INES 1.0
.000
.000 T
I I
I I
I I
I I
I
I I
TABLE 6-12 CONTROLLING LORD COMBINATIONS AND SERVICE LEVELS OF 3
00HNCOMER/TIEBAR INTERSECTION I
EVENT COMBINATION SERVICE LEVEL B
27 B
l 21 DBA + SSE EQ + C0 25 DBA + SSE EQ + S/RV + PS B
I 15 SBA + OBE EQ + S/RV + CH B
I I
I I
TABLE 6-13 MAXIMUM STRESS INTENSITIES ON DOHNCOMER/TIEBAR INTERSECTION g
EVENT _
STRESS CATEGORY _
STRESS ALLOHABLE 21.7 KSI 28.95 KSI l
21 PL 30.6 34.74 27 PL 21 PL+Q 30.9 57.9 l
l 4
28.4 28.95 25 Pt 15 PL+Q 53.4 57.9 27.97 28.95 15 PL I
I lI I
I 1
I I
TABLE 8-14 l
STRESS EVALUATION ON TIEBAR EVENT STRESS CATEGORY STRESS ALLOHABLE 20.7 KSI 23 KSI 27 PL 8.7 23 g!
25 PL 18.0 23 3
15 PL I'
I I
I I
I l
I I
I I
l I
~I I
TABLE 6-15 I
CONTROLLING LOAD COMBINATIONS AND SERVICE LEVELS OF THE VENT COLUMN SUPPORTS I
EVENT COMBINATION SERVICE LEVEL 27 DBA + SSE EQ + S/RV + CO B
l I
I I
I TABLE 6-16 STRESS AND BUCKLING EVALUATION ON VENT COLUMN SUPPORTS l
BUCKLING STRESS I
EVENT CATEGORY STRESS ALLOHABLE FACTOR 7.1 KSI 16.6 KSI
.43 27 PL 10.4 16.6
.63 25 PL 7.6 16.6
.46 15 PL I
I I
I I
I
I I
TABLE 6-17 s
MAXIMUM STRESS INTENSITIES ON 5
MAIN VENT MITER BEND I
STRESS SERVICE EVENT CATEGORY STRESS LEVEL ALLOWABLE 3.9 KSI B
28.95 KSI 27 Pt 1.8 8
28.95 25 PL 1
2.4 B
28.95 15 PL I
I TABLE 6-18 a
MAXIMUM STRESS INTENSITIES ON 5
VENT HEADER MITER BEND I
STRESS SERVICE EVENT _
CATEGORY STRESS LEVEL ALLOHABLE 18.0 KSI B
28.95 KSI 27 PL 25 PL 15.5 8
28.95 15 PL 18.7 8
28.95 I
I 4
TABLE 6-19 a
MAXIMUM STRESS INTENSITIES ON 5
00HNCOMER MITER BENDS I
STRESS SERVICE EVENT CATEGORY STRESS LEVEL ALLONABLE 16.5 KSI B
28.95 KSI 27 PL l
25 PL 13.8 8
28.95 20.6 B
28.95 l
15 PL I
I I
TABLE 6-20 I~
STRESS EVALURTION RT KEY LOCRTIONS FOR ZERO AP (LORD COMBINRTION P + H + T + PS ZERO AP)
WMIDBRY I
lts' %oml== ao y = oAml==
[
s.
4
-s es f
/
/
5
\\
('
~
~
h E
F
- r. VENT C
e I
Q =P.T LOCATIM INTBS TY STE 96 1
STMES Il I
W
- L+#H <1.0
~ 55 00TTm PM 12.0 41.8 MI m 13 pL+Pb 23.1 82.5 LA WL
- L 12.0 25.5 3.1 18.9 0.7 H
Ty PM 8.0 41.8 PL+Pb 17.8 S2.5 L
4.6 25.5 0.2 I
PM 9.0 41.6
- H 0.5 16.9 BOTTW BETEEN (Aj PL+Pb 41.2 82.5 I
L 11.0 25.5
- H 0.7 16.9 0.5
~
PM 15.5 41.6 90TTOM BETW EN
'I.
30 M
pL + Pb 24.1 S2.5 13.9 22.5 L
'I H
0.7 13.6 0.7 BOTTm Pg 17.4 41.6 41.6 82.5 PL+Pb B. Den 2062 17.2 22.5 I
H 3.4 13.6 1.0 L
OM PM 5.3 41.6 PL+Pb 19.4 62.5 I
N 3062 5.0 22.5 L
H 1.7 13.6 0.3 I
I
,a 4
a
-.4 s
a 4
I I
TABLE 6-21 STRESS INTENSIFIC ATION FACTORS l
LOCATION FACTOR MAIN VENT MITER BEND 3.85 3
VENT HEADER MITER BEND 8.2 5
DOWNCOMER MITER BEND 3.82 I
I l
I I
I I'
I i
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I I
I I
I I
E E
E E
M M
M M
g g
g g
g g
g 9
9 9
9 6
mm.awm-g g
g g
g 88%E983GaMEEE
=
a w
EE-mea?PE"BE ZOEFga"smR"-'
EMr m"$dM" MSS mm gzg emogrmrg mm x
x x
x DEADHEIGHT a
gQm E3 Eh$$SE3 nY d
k
-m mg om a
x x
x x
PRESSURE 32$so3E-ogj
$2 "E5sdE2PREE"?
Fy 5*C'5EEEEM5224 7seagat (1) r-2 520E*r5Mm x
o tn A4 r
2"dE5,04?
a GP (2)
T1 8 !mFj?
E x
x x
NOC S/RV
=598$
n EP5bg oq
$355$ 3$ gdze x
IBA S/RV (3)(6) hl E5 502 En E'm-(SUBSEQUENT ACTUATION, y6^
i pg STEAM IN DRYMELL) x2 gx5-hE$g
..,,E ma EgE m
x OBE I43 4550 g
H mEE m
r55 BIzsEPzE x
x x
SSE I43 GEE *d3% @$
F9ERP 86 *M>o rormg n POOL SHELL (5) grg M Eg Fy X
- MW9Rz -5 m
mar-w F
o o
o m
SERVICE LEVEL 5 s' i
zg
Yo?Q 0
0 0
0 ro
==cm m
m m
m gpg<
m m
m m
m
~ ~og
+
w m
- sa 3aa
@z x
x x
x DEADHEIGHT 08 m@g n,
mm H
zCm
- j x
x x
x PRESSURE Em 3 og;5 D$ W 02 Em" NOC S/RV (13 x
x iST. ACTUATION 80 9*
$"h EVENTS I: A y
x58 Br i
NOC S/RV (11 h
N x
x mEd 2M) ACTUATION
--I r m
EVENTS Bu) n$c m' "
$tD mmm arm x
x SSE o
o e
m SERVICE LEVEL ja i
-s n-m, x
-we
-4,--
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WH I
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Tcv wisv nev g :
?5%
~
I
/,y na.a dtecst?ci I
caso
~
g-
,y 3
~
g n -t l l =""
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Sygm g
- -- - - - H I
FIGURE 10-1 COUPLED REACTOR AND SUPPRESSION POOL MODEL I
I
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\\
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\\
9 l
I N
8T [10
/
\\
ii i
z s
s I
e iz s.,
-h utsak
~
g RHR DISCHARGE
~ ~ ~.
x
/
\\
I 2
ie s
i
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FIGURE 10-2 I
PLAN VIEW OF BROHNS FERRY SUPPRESSION POOL HITH T-QUENCHERS AND RHR DISCHARGE LOCATIONS USED IN THE LOCAL POOL TEMPERATURE HODEL I
I 1.
l.y..
DFN-PUAR J
I
SUMMARY
AND CONCLUSIONS 11.O 11.1 General I
The BFN Torus Integrity LTP has been underway since 1977.
to upgrade the containment systems The program objective was of each BFN unit for suppression pool hydrodynamic loads I
which were not explicitly included in the original design specification.
1 Owners Group and NRC activities resulted in Mark generic load definitions and corresponding structural acceptance criteria to be applied for each domestic Mark I for the plant.
NRC's generic safety evaluation report I
Mark I containment system long-term program, NUREG 0661 (Reference 1), was published in July 1980.
for completion of DFN LTP modifications The current orders were issued on January 19, 1982.
Those orders require installation of all modifications necessary for compilance I
with NUREG 0661 before the start of Cycle 6 operations of each DFN unit, 11.2 Browns Ferry Design Criteria design criteria (Section 4.0) defined The DFN LTP general the basis for structural analysis of DFN containment system I
components as well as structural design of required It also ensured compliance with the intent modifications.
of NUREG 0661.
The detailed design criteria for analysis of torus attached piping systems (Appendix A) supplemented the general design criteria and defined specific requirements and procedures for analysis of torus attached piping systems.
Structural Analyses and Design of Required 11.3 Modtfleattons Structural analysis and design of required modifications for each basic category of DFN containment system components in Sections 5.0 through 9.0.
were performed as described I
The analyses addressed containment systems as configured for including the start of Cycle 6 operations of each unit, installed for NUREG 0661 comp plant modifications Modifications related to suppression I
local and bulk temperature requirements in NUREG 0661 for other reasons.
pool All were designed as described in Section 10.0.
modification designs comp!!cd with the Drowns Ferry LTP I
design criteria and NUREG 0661.
I PUAR.II 11-1 I
I I
BFN-PUAR Analysis and design of associated modifications to the 10-inch S/RV discharge line vacuum breakers and the drywell/wetwell vacuum breakers have also been completed.
Permanent documentation of all analysis and design activities associated with the BFN LTP was accomplished, in compliance with Section 4.6.
Future modifications to BFN containment system components will be designed in accordance with the BFN LTP design g
criteria, when the modi fications are wi thin the region of g
in fluence o f torus hydrodynamic loads.
11.4 S/RV Confirmatory Test An in-plant S/RV confirmatory test was successfully completed in BFN Unit 2 during April 1983 in accordance 3
with Section 4.6.
The test results were documented by 3
Reference 41.
Correlation of analysis and test results, including definition of selected load reduction factors, was accomplished as described in Appendix C.
11.5 Installation of Modifications and Final Conclusions At this time (December 1984), all LTP modifications and 10-inch S/RV vacuum breaker modifications have been installed in DFN Units 1 and 3.
Major LTP modifleations and 10-inch S/RV vacuum breaker modifications have been installed in Unit 2.
Unit 2 is in its Cycle 6 refueling outage.
All DFN containment system modifications for compliance with NUREG 0661 will be completed before restart for Cycle 6 operations, in accordance with NRC's orders (Reference 12).
Other modifications will be installed according to a NRC-approved integrated schedule.
DFN LTP total costs are currently estimated at $105,000,000 excluding interest payments and lost power revenues while installing modifications.
The extent and scope of those modifications are sunmarized by Table 11-1 and the construction photographs in Appendix 0.
This PUAR provides an accurate and sufficient sunina ry o f LTP activities for all three BFN units.
Il-2R2 PUAR ll
TABLE A-4 (CONTINUED), SHEET 8 ANALYSIS CRITERIA FOR TORUS ATTACHED PIPING-LOCA EFFECTS (1, 2, 5)
S gQTCmyDTm LgT g
p]
s rat 4TIONS AND STpFM LIMITS l
158.(CONTINUED) l POSTPROCESSOR 8 PUAAS COIG. 16 i
O.0 POOL SELL PRIMARY y)4 2
e 2.4 S ESSENTIAL (PRESSt#tE +
P- ~ d
+ 0.751 (M + M g
SUSTAIED + DBA )
Mg = M (DM + PL)
(D,2_g 3 Z
& MSS.
2 Y " N EIh *phfID ACT. COMP.
21.2 Sh FI MONESS.-SVC LEVEL D
+ PSOL + PSFB Poem i 2.0 P N0E SS.
ESS. - SVC LEVEL 8
+ DNJOL) g p
SEC0peARY 2
+ 0.7EI A+
ISg+Sh M
PJ
=M(k+)PS2 (D,2_g 3 (PRESSt#E + SUSTAIED MC 2
I
+
+ EXPANSION + (BA)
OR 1
= 1,2 iMc ISg Z
DR MC = M[TI)
AND M
= M(PS2 + PG23 1MD 1 3Sc D
Z
I TABLE A-4 (CONTINUED), SHEET 9 I
NOTES:
- 1. THESE EQUATIONS REPRESENT THE NORST CASES FROM PUAQG a
TABLE 5-2 AND MARK I CONTAINMENT PROGRAM LOAD DEFINITION g
REPORT NEDO-21898,SECTION 3, TABLE 3.0-3, FIGURES 3.0-1,
-2,
-3,
-4, AND -5.
- 2. ALL DYNAMIC ANCHOR POINT MOVEMNTS ARE INCLUDED IN EQUATIONS S, 10, AND 11. FATIGUE ANALYSIS REQUIREMENTS WILL BE SATISFIED BY DEMONSTRATION OF COMPLIANCE MITH ASME CODE SECTION III NC 3600, EQUATIONS S, 10, ANO 11.
- 3. THE PUAAG (MFEMNCE 13) PERNITS THE PIPING STRESS l
E ALLONAOLES, PIPING DAWING VALUES, ADW PIPING SUPPORT 5
ALLONASLES TO MEET THE REQUIREMNTS OF SERVICE LEVEL C.
THIS DOES NOT APPLY TO ACTIVE CONPONENTS.
- 4. -THE PUAAS (REFERENM 13) PE MITS TM PIPING STMSS l
ALLONAOLES, PIPING DRM ING VALUES, AND PIPING SUPPORT ALLOMABLES TO MEET THE REQUIREMENTS OF SERVICE LEVEL D.
THIS DOES NOT APPLY TO ACTIVE COMPONENTS.
- 5. A STRESS RANGE EVALUATION NWST BE PERFORNED FOR ALL THERNAL CYCLIC CONDITIDIS Ape ALL DYNANIC DISPLACENENT CYCLIC CONDITIONS THAT ARE QUALIFIED BY CODE EQUATIONS 10 OR 11.
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t v
t BFN-PUAR APPENDIX H FINAL TVA RESPONSES TO NRC AND BROOEHAVEN NATIONAL LABORATORY QUESTIONS I '
l I l
i l
l I
l l-
General Response to PUAR Questions BFN LTP analysis and design activity has proceeded on I
a schedule necessary to support installation of all modi fications during the Cycle 4 and 5 refueling outages of each unit, as required by NRC.
The first BFN Cycle 4 refueling outage began in April 1981, and most of the major modi fication designs were complete by May 1981.
Remaining modi fication designs, primarily for torus attached piping external supports, were complete in time to support l
I installation during the Cycle 5 refueling outages.
In order to satis fy schedule commitments, it was necessary i
to make interpretations of LDR and NUREG 0661 requirements I
based upon the best available inforna tion a t the time of analysis.
Most of the interpretations were originally established in 1979 and early 1980.
A continuing ef fort
. I to~ remove excessive conservatism from load definitions and analysis methods was made, particularly when that conservatism would result in unnecessary, impractical mod i fi ca t ions.
When la ter information on load definitions and associated analysis methods became available, it was compared to the I
previous interpretations.
The later information was used for reanalysis and associated design work i f a signi ficant unconservatism in the previous interpretation was indicated.
For example, the final downcomer t iebar/V-bracing modi ficat ion I
resulted from November 1981 changes in the DBA condensation oscillation lateral load definition.
Sometimes, later information was used to remove excessive conservatism in remaining analysis and design work.
For example, the 1.1 SRSS load combination technique was I.
permitted for torus attached piping analysis after NRC's final position on this subject was defined in April 1983 by PUAR Reference 58.
An absolute summation combination technique was required prior to that time.
Finally, when the later information showed the previous load definitions and analysis methods to be adequately (but not excessively) conservative, the original interpretations were I
retained.
In these situations, reanalysis utilizing the later information would have been unnecessary and costly, and, in some cases, would have resulted in delays in the I
installa tion of modi fica t ions.
Many of the PUAR questions derive from situations where the original interpretations stated in PUAR Section 4.were used for analysis.
Justification for these interpretations was
-provided in PUAR Section 4, Section 5, and Appendix C.
Additional technical justification follows in the responses I
to specific questions on these topics.
Other PUAR questions simply request additional in forma t ion, which is provided in the responses.
H-GR-lR2 PUAR.00
It is TVA's position that the BFN PUAR and our review question l
responses demonstrate compliance with the intent of the Mark I Containment Long-Term Program and NUREG 0661 (i.e.,
to upgrade the containment sys tem sa fety margins, for all postulated hydrodynamic loading conditions, to those intended by the original design specifications).
On this basis, we feel that all indicated safety concerns are fully and satisfactorily addressed, and we respectfully request a favorable final evaluation for the BFN LTP.
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I II-GR-2R2 PUAR.00 v
ITEM 1:
According to Section 4.2.5 of the PUAR, BFN used the loads defined by the PULD and the LDR Section 4.3.2 for pressure loads on the torus.
Ilowever, BFN applied a much smaller margin on the LDR load than s t ipula ted in NUREG 0661, page I.
A-6.
2 The BFN margin of 6.5 percer.t on the LDR upload i s justified I
in the BFN PUAR on the basis that the 15 percent margin recommended on page 39 of NUREG 0661 is unnecessary because the EPRI 1/12-scale model had the BFN geometry, and that I
the Acceptance Criteria (AC) margin of 21.5 percent should therefore be reduced by 15 percent to yield 6.5 percent.
This does not meet the intent of the AC.
The 15 percent margin of NUREG 0661 was imposed for several reasons (see I
pp. 36-38 of NUREG 0661), the geometry being only one of the concerns.
Consequently, a full justification for the reduction of the margin from 21.5 percent to 6.5 percent is needed, or the ability of the torus to withstand a 15 percent load increase must be demonstrated.
RESPONSE
The uncertainty margins used for BFN pool swell load definition and the BFN pool swell analysis procedure ensured conservative structural response predictions.
Some justifi-cation for this fact is given in Section 4.2.5.2 of the PUAR.
Additional justification follows:
1.
Uncertainties regarding the 2D/3D test model results were minimized because the 1/4 scale 2D and 1/12 scale 3D models for the generic Mark I LTP tests were prototypical of BFN geometry.
2.
Signi fican t conservatism was added to the BFN pool swell load definition because fluid compressibili ty I
effects in the vent system were not considered in the-1/4 scale plant unique tests.
This conservatism is recognized and quantified in Section 2.4 of Supplement I to NUREG 0661.
3.
BFN plant unique 1/4 scale tests for nornal operating conditions were conducted at minimum AP and maximum I
downcomer submergence, thereby ensuring upper-bound pool swell load predictions.
4.
BFN plant unique 1/4 scale tests for 0.0 AP conditions I
were conducted at 0.0 AP and maximum downcomer submergence, thus ensuring upper-bound pool swell load predictions.
I II-DNL I-IR2 PUAR.00 I
'2--
t-
--w,,-y.-,-.-g,,,~,,_y
r 5.
The BFN vent system and torus analysis procedures fo r pool swell loading included signi ficant conservatisms.
Sections 4.4.5 and 6.11 of the PUAR sumnarize the BFN vent system analysis procedure.
Sections 4.4.4 and 5.4.2.7 of the PUAR sumnerize the BFN torus analysis procedure.
Signi ficant analyt ical conservatisms included the two percent damping assumption for vent system analysis, the 80 percent wa ter mass assumpt ion for torus analysis, and the two percent torus damping assumption for all opera ting AP pool swell load combinations.
6.
The BFN uncertainty margin for pool swell loads (6.5 percent) was conservatively applied to the predicted torus response including vent system input effects, g
whereas the download and upload margins in NRC's g
acceptance criteria (Appendix A of NUREG 0661) are applicable for torus hydrodynamic pressure loads only.
7.
The BFN uncertainty margin (6.5 percent) exceeds NRC's recommended download margin (5.4 percent).
It also exceeds one standard deviation of the BFN 1/4 scale results for operating AP conditions.
Those standard devia tions were approxima tely 3.6 percent and 4.0 percent for upload and download respectively.
8.
BFN operating AP pool swell dynamic responses were conservatively combined with dynamic responses from other load sources by the methods described in Section 4.4.2 of the PUAR.
Additional assurance regarding any remaining upload concern is provided by the fact that the BFN torus tiedown design is 3
not controlled by a pool swell load combination.
This would E
remain true even i f an addi t ional 15 percent upload margin were added for pool swell loads.
I I
I H-BNL I-2R2 PUAR.00
I ITEM 2:
What margin was applied on the LDR download?
Is the download speci fication consistent with Section 2.3 of the the Acceptance Criteria?
RESPONSE
A 6.5 percent uncertainty margin was applied for both down-I load and upload as described in the response to BNL item 1.
The specification in Section 2.3 of NRC's acceptance criteria requires a download margin of 5.4 percent, based I
upon a peak download of 2700 po:inds for BFN 1/4 scale opera t ing AP tests.
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I II-BNL 2-lR2 PUAR.00
I ITEM 3:
For what structures would the load exceed acceptable levels if the torus pressure loads were made consistent with NUREG 0661?
By how much, and for what load combinations?
RESPONSE
To make the BFN torus pressure loads consistent with NUREG 0661 an additional 15 percent margin would be added to the g
upload phase--if the other conservatisms in the BFN pool E
swell load definition were disregarded.
The download margin would be reduced by I percent.
Assuming that the same conservative analytical procedure was applied, maximum stresses in the download phase would decrease slightly and maximum stresses in the upload phase would increase by less than 15 percent.
(An increase of 5 to 10 percent g
is estimated.)
This level of potential stress increase E
could readily be compensated by removal of some of the conservatism in the analytical procedure and load combination technique.
Therefore, realistically, there is no potential to overstress a BFN structure by changing the torus pressure load definition to comply with NUREG 0661 generic margins.
~I I
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H-BNL 3-IR2 PUAR.00
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ITEM 4:
Was the vent header impact load defini t ion of pages 6-17 of I
the PUAR in accordance with Section 2.10.1 of NUREG 0661?
If not, explain the dif ferences and provide estima tes showing that sufficient margin exists to acconmoda te the I
NUREG load.
RESPONSE
The vent header impact load definition was not in accordance with Section 2.10.1 of the NUREG 0661 acceptance criteria.
I Section 2.10.1 addresses loads on a vent header deflector.
DFN does not have a vent header deflector; however, the BFN vent headers were reinforced near the center of each non-vent bay as a result of pool swell impact loading I
analysis as described in Section 6.11.3 of the PUAR.
A typical BFN header reinforcement installation is shown by PUAR Plates 7 and 8.
The BFN vent system pool swell impact load analysis (PUAR Reference 17) and header reinforcement modification design were per formed in 1979, prior to the release of NUREG 0661.
I The longitudinal velocity and impact timing profiles were based upon EPRI 1/12 scale split orifice tests for operating AP and 0.0 AP pool swelI conditions.
NUREG 0661 speci fied the use of a single " conservative" profile for impact velocity and timing for alI conditions.
Ilowever, a comparison of the resulting peak impact pressures
.I on the BFN vent header showed that the existing analytical values were more conservative for the entire non-vent bay.
This was particularly true in the critical region where the I
DFN reinforcement modification is located.
Therefore, within the non-vent bay it was concluded that the existing analysis results were conservative relative to the revised load definition from NUREG 0661.
WIthin the vent bay the peak impact pressures would be somewhat higher with the revised load definition.
- However, I
conservative estimates of the increased vent system stresses in this region showed all stresses to be less than 24 percent of allowables for the operating AP case and 37 percent for the 0.0 AP case.
Further consideration of this information leads to the basic conclusion that the 1979 analysis was appropria tely I
conservative and sufficiently accurate to address all structural concerns of the BFN vent sys tem for pool swell impact and drag loads.
Additional analysis was not necessary.
I
"- = 4->"2
" " ^ ^ "
I
I I
ITEM 5:
Were the LOCA Jet and bubble drag loads for BFN evaluated in accordance with the LDR and NUREG 0661?
I
RESPONSE
Yes, LOCA jet and bubble drag loads for BFN were evaluated in accordance with the LDR and NUREG 0661 (See PUAR, Appendix D, Sections D.I.l.2 and D.I.l.1, respectively, for discussions).
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H-BNL 5-IR2 PUAR.00 l
I'
I ITEM 6:
For analyzing structures affected by CO loads, the LDR and I
NUREG 0661 prescribe absolute sunma tion of the CO load harmonics at 1-Hz intervals from 1 to 50 Hz.
BFN used an alternate approach where:
(1) forcing frequencies above 31 Hz were neglected, and (ii) four particular load harmonics (the ones at 4-5, 5-6, I
10-11, and 15-16 Hz) were added absolutely and added l
to the SRSS of the remaining 26.
l Justi fy the neglect of forcing f requencies above 31 Hz for (a) torus shell loads, and (b) submerged structure drag loads.
(Arguments about small torus response do not apply for drag loads.)
Why were CO drag loads (page 4-4) analyzed fo r 1-31 Hz only, but post-chug drag loads (page 4-5) for 1-50 Hz?
RESPONSE
When the DBA CO load definitions were provided by the LDR, it soon became apparent that there were significant inherent conservatisms, not the least of which was the lack of any I
information about the phasing relationships between the Fourier harmonics.
Clearly, conservatisms could have been maximized by applying all 50 CO harmonics using an absolute I
summation rule, and while some might infer this approach from the LDR and NUREG 0661, it was not prescribed.
Va r ious experts ident ified speci fic conservatisms and reconmended approaches that would allow more realistic accounting for I
the potential DBA CO event.
Some of the key findings by these experts fo' low:
From PUAR Reference 19 (or equivalently:
GE/NEDE-24840),
Section 3.4:
(1)
The 5.5 Hz harmonic was the dominant content of the I
loading.
... all harmonics appear to be randomly phased (2)
I rela tive to the dominant harmonic at 5.5 Hz."
(3) investigators have seen a tendency for a fixed phase relationship between the dominant harmonic and I
one.or two multiples of the dominant (e.g., 5.5, 11, and 16.5 Hz) from examination of data from individual I
"-""' a-'"
""^^ aa I
I
pressure transducer records for the FSTF test, but g
showed random phasing for all other harmonics."
g (4)
While it was admitted that there appeared to be some rela tive periodici ty of the harmonic ampli tudes tha t would preclude the appropria teness of pure SRSS combination of the harmonic amplitudes, it was stated that "... It is highly improbable for more than about g
three harmonics to be worst-case phased at any one g
~
time..."
... a rule which requires about three harmonics to be l
(5) absolute combined with all additional harmonics SRSS a
combined is consistent with the assumption of steady-state periodic amplitudes and random phasing."
(6)
The LDR ampli tudes are defined wi th signi ficant conservatisms as can be seen from Figure 3-11 (especially in the frequency range from 40 to 50 Hz where most LDR amplitudes are much more than 100 percent greater than the average FSTF amplitudes).
(7)
Also from Figure 3-11, it can be seen that actual FSTF ampli tudes seem to show that CO has rela tively li t tle frequency content above 30 Hz.
From PUAR_ Reference 42, Section 4:
(8)
Conclusion No. 2 states that:
"For structures with frequency content similar to the FSTF or Oyster Creek torus and supports, only the harmonic responses below 30 Hz need to be computed and included."
Based on the findings listed and our own best technical judgment, TVA feels that neglect of the forcing frequencies above 30 Hz is justified for:
(a) torus shell loads (b) submerged structure drag loads An explanation of our reasoning follows:
As stated in finding (7), there is little frequency content of the CO loading above 30 Hz; this is the main reason for neglecting it.
Additionally, for the BFN torus (even after substantial modifications that resulted directly from CO loads analysis), the responsive structural modes occur at frequeneles below 30 Hz.
Shapes of the higher frequency modes of the torus will not participate significantly with the shape of the CO pressure distribution.
This argument I
H-BNL 6-2R2 PUAR.00 I
y I
I for the post-chug loading on the torus since it the same distribution shape as CO with only the harmonic also applies has ampli tude coef ficients being di f ferent.
frequencies more than For harmonic forcing components with 1.5 times the natural structural frequencies, the dynamic I
load factors (DLFs) become less than 1.0.
For harmonic than twice the natural forces with frequencies greaterthe DLPs are small and structural frequencies, In this range, the forcing asymptotically approach zero. components would have negligible effect on the s I
for high frequency This is the case with the torus harmonics.
I evidence of the adequacy of the BFN Further, empirical the DDA CO loading on the torus for analytical approach (see Table is provided in the responses to BNL item 7 is provided for the chugging I
DNL-7-1).
Similar evidence in the response to FRC item 7 (see loading on the torus Table FRC-7-1).
I Submerged structures are a di f ferent ma t ter, however.
submerged structures were primarily frequency ranges and were dominated
~
Initially, most In order to avoid highly amplified responsive in the lower I
by the DBA CO harmonics.
virtually every submerged structure required substantialresp These modi fication to stif fen and strenghten it. modifications took some o I
While it was responsive range with post-chug fluid drag.to stif fen submerged structures (
above the post-impracticalinternal portions of large piping systems) chug forcing frequencies and still maintain viable designs
=
for thermal loads, it was possible, after many iterations, strength to meet g
to obtain designs that had sufficient This left BFN with allowables for all load combinations.
g stiff submerged structures that were well within a range of full 0 to 50 necessitating use of the
- thus, in the LDR.
post-chug drag,liz range of the post-chug definition prescribed 20, and 21 I
PUAR Plates 11, 18, 19, stiffened submerged structures.
I I
1 I
PUAR.00 Il-BNL 6-3R2 A
uws J
E I
I ITEM 7:
The approach of GE/NEDE 24840 - which is i tsel f a departure i
from the LDR - calls for taking the sum of the fou r harmonics which produce the highest structural response, and adding them to the SRSS of the remaining harmonics.
Were the forcing functions a t 4-5, S-6, 10-11, and 15-16 Hz the ones which produced the highest structural response for bo th torus shell and all drag loads?
In the work done by SMA l
(References 19 and 42 of the BFN PUAR), the absolute sunina t ion o f t he four highest harmonics had nothing to do with phase relationships, but was an artifice used to arrive B
a t an 84 pereen t nonexceedance probabi1i ty (NEP).
Based on the discussion in the PUAR, BFN's procedure does not E
guarantee an NEP of 84 percent.
Jus t i fy BFN's depar ture f rom the recorrmended procedure and/or demons t ra te s t ructu ral margins which would adequately cover increases in the CO loads.
Was Alternate 4 of the CO baseline rigid wall pressure spectrum applied to BFN?
RESPONSE
I While the approach recorrmended in PUAR Reference 19 (or GE/ NEDE-248 40 ) is a departure from the LDR, TVA belleves it is well justified by the thorough studies and findings of many experts.
Those findings clearly stress the extreme conservatism, hence inappropria teness, of absolu te sunma t ion of all response harmonics.
As the approach is speelfied, it requires the identification of the three or four highest response harmonics of a structure to be combined absolutely with the SRSS of those remaining.
This identification, however, is an impracticable task when one considers the number of structures to be analyzed and the response quantities of interest for each (e.g., displacement, acceleration, force, stress, stress intensities).
Also, while this approach may guarantee 50 or 84 percent NEP for the CO loading alone, there can be no such claim for the controlling design load combinations involving CO since the points of maximum responses for CO load combina tions are likely to be di f ferent t han fo r CO alone.
Therefore, TVA chose to vary slightly from the approach suggested in Sec t ion 6 o f PUAR Re f e renc e 19.
The approach used was justifiable and practical for timely, cost ef ficient implementation.
-I t cannot be guaranteed for both the torus shell and all submerged structures that forcing funct ions a t 4-5, 5-6, 10-11, and 15-16 Hz were the ones producing the highest structural responses, although for some they may be--
especially for the torus shell since its primary response I
H-BNL 7-lR2 PUAR.00 g
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occurs around these frequencies.
Ilowever, it is inportant to I
note that the suggestion to use the four highest response harmonics was an artifice to obtain an 84 percent NEP of an artificial load definition--one which Dr. Alan Bilanin has stated is 33 percent conservative just due to the presence of r
I the bulkheads on the FSTF (see Reference BNL-7.1, Section 1, Equation 1.7).
There are the additional conservatisms of the LDR prescribed amplitudes, as already addressed by finding 6 listed in our response to BNL Item 6.
And, specifically concerning torus I
shell responses, TVA has the conservatism of having applied a maximum envelope of the three LDR al terna t ives speci fied fo r harmonics between 4-16 liz rather than selecting the one alternative producing maximum response.
To pursue the issue of why TVA chose forcing funct ions a t 4-5, 5-6, 10-11, and 15-16 liz, the following arguments are of fered:
(1)
As contended, it is a practical impossibility to identify the three or four highest CO response harmonics I
for all structures and response quantities.
Th e r e fo r e,
TVA sough t a prac t ical al terna t ive tha t would nain tain some conservatism over a pure SRSS combination of the CO harmonics.
Because of the reported indications that I
there may be some fixed phase rela t ionship between the dominant harmonic at 5.5 liz and its first few multiples, we decided to use the 5-6, 10-11, and 15-16 Hz LDR I
harmonics.
The 4-5 liz harmonic was added to this list since it was the next largest amplitude harmonic in the LDR definition.
It should be noted that three of these (4-5, 5-6, and 10-11 Hz) are the highest of all the I
anplitudes provided in the LDR.
(2)
Conpa r ing the LDR prescribed DBA CO and post-chug I.
amplitudes, it can be seen that for structures having prinary response modes below about 20 liz, CO load combinations would be expected to control for design since the CO ampli tudes below 20 IIz are generally larger I
than the post-chug anplitudes.
Structures having prinary response modes above 20 IIz would most likely be controlled for design by post-chug load combina tions
.I since post-chug amplitudes are higher in this range.
Therefore, by picking three of the highest CO amplitudes for absolu te summa t ion, the three potentially most damaging load components are assured of receiving conservative combination in the response predictions for structures likely to be controlled in their design by CO load combinations.
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II-DNL 7-2R2 PUAR.00
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(3)
PUAR Reference 19 recommends a procedure for obtaining 50 percent or 84 percent NEPs and shows that this is g
achieved when the three or four highest response g
harmonics are combined absolutely.
This is based on comparisons with predicted response values a t these probabilities as taken f rom cons t ruc ted CDF cu rves fo r l
both the FSTF and Oyster Creek torus.
As can be seen E
from the CDF curves of Figures 4-6 through 4-10 from Reference 19, all have rela t ively small variance as 3
indicated by their steep slope, it can also be seen 5
that the-response values predicted by total absolute sum of all harmonics is well above the response value associated with 100 percent NEP.
Therefore, while even a total SRSS combina t ion of all harmonics would be only slightly unconserva t ive rela t ive to the CDP 50 percent and 84 percent NEP response values, a total absolute combination would be grossly overconservative.
(4)
While there is no reason to suspect that a total SRSS comb i na t i on o f a l l harmonic resporises would produce a response value as low even as that associa ted with a O percent NEP, it is interesting to observe from PUAR Reference 19, Figures 4-6, 4-7, 4-8, 4-9, and 4-10, g
that the percentage differences between the 0 and g
50 percent NEP response values are 11.3, 15.6, 16.3, 24.0, and 18.2 percent, respectively.
Therefore, even a pure SRSS combination rule would result in at leas t a O percent NEP response value, meaning the potent ial unconservatism could be no more than the above percentage differences between 0 and 50 percent NEP g
values.
Further, since TVA's approach is more E
conservative than pure SRSS, our predicted response values would be still less of a percentage di f ference.
All of these arguments mean that any slight uncon-servatism there may be in TVA's approach is much more than offset by the inherent conservatism in the FSTF based load definitions.
Inherent in those definitions, g
as already mentioned, is at least a 33 percent I
conservatism according to Dr. Alan Bilanin.
So, while our approach does not rigorously assure 84 percen t NEP of t he conserva t ive LDR load de fini t ions per se, we feel very confiden t that this level or more would be achieved if more realistic load definition and g
analysis techniques were possible.
A strong indication g
of the conservatism of the BFN DBA CO analysis is seen in the attached Table BNL-7-1.
BFN stresses and reaction forces are presen ted, factored as nearly as possible to an FSTF-equivalent basis, and compared to measured and calcula ted NEP values fo r the FSTF.
I II-DNL 7-3R2 PUAR.00 I
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I Looking at Table BNL-7-1, the factored BFN cradle support pad reaction forces are seen to be conservative I
with respect to the 84 percent NEP values (this would indicate similar conservatism of the stresses in the critical cradle regions which required extensive I'
modifications).
Also, the BFN BDC membrane stress intensity (factored to an FSTF-equivalent basis) is almost exactly the same as the 84 percent NEP value for the FSTF.
While this degree of closeness may be I
-coincidental, it does provide additional evidence that there is no large deficiency in stress intensity predictions in the BFN analysis.
Finally, there are two additlonal conservatIsms worth noting about the BFN analytical approach.
First, 2 percent damping was used for the DBA plus S/RV load
~I.
combination analyses of the torus and submerged structures even though higher damping is justifiable because of the higher service level allowables.
I Second, conservative load combination techniques were applied (see PUAR Section 4.4.2).
Concerning the final question about Alternate 4 of the I
CO baseline rigid wall pressure spectrum, we do not know to what this refers.
Addi t lonal Re ference:
BNL-7.1 Structural Mechanics Associates, "A Statistical Basis for Load Factors Appropria te fo r Us e wi th CO I
Harmonic Response Combination Design Rules," Report No. SMA 12101.04-R003D, March 1982.
I Addendum In the September 5, 1984 meeting, BNL expressed remaining concerns over TVA's use of the absolute sum of - the four highest DBA CO harmonic amplitudes, rather than the four harmonics causing the greatest response.
BNL also asked:
I-
"How much greater could the DBA CO load be without exceeding allowables"?
(paraphrased)
I Our response considers the torus separately from the tiedown system for reasons which are explained below.
The most highly stressed regions of the torus, rela tive to I
allowables, are in the cradle adjacent to the scab plates
. described in PUAR Section 5.2.4.3.
(Also see the response to FRC Item 11.) The controlling load combination is number 14 which does not include DBA CO.
There are large margins II-BNL 7-4R2 PUAR.00 I
I I
for al1 load combinations involving DBA CO.
If one assumes 3
that cradle stresses are directly proportional to the net g
compressive loads, the DBA CO reactions could be 2.5 times the values computed by the TVA analysis without exceeding allowables.
The tiedown system, on the other hand, is controlled by load combination number 27 which includes DBA CO.
Tiedown g
stresses are directly proportional to net uplift loads.
g From computer calculated responses to the 0 to 30 Hz unit amplitude harmonics and hand calculations, the increase in reactions by taking the four highest responses, rather than the responses to the four highest amplitudes, has been quantified.
The conservatism of TVA's method of enveloping the three al ternate sets of ampli tudes has also been quan-g tified.
These calculations show that the DBA CO reactions E
presented in Table BNL-7-1 would be 9 percent higher i f the four highest harmonic responses had been used, while retaining the conservatism of enveloping the alternate amplitudes.
If the individual alternate amplitude sets are used, the reactions would be 5.4 percent higher than those in the table.
With the 5.4 percent increase, the tiedown system stresses do not exceed allowables.
It is important to reemphasize that the SMA method of Reference 19 provides large margins for support reactions.
It is also noteworthy that the limiting load combination (number 27) includes the highly unlikely simultaneous occurrence of the maximum responses g
due to the safe shutdown earthquake, DBA CO, and a single g
valve S/RV actuation.
TVA combined these three dynamic events absolutely.
If, for example, a 1.1 SRSS combination of the three dynamic loads had been used, the uplift loads would have been barely great enough to overcome the deadweight.
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H-BNL 7-5R2 PUAR.00
TABLE BNL-7-1 COMPARISON OF FSTF AND BROWNS FERRY DBA'C.O. RESPONSES BFN
' BFN RESPONSES NEASURED 50% NEP 94% NEP FACTORED TO FSTF, PER-FSTF, PER FSTF, PER CALCULATED EQUIVALENT REF.I9 REF.19 REF.19 RESPONSE QUANTITY RESULTS FSTF TABLE 7-2 TABLE 6-2 TASLE 6-2 NENORAME S
INTENSITY 1.98 2.77 m 2.6 2.47 2.7S (KSI)
INSIDE REACTION 306 2022 S3-122 140 (KIPS)
SUTSIDE IEACTION 333 2208 110 140 159 (KIPS)
"I 4 #BFN (1) FSTF EQUIVALENT SHELL S.I. = E gpy PLANT UNIQUE PRESOURE FACTOR =0.85 NNERE, R = NINOR RADIUS OF THE TONUS, T = SELL THICKNESS, AND S = STRESS INTENSITY (2) FSTF EQUIVALENT SUPPORT REACTIONS = (9FN REACTION)
POOL A PE CRA L PLANT UNIQUE PRESSURE FACTOR,
~
s
~ ~ ~ ' - ~ -
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ITEM 8:
Were pre-chug loads applied to BFN according to the LDR and NUREG 0661 speci fica t ions regarding ampli tude, ci rcumferent ial and vertical dis tribut ion and cycle duration?
If not, provide quantitative justification.
RESPONSE
The pre-chug loads were applied in complete accordance with the LDR and NUREG 0661, including considerations of amplitude, circumferential and vert ical dis t ribut ion, and cycle duration.
With regard to the latter, the responses o f 71 shell modes due to six independent harmonic (implying 3
infinite duration) forcing functions, fine-tuned to the E
structural frequencies, were enveloped for all points in the model.
No credit was taken for the finite duration of the actual event.
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H-BNL 8-lR2 PUAR.00 I
I ITEM 9:
For post-chug loads, were the harmonic forcing funct ions used in the 1-30 I!z range the ones specifled in the LDR, and were they applied in the manner prescribed in the LDR?
If not, jus t i fy departures.
RESPONSE
For both the post-chug pressure loads applied to the torus and the drag loads applied to internal structures, the harmonic forcing functions used in the 1-30 IIz range were I
those specified in the LDR, and they were applied in the manner prescribed by the LDR.
Appendix D, Section D.I.2.4.2 of the BFN PUAR explains how the LDR prescribed method was I
applied for submerged structures by considering the closest downcomer load sources together with worst-case phasing between the sources.
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0 ITEM l_:
The finite element model of Figure 6-7 shows amputa ted downcomers.
How were the CO and CH loads applied to these amputated downcomers?
RESPONSE
The finite element model in Figure 6-7 was used to examine more closely the vent downcomer/ header intersection.
Loads for the di f ferent combination events which included g
condensation oscillation and chugging were extracted from E
the 450 vent system beam model (PUAR Figure 6-2) at the node representing the downcomer/ vent header shell intersection.
These loads were then input into the truncated model at the end of-the downcomer.
(The downcomer end is comprised of rigid beams connected by a node in the center.) The appro-priate stresses were then extracted and compared to the stress allowables.
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I H-BNL 10-IR2 PUAR.00 L__
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ITEM 11:
Were the CO loads applied to the downcomers in accordance with the LDR and NUREG 0661?
Were the eight load cases of
'ection 4.4.3.2 of the LDR analyzed for all relevant vent system parts (including main vent / vent header intersection, I
drywell/ main vent interaction, downcomer/ vent header intersection, etc.)?
The PUAR explicitly mentions considering different load cases only for the downcomer/
tiebar intersection, and in that case refers only to four load cases rather than the eight of the LDR (Section 6.7.1.2.1).
Why is 2.5 percent damping justified for DFN for CO lateral load analysis?
Note that Table 6-10 shows no margin for the downcome r/ven t header intersection in Load Combination 27 which involves CO.
RESPONSE
The CO downcomer loads were applied in a manner consistent with the intent of the LDR and NUREG 0661.
The load from the differential pressure for one downcomer was added to the internal pressure, that occurs simultaneously in all downcomers, thereby producing a higher load in one downcomer in each pair.
Thus, f rom Figure 4.4.3.4 of the LDR, a darkened downcomer indicated that the di f ferential and I
internal pressures were working together simultaneously, whereas the other downcomer in the pair experienced only the internal pressure.
Based on the primary downcomer swing f requency extracted from a modal analysis of the system, sinusoidal forcing functions in were applied to downcomer pairs defined by Figure 4.4.3-3 I
the LDR.
Since the primary swing mode for the BFN system occurs at approximately 8 Hz, the 1st, 2nd, and 3rd harmonics were applied in the 4, 8, and 12 Hz ranges, respect iv ely.
I Another aspect of the BFN analysis was the application of the first harmonic forces to the coincident 8-IIz swing frequency.
Response of the system to this single frequency I
load envelops the sum of the three harmonics defined by the LDR.
This load was subsequently applied in the stress evaluation.
Also, the first harmonic force amplitudes were I
applied with 16 and 24 IIz sinusoidal functions to verify that higher frequency responses do not impact the total CO
- response, in actuality, 30 individual sinusoidal fu nc t ion s were applied for each load case to account for potential I
response at the 1/2, 1, 1-1/2, 2, and 2-1/2 harmonics of the six discreet primary swing mode frequencies in the 8-9 IIz range.
Note that this load was in addition to the vent I
system CO loads which were applied in a separate analysis.
II-BNL 11-lR2 PUAR.00 I
Four load cases were ini tially analyzed for downcomer CO lateral loads as indicated by PUAR Figure 6-12.
From inspection of the first four load cases and resulting stresses, it was evident that in each instance, the worst effect on a downcomer would occur on one that was located on the inboard side of the vent header.
Furthermore, the highest loaded downcomer (unreinforced shell) resulted from the application of Load Case 1 (di f ferential pressure applied to all inboard downcomers).
Since the second four load cases defined in revision 2 of the LDR are mirror images of the first four cases and greater response resulted from Case 1, it was resolved that the worst loading had already been analyzed.
Therefore, no further analysis was perforned.
All critical locations of the vent system were evaluated fo r DBA CO la teral load combinations.
As noted in item 11, the s tress margin rela tive to Service Level B allowables for the CO combination in Table 6-10 of the PUAR is close to 1.0 for
'the prinary plus secondary stress category.
This stress level should be evaluated with consideration of the conservatism in the load combination (event 21 is a Service Level C combination) and the fact that downcomer lateral Load Case 1 is the worst of the eight potential load configurations.
Preliminary analysis of the BFN containment vent sys tem for the DBA CO la teral load definition Indicated a surface stress level in the vent header shell near the downcomer that approached the yield stress value for SA-516 Grade 70 steel.
Under this situation the total stress level for load combination No. 21 would not meet the allowable for primary plus secondary stress range.
In an effort to avoid additional modification (i.e., downcomer/ vent header reinforcement gussets), the 2 percent recommended damping ratio was investigated as a source of excessive conservatism.
Per Regula tory Guide 1.61, a 3 percent damping value is recommended for analysis of large diameter piping systems for the safe shutdown earthquake.
Fu r t he rne r e, the calculated damping value resulting from the snap pull test of a tied downcomer arrangement (see Figure 4-5 in Reference 5 to supplement 1 o f NUREG 0661 ) wa s found to be approximately 2.2 percent for a 50 percent yield.
Also, the damping versus strain curve indicates a rapid increase in damping above the 50 percent yield level.
Based on these findings, it was concluded that a damping value greater than 2 percent but less than 3 percent is appropriate for the Browns Ferry configuration.
The 2.5 percent value was selected as the midpoint of the 2-3 percent range and utilized for the DBA CO downcomer lateral analysis.
Resulting surface stresses in I'
I
,,.BNL ll.2R, pUAR.00 I
I the vent header shell at the downcomer intersection are 20 3
to 24 ksi ( for DBA CO loading alone) as compared to a yield 3
stress value of 32.6 ksi for SA-516 Grade 70 steel at 4000F.
Therefore, the 2.5 percent damping value is justified based on:
(1)
The structural response of the Browns Ferry downcomer/
vent header configuratlon.
(2)
The projected results of the snap test for higher initial stress levels.
(3)
The damping criteria delineated in Regulatory Guide 1.61.
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II-BNL 11-3R2 PUAR.00 I
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ITEM 12:
Were the chugging loads applied to the downcomers in accordance with the LDR and NUREG 0661?
Were the multivent chugging loads accounted for on all vent system parts in accordance with the LDR and NUREG 0661?
Note that according to Table 6-10, Load Combination 15, which involves CH, has relatively little margin.
RESPONSE
Yes, the chugging loads were applied to the vent system in accordance with the LDR and NUREG 0661.
LOCA chugging loads included post-chug drag, chugging lateral, acoustic vent system pressure oscillation, and gross vent system pressure oscillation, which were applied to the beam models shown in PUAR Figures 6-2 and 6-3.
Responses were determined from those models and local stresses were calculated for the eritical locations.
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H-BNL 12-1R2 PUAR.00
ITEM 13:
What hydrodynamic load definition was used for the vent pipe drain referred to on page 6-10 and shown on Figure 6-8?
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RESPONSE
From a Stardyne model of the vent drain and support system, I
the fundamental na tural frequency of the system was found to be 35.1 Hz.
Clearly, high amplitude harmonics of post-chug l
around this frequency would lead to a strong expectation that a load combination involving post-chug would be I
controlling; although, the possibility of a combination with pool swell was also considered.
From investigations of all potentially controlling design load combinations (including associated service level allowables) it was determined that combination 11 (see Figure 4.3-1 of NUREG 0661) was controlling.
Specifically, the design case determined to be l
controlling was the SBA combination of S/RV plus post-chug 5
fluid drag loads under service level A allowables.
The dynamic loads were very conservatively accounted for by I
the " Equivalent Static Load Method" explained in Appendix D, Section D.I.2.2 of the BFN PUAR report.
All 50 post-chug bubble source amplitudes and FSI acceleration coef ficients l
were summed and used as multipliers of the unit forces in Equations D.I.2-6 and respect 5v)psgdescribed t -
(For) DUB and (F o ely.
To these a resonant DLF = 25 D.I.2-8, (assuming 2 percent damping) was conservatively applied.
I The S/RV load contributions were applied with a harmonic DLF = 1.2 based on the ratio of maximum S/RV bubble frequency-to-system frequency of 14.7/35.1 (again assuming 2 percent damping).
Addendum I
1984 meeting, BNL's consultant, Professor in the Septenber 5, Sonin of MIT, asked if the potential for chugging through the drain pipe and the ef fects of the result ing la teral vent loads had been considered.
TVA responded that the LDR did not include a method fo r I
defining such loads, and the effects could therefore not be evaluated.
Professor Sonin then asked to be provided the properties and dimensions of the drain pipe and its support and the drag loads which had been applied to them.
I H-BNL 13-lR2 PUAR.00 I
A set of five pages of calculations on the ef fects of S/RV and post-chug drag loads were subsequently transmitted to Pro fe ssor Sonin through NRC.
The calculations show the post-chug drag loads, particularly, were defined in an extremely conservative manner which should compensate fo r the lack of a directly applied lateral chugging load.
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H-BNL 13-2R2 PUAR.00 i
I ITEM 14:
- I.
Combining individual S/RV shell pressures by SRSS to obtain multiple valve shell pressures is an exception to the AC.
Justify this procedure for BFN.
RESPONSE
The use of SRSS to obtain multiple valve shell pressures for analysis of S/RV discharges is justified by the BFN plant unique S/RV tests (PUAR Reference 41) and the correla tion of analysis and test results (PUAR Appendix C).
Section C.3.2 of the PUAR specifically addresses this issue.
The measured peak shell pressures during multiple valve tests were approximately 45 percent of the analysis values for single valve tests and 54 percent of the analysis values for multiple valve tests.
Thus the nmitiple valve test pressures are correlated by a 1.2 SRSS of single valve test pressures, but the overall BFN analysis and design approach was clearly conservative.
It is also noteworthy that considerable care was taken in the BFN test to ensure simultaneous actuation of three S/RVs with adjacent discharge locations in the torus.
This represents a " worst location" in the torus.
Referring to I
PUAR Figure 7-3, S/RVs D, E, and M were actuated simultaneously for multiple valve tests.
S/RV E was actuated for single valve tests.
Excellent repeatability was demonstrated for both test series (five single valve tests and four multiple valve tests.)
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II-BNL 14-IR2 FUAR.00 I-
ITEM 15:
Cla r i fy the statement that the torus was analyzed quasi-statically for S/RV hydrodynamic shell pressures.
Wh e r e does g(t),
i.e.,
the wave form of the pressure his tory, in the expression on page 5-13 of the PUAR come f r om?
Are pressures applied statically as stated on page 5-12 or is there a time variation as implied by the expression on page 5-13?
RESPONSE
The wave form of the pressure history, g( t), was genera ted by the QBUBS02 computer code.
The pressures were applied statically to the torus shell to determine torus stresses, deflections, and support loads.
Subsystems, such as attached piping, were analyzed dynamically for the accelera tion response resulting from the assumed shell motion (see page 5-14 of the PUAR).
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II-BNL 15-lR2 PUAR.00
ITEM 16:
Provide the following additional information regarding the in-plant S/RV tests conducted at BFN and the S/RV design loads extrapolated from the tests:
1.0 Description of the tested Quencher Device -
.g 1.1 Drawings showing details of the quencher geometry -
'g plan, elevation, arm length, arm diameter, hole arrangement, spacing, size, etc.
'I 1.2 Location of quencher device relative to suppression pool boundaries and suppression pool surface.
1.3 Any difference between the tested quencher configu-ration and the Monticello version (as described in GE/ NEDE-245 42-P ) highlighted and quantified.
2.0 A description of the loads observed during testing -
2.1 Peak overpressure (POP) and underpressure (PUP) recorded on the torus shell during each relevant S/RV actuation.
2.2 A measure of the frequency content of each pressure signature.
3.0 A description of the test conditions -
3.1 Geometry of the tested SRVDL (diameter, length, free volume, and routing below pool su r fase).
I 3.2 Geome t ry o f any SRVDLs in the plant that dif fer significantly from the tested SRVDL.
13 3.3 S/RV steam flow rate (MS), pool temperature (TPL),
g pipe temperature (TP), water leg length (LW) and pressure differential (AP), if any, for each test.
3.4 Minimum AP permitted by NRC Technical Specification and corresponding LW for all SRVDLs.
4.0 A description of the design conditions for each load I
case used for design -
4.1 Geometry of all SRVDLs involved and their azimuthal location in the torus.
4.2-TP,.TPL, MS, AP, and LW for all SRYDLS involved.
H-DNL 16-1R2 PUAR.00 I
I 5.0 A descript ion of the design loads for each load case -
5.1 Normalized pressure signature.
5.2 Single valve POP / PUP values.
5.3 Spa tIal a t tenua tion of the POP / PUP values (i f this differs from the LDR methodology, sufficient E
additional torus shell pressure data must be a
supplied to jus t ify such devia t ion).
5.4 Frequency range considered.
RESPONSE
1.0 Description of the tested quencher device -
1.1 The plan view of all BFN quenchers in Units 1 and 2 is shown on TVA drawing 47W401-7.
For Unit 3 the l
plan view is shown on drawing 47W401-3.
The tested quenchers were in Unit 2 at azimuths 780-45' (D),
g 1010-15' (E), and 1230-45' (M).
S/RV E was E
actuated for single valve tests and all three (D, E, and M) were. actuated for mult iple valve tests.
These plan views correspond to PUAR Figure 7-3.
BFN quencher arm details are shown on TVA drawing g
All BFN quer!chers are identical in 5
design.
Copies of all referenced drawings are available for review.
1.2 The BFN quencher device locations are shown on TVA E
drawings 47W4 01-3, 47W401-5, and 47W401-7.
Each 5
quencher centerline is at elevation 526.5 which is 5.0 feet above the bottom of the torus shell.
This g
yields a submergence of approxima tely 10.0 feet.
5 Typical BFN quencher installations are shown on PUAR pla te s 10, 11, and 13.
1.3 The BFN quencher device utilizes the previously existing 10-inch ramshead as indicated on drawing 47W401-5, while the Monticello version has a 12-inch ramshead.
The BFN device has a 10-inch x 12-inch reducer between the ramshead and quencher arm while the Monticello version does not.
BFN and Monticello quencher arm designs are identical I
H-BNL I6-2R2 PUAR.00
I except for minor varia tions in support type and location.
The BFN weld cap hole pattern amtches
- I the pattern shown in Figure 1-2 of NEDE-24542-P; i t does not na tch the pa t tern in Figure 1-3 of that report.
2.0 A description of the loads observed during testing -
2.1 Torus shell pressures resulting from the S/RV test
- I are discussed in Appendix C, Paragraph C.S.2 of the PUAR.
Table C-2 of the PUAR presents a comparison of the analytically predicted pressures versus the average of the peak pressures from each test.
This I-information is from the TES Report No. 5172 (PUAR Reference 41).
Pages 1 through 54 of Volume III of the TES report show the pressure traces of the g
torus shell for each test.
A summary of the g
maximum and minimum pressures (POP and PUP) recorded during each test are shown in Tables BNL-16-1 and BNL-16-2.
2.2 The pressure traces for all locations and each test are found in the TES Report (PUAR Reference 41).
I All pressure traces are similar in shape and primary frequency.
The primary frequency of the pressure traces ranges from 5.5 to 6.5 Hz.
Typical pressure traces are shown in Figure BNL-16-1.
3.0 A description of the test conditions -
3.1 & 3.2 The geometry of all SRVDLs is shown on the TVA 47W401 drawing series.
All discharge lines are 10" SCH 40 in the drywell and 10" SCII 80 or 10" SCH 60 in the wetwell.
The routing for all lines below the pool is the same.
The routing in the torus above the pool can be grouped into two categories--
- gg long lines and short lines.
SRVDL E, a long line, was chosen for the single valve tests since the long line should represent the worst case for S/RV blowdown.
S/RVs D, E, and M were actuated simultaneously for the multiple valve tests.
The initial gas volume for all lines is shown in Table BNL-16-3.
Also, see 1.1 above.
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Il-eNL. l e-3R2 PUAR.00
I 3.3 Test Conditions Line D Line E Line M MS, Ib/see 268 268 268 TPL OF 78-81 78-81 78-81 TP*,
OF 238-355 220-379 246-361 LW, FT 7.0 7.0 7.0 A P, psid 1.2-1.33 1.2-1.33 1.2-1.33
- Initial and final tempera tures of drywell pipe gauge.
3.4 The minimum AP permitted by technical speci fi-
' cations is 1.10 psid.
The corresponding water leg length is approximately 7.5 feet measured from the quencher centerline elevation.
4.0 A description of.the design conditions for each load case used for design -
4.1 Figure 7.3 of the PUAR shows a plan view of the S/RV discharge in the torus.
This information as g
well as other in forma t ion regarding geome t ry is 3
available from TVA drawing series 47W401.
- Also, see 1.1 above.
4.2 The following parameters were ext racted from selected RVRIZ, RVFOR input.
Cas'e A1.1 (NOC)
SRVDL TP,0F TPL,0F MS,Ib/see AP,psid LW,ft E
115 75 308 1.2*
7.13 L
115 75 308 1.2*
7.13 Case C3.3 (IBA with Steam in DW, Second Actuation)
'SRVDL TP,0F TPL,0F MS,1b/see AP,psid, LW,ft E
350 90 308 1.0*
40.12 L
350 90 308 1.0*
31.56
- These' values were used in RVRIZ, a value of zero was used for RVFOR.
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H-BNL 16-4R2 PUAR.00
5.0 A description of the design loads for each load case -
5.1 Normalized pressure signature -
We interpret the term " normalized pressure I
signature" to mean the variation of the shell pressures with time.
Therefore, it is the same as the variable g(t) defined in Section 5.4.2.8 of the PUAR, and generated in accordance with the LDR by I
GE computer code QBUBS02.
Also, see the response to BNL Item 15.
5.2 The la rgest magni tude POP and PUP values genera ted by QBUBS02 were applied for each SRVDL in the torus.
Per the LDR, first actuation pressures were conservatively assumed to be possible for second actuation (reflood) conditions.
The single valve values are as follows:
Pressure (psig)
Event POP PUP I
NOC or DBA 16.2 11.9 SBA or IBA 19.7 16.6 I
5.3 The spa t ial a t tenua t ion functions used were as defined in the LDR and generated by QBUBS02.
The only deviation from the LDR in this regard was the use of SRSS for combining the ef fects of mult iple valve actuations.
The SRSS issue is addressed by the response to BNL Item 14.
5.4 The frequency range used for design, as provided by E
Section 5.5.2 of the PUAR, is as follows:
Frequency (IIz )
Event Minimum Maximum NOC or DBA 4.16 10.29 SBA or IBA 5.58 14.69 1
II-UNL 16-SR2 PUAR.00
8.0 --
6.0 --
f 4.0 --
p 2.0 --
3 O.
- 2.0 --
- 4.0 --
- 6.0 --
-8.0 0.0 0.2 0.4 0 '. 6 0 '. 8 If0 lj2 If4 If6 If8 2f0 6.0 --
5.0 --
4.0 --
3.0 --
p 2.0 --
1 10--
4 0.0
- 1.0 --
-2.0 --
- 3.0 --
- 4.0 --
-5.0 0.0 012 0.4 0 '.6 0.8 If0 if2 If4 If6 18 20 X100 TIME SECONOS FIGURE BNL-16-1 TYPICAL PRESSURE TRACES FOR BFN TORUS SHELL i
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TABLE BNL-16-1 MAXIMUM AND MINIMUM PRESSURES (PSI) g I
SINGLE VALVE ACTURTION TESTS 4
S1 S2 S3 S4 S5 E
1 5.648 6.417 6.714 6.695 5.971
-5
-4.950
-5.596
-5.538
-5.545
-5.293 2
5.870 6.743 7.233 7.112 6.232
-5.203
-5.898
-5.827
-5.849
-5.529 3
5.393 6.455 6.979 6.550 5.800
-4.825
-5.386
-5.451
-5.429
-4.949 4
4.593 5.285 6.004 5.709 5.183
-3.950
-4.477
-4.650
-4.509
-4.195 7
3.105 4.077 5.297 4.620 4.151 zo
-2.354
-4.097
-4.258
-3.929
-3.467 g
8 5.195 5.746 5.501 5.814 5.113 g
-3.554
-3.990
-3.969
-4.112
-3.915 J
9 1.757 1.670 1.878 1.810 1.650 I
-1.113
-1.482
-1.341
-1.354
-1.341 10 2.539 2.737 2.621 2.751 2.512
-1.716
-1.988
-1.961
-2.008
-1.900 12 2.226 2.369 2.580 3.057 2.764
-1.648
-1.859
-2.036
-1.838
-1.634 13 6.158 6.996 6.757 6.919 7.046
-4 755
-5.397
-5.249
-5.333
-5.383 14 4.887 5.615 5.517 5.496 5.461
-3.655
- 3.907
-3.732
-3.858
-3.970 I
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I
I TABLE BNL-16-2 l
MAXIMUM AND MINIMUM PRESSURES (PSI)
I MULTIPLE VALVE ACTUATION TESTS I
M1 M2 M3 M4 1
4.685 6.029 8.123 7.005
-5.997
-6.062
-6.333
-6.359 2
6.211 7.283 9.135 0.617
-6.310
-6.438
-6.771
-6.786 3
7.787 8.857 9.956 10.00
-6.150
-6.434
-6.688
-6.66 4
9.204 10.01 10.79 1 0.66
-6.016
-6.485
-6.760
-6.491 7
6.471 6.310 8.523 6.216 2
o
-4.922
-4.613
-5.659
-5.686
.(
8 5.828 6.795 8.224 8.007 g
-4.623
-5.174
-5.331
-5.590 J
9 1.348 1.764 2.642 2.903
-1.831
-1.777
~ - 1.837
-1.509 10 2.151 2.669
-3.288 2.914
-2.192
-2.185
-2.383
-2.240 12 11.98 11.00 10.36 11.16
-7.482
-7.639
-7.162
-7.196 13 7.616 8.328 7.919 7.264
-6.108
-5.876
-6.524
- 5.953 14 5.314 6.140 6.343 5.076
-5.300
-5.356
-5.097
-4.579 I
g.
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I TABLE BNL-16-3 I
3 INITIAL GAS VOLUME (FT )
1 -
LINE A1.1 NORMAL OPERATING CONDITION BLONDOWN D
73.80 H
75.37 LONG LINES E
66.16 A
58.31 F
53.33 L
55.07 C
52.14 K
54.82 SHORT LINES 8
51.00 C
52.69 J
55.60 l.
N 54.82-M 54.93
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I ITEM 18:
What is the vertical location of the suppression pool temperature sensors in relation to the S/RV T/ Quencher centerline?
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RESPONSE
The sensors are located approximately 20 inches above the
(
T-Quencher centerline and at mid-bay.
(See PUAR Figure 10-7.)
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,,-Im le-le2 Po,e.oo
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ITEM 19:
Were there any exceptions to the AC for the hydrodynamic loads applied for analysis of the Torus At tached Piping?
If so, elaborate.
RESPONSE
Other than the general in terpre ta t ions elabora ted in Section 4. 2 o f t h e BFN PUAR, there were no speci fic except ions to the NUREG 0661 AC for the hydrodynamic loads applied for analysis.of the torus attached piping.
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I H-DNL 19-IR2 PUAR.00 l
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I ITEM 20:
I In the calculation of vartcus drag loads for DFN, the computer codes LOCAFOR, CONDFOR, TQFORDF, and TQFORO3 were used.
Do the algorithms of these codes follow approved AC procedures?
Sta te any except ions and jus t i fy them.
RESPONSE
The GE compu ter codes LouAFOR, CONDFOR, TQFORDF, and TQFORO3 were used in the calcula tion of various drag loads for DFN.
These codes were put up on Control Da ta Corpora t ion I
computers around the country for access by the di f ferent AEs per forming Mark I plant unique long-term program evaluations.
These codes were developed, documented, and verified by I
consultants under contract with GE, not by the AEs performing the Mark I analyses.
The codes are proprietary to GE and were only provided as " black boxes" with instructions on their use (including description of required input data)
I provided in the form of Application Guides.
The AEs (including TVA) therefore do not have the direct access to the speci fic algorithms of these codes which would be I
necessary to answer your quest ion defini t ively.
It is TVA's understanding, however, that the codes LOCAFOR, CONDFOR, and TEEQFOR, used for evaluation of pool swell, CO and chugging, and S/RV drag loads, respectively, follow approved NUREG 0661 I
AC procedures.
That means that TVA has defined only S/RV drag loads with codes not thought to speci fically follow all NRC-approved AC procedures.
The only signi ficant di f ferences between the approved code TEEQFOR (not used by TVA) and the TQFORBF and TQFOR03 codes, to TVA's knowledge, a re as follows:
TQFORBF - Th i s code is di f ferent in that two bubble pressure factors, BFAC(l) and BFAC(2), were incorporated to I
be used as multipliers of the negative and positive bubble pressures, respec t ively.
These were empirically developed factors used to obtain more realistic comparisons of code predictions to I.
Monticello test resul ts (see Appendix B of GE Application Guide 5, Revision 3 - a later revision o f PUAR Re f e r ence 6 3 ).
TQFOR03 - Th i s code is di f ferent in the bubble dynamics portion of the code which uses QBUBS03 instead of QBUBS02.
The result is that fa r more realis t ic load predictions are obtained from this code due I
I
"-8"' 2a->"2
~ ^ " "
I
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to the attenuation in bubble energy as it rises to 3
the surface of the torus pool.
Pa r t icula rly for E
structures located high in the pool, this code predic ts drag loads tha t are signi fican tly a t tenua ted in ampiitude w!th time.
While both of these codes are felt to be more realistic than the extremely conservative TEEQFOR code, it should be noted g
that the least conservative code, TQFOR03, was only used fo r 3
the downcomer S/RV drag load predictions.
That was because the downcomers are located very high in the pool where, realistically, the S/RV bubbles have a t tenua ted subs tan t ially from their exit strength.
Loads predicted for this structure using even the M FORDF code (which is slightly less conser-vative than TEEQFOR) were found to be unrealistically high.
For all submerged structures other than the downcomers, the still very conservative TQFORDF code was used to obtain peak S/RV drag force ampli tudes.
This code was used fo r the other submerged structures because it is cheaper to run than TQFOR03 and the degree of overconservatism in N FORDF versus TQFOR03 is not too signi ficant for structures in lower 3
elevations of the pool.
Other than the downcomers, most BFN g
submerged structures are in the lower pool elevations.
The use of TQFORBF and N FOR03 was belleved to be welI justified by' good engineering judgment and especially by the fact that TVA planned S/RV tests which were expected to support that judgment.
Additionally, the S/RV drag (except g
for the downcomers) were conservatively defined assuming g
worst-case peak load amplitudes, applied as s teady-s ta te harmonics at worst-case frequencies.
As expected, the S/RV tests provided conclusive evidence of the adequacy of the analytical appraoch for S/RV fluid drag loads.
Appendix C of the PUAR describes the correla tion of 3
analytical and test data.
For example, the analytical 3
approach for_the downcomers (using N FOR03 loads) was shown to overpredict stresses by a factor of four rela tive to single and multiple S/RV test results.
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H-BNL 20-2R2 PUAR.00
I ITEM 21:
Are there any di f ferences between Browns Ferry Uni ts 1,
2, I
and 3 which were significant enough to warrant separate analyses for any uni t ?
If so, state the differences and the analyses used.
RESPONSE
Torus I
The Units I and 2 tori are virtually identical.
Unit 3 used lighter construction in the following areas:
Location Units I and 2 Unit 3 Ring girder inside flange 1-1/2" x 12" 1" x 10" I
Ring girder web l-1/4" y 12" 3/4" x 12" Cradle edge plates 1-1/4" 12" 1" x 12"
'g All dynamic torus analyses and the defini t ion of modi fi-3 cations were based on the Unit 3 properties.
Mod i f i ca t ion studies showed that stiffening the ring girder-cradle system always improved performance.
Hence, the definition of I
modi fica t ion for Units I and 2 from the Unit 3 analysis results is conservative.
Vent System The vent systems for all BFN' units are virtually identical.
Torus Attached Piping External BFN torus attached piping configurations are 5
di f ferent for each BFN unit.
Therefore, generally a 3
separate piping analysis was required for every piping system on each unit.
Internal torus attached piping configurations are virtually identical from unit to unit but they are included in the external piping analytical models.
S/RV Discharge Piping Two basic S/RV piping configurations are used in each DFN I
torus as shown by PUAR Figures 7-8 and 7-9.
The arrangement of all 13 lines in each BFN torus is shown in plan view by PUAR Figure 7-3.
DFN S/RV drywell piping configurations vary from line to line and and in some cases unit to unit.
I Therefore, various analytical models were used for the S/RV piping in the drywells.
Nonsafety-Related Internal Structuies The nonsafe ty-rela ted internal structures are virtually identical for all BFN units.
Il-BNL 21-IR2 PUAR.00 I
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ITEM 22:
This is an additional item to respond to a verbal inquiry from BNL at the September 5, 1984 meeting.
DNL asked how g
close to impact with the main vent bellows does the pool g
swell come.
RESPONSE
Ca l cu la t ion s, based on the conservative LDR methodology, a
predict that the pool swell profile would come within g
1.1 inches of the bellows at the closest point.
If the pool swell were to slightly impact the bellows, the velocity would be very small, approaching zero at incipient impact.
Examination of the construction of the bellows leads to the assessment that it has very good local impact resistance.
g Any potential for damage or leakage would require impact g
over a large area, which, in turn, would require the pool to rise at least a foot or more above the bottom of the bellows.
TVA concludes that there is no signi ficant safety concern for pool swell impact on the vent system bellows.
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I II-BNL 22-IR2 PUAR.00 I
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Y
....... MX POOL SELL PROFILE l
- - 4.46 ' +
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'N
.s EL542'-10$*
u s_ _ g'3 s
e 5.07' 3.87*
3y-0*
u_
y l
,,,s, n
l Jg
=
7.75'
=
~
FOR POINT S ON VENT:
W Yg = 5.07' POOL AT X/R = 0.5 ys. VMAX YMAX
- 4*88'
.. POOL MisgES BY.00' OR Ae0UT l' AT TME CLOSEST POINT.
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FIGURE BNL-22-1 TYPICAL MID VENT BAY CROSS-SECTION I
s e
s
~
A a
&-a
- I 4
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I BFN PUAR APPENDIX I I
FINAL TVA RESPONSES TO NRC AND FRANKLIN RESEARCH CENTER QUESTIONS d
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l General Response to PUAR Questions DFN LTP analysis and design activity has proceeded on a schedule necessary to support installa tion of all I
modi fications during the Cycle 4 and 5 refueling outages of each unit, as required by NRC.
The first BFN Cycle 4 refueling outage began in April 1981, and most of the major modi fica t ion designs were comple te by May 1981.
Remaining modi fica t ion designs, primarily for torus attached piping external supports, were complete in time to support installa tion during the Cycle 5 refueling outages.
In order to sa t is fy schedule conmi tmen t s, it was necessary to make i n t e r p r e t a t i on s o f LDR a n d NUREG 0661 requirements I
based upon the best available in forma t ion a t the time of analysis.
Most of the interpretations were originally established in 1979 and early 1980.
A continuing effort to remove excessive conservat ism f rom load defini t ions I
and analys,is methods was made, particularly when that conservatism would result in unnecessary, impractical modi fica t ions.
When la ter information on load definitions and associated analysis methods became available, it was compared to the previous interpretations.
The la ter in fo rma t ion was used I
for reanalysis and associated design work i f a signi ficant unconservatism in the previous interpreta tion was indicated.
For example, the final downcomer t lebar/v-bracing modi fica t ion resulted from November 1981 changes in the DBA condensation I
oscillation la teral load definition.
Sometimes, later information was used to remove excessive I
conservatism in remaining analysis and design work.
For example, the 1.I SRSS load combination technique was permitted for torus attached piping analysis after NRC's I
final position on this subject was defined in April 1983 by PUAR Reference 58.
An absolute summation combination technique was required prior to that time.
Finally, when the later information showed the previous load definitions and analysis methods to be adequately (but not excessively) conservative, the original interpretations were retained.
In these situations, reanalysis utilizing the later information would have been unnecessary and costly, and, in some cases, would have resulted in delays in the ins talla t ion o f modi ficat ions.
Many of the PUAR questions derive from situations where the original interpretations stated in PUAR Section 4 were used for analysis.
Justification for these interpretations was I
provided in PUAR Section 4, Section 5, and Appendix C.
Additional technical justification follows in the responses to speci fic quest ions on these topics.
Other PUAR questions I
simply request additional information, which is provided in the responses.
I l-GR-IR2 PUAR 04 I
I It is TVA's position that the BFN PUAR and our review question l
responses demonstrate compliance with the in ten t o f the Ma rk I Containment Long-Term Program and NUREG 0661 (i.e.,
to upgrade the containment system safety margins, for all postulated hydrodynamic loading conditions, to those intended by the original design speci fications).
On this basis, we feel that all indicated safety concerns are fully and satis factorily addressed, and we respectfully request a favorable final evaluation for the BFN LTP.
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1.co.2n2
,Uso.o.
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I ITEM l_:
Provide a more detailed description of the vent system analysis regarding downcomer lateral loads (Section 4.4.5 (5)).
RESPONSE
As described in the LDR, the condensation oscillation lateral load is simulated for each downcomer pair by adding a di f ferential pressure for one downcomer to the internal pressure, that occurs in both downcomers, thereby producing I
a higher load in one downcomer than in the other.
- Thus, from Figure 4.4.3.4 of the LDR, a darkened downcomer indicates that the differential and internal pressures are I
working simultaneously, whereas the other downcomer only experiences the internal pressure.
A 450 beam model, Figure 6-2, was used to analyze both IBA CO and DBA CO.
Based on the primary downcomer swing frequency extracted from a modal analysis of the system, sinusoidal forcing functions were applied to the downcomer pairs considering the load cases defined by Figure 4.4.1-3 in the LDR.
Since the primary swing mode for the DFN Sys tem occurs in the 8 !!z range, the 1st, 2nd, and 3rd harmonics were addressed by sinusoidal functions in the 4, 8, and 12 liz ranges, respectively.
An added conservatism in the DFN analysis was the application of the first harmonic forces to the coincident 8 IIz swing frequency.
The response of the system to this single frequency load enveloped responses frem the sum of the three harmonics defined by the LDR.
Also, the first harmonic force ampli tudes were applied with 16 and 24 lfz sinusoidal functions to verify that higher frequency responses I
do not impact the total CO response.
In fact, 30 individual sinusoidal functions were applied for each load case to account for potential harmonics at the 1/2, 1,
1-1/2, 2, and I
2-1/2 harmonics of the six discreet primary swing mode frequencies in the 8 to 9 IIz range.
The DFN vent system was analyzed for the four ini tial I
di f ferential pressure cases specified by a May 1981 draf t o f LDR Sec t ion 4.4.3.
(See PUAR Figure 6-12.)
Subsequently, four addltlonal mirror image cases were included in the I
final CO la teral load definition.
Evaluation of both the initial four cases and their four mirror images demonstrated I
g
- i. pac g.In2 PUAR.04 I
that Load Case I is controlling and additional rigorous analysis of the mirror images was unnecessary.
(See response to DNL Item 11 for further discussion.)
The DBA CO downcomer lateral load effects were combined with DBA CO fluid drag loads and other loads in the controlling load g
combinations and evaluated to the governing stress levels.
g The chugging lateral loads were calculated in accordance with the LDR and NUREG 0661, using frequencies from the modal analyses performed on the 450 and 1800 vent system beam models.
These loads were applied to single downcomers chugging exclusively and to multiple downcomers chugging synchronously.
The 450 beam model (PUAR Figure 6-2) was used for analysis of single downcomer chugging la teral loads, and the 1800 beam model (PUAR Figure 6-3) was used for analysis of downcomer synchronous chugging la teral loads.
The result ing ef fects were then combined with other loads, including chugging fluid drag loads on the tlebars and V-bracing for s tress and fa tigue evalua tion of the entire vent system.
Stresses were determined by applying approprl_ ate intensi-fication factors at intersections and by direct application of loads to finite element' models shown on PUAR Figures 6-5,,
6-7, and 6-11.
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,UAR.o.
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I ITEM 2:
Provide the physical details of the seismic lugs that restrain the torus against horizontal seismic motion yet allow thermal growth.
RESPONSE
The erection drawing for the seismic lugs is PDM-E12.
Fabrication details for the components are shown on PDM drawing 41.
(Copies of the drawings are available for review.)
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I I-FRC 2-IR2 PUAR.04
I ITEM 3:
Indicate how the ring girders were analyzed for loads from attached internal structures.
Any dynamic load factors that may have been used in the analysis must be provided and justified.
RESPONSE
The ef fects of the larger systems on the ring girder and other portions of the torus were considered as follows:
1.
Vent System:
The support column reactions due to vent header pool swell inpact were considered directly, as described in Section 5.4.2.7 of the PUAR.
In addition, the vent system masses were included in the dynamic 22-1/20 torus model, so that the ness t imes accelera t ion (rigid body) inertial effects were developed for all dynamic loads.
2.
ECCS IIcad e r :
The torus cradle stresses due to the reaction loads at the ECCS header supports were added to the stress intensities in that region for the torus model, without exceeding allowables.
The mass of the ECCS was also included in the dynamic torus model.
Thus, the rigid body portion of the ECCS header support reactions were conservatively included twice.
3.
IIPCI and RIIR:
The nesses of these sys tems were also included in the dynamic torus model.
The other systems were judged not to af fect overall torus behavior, but to produce only localized effects.
Support connections to the ring girder were heavily reinforced.
The line-of-action of pipe bracing members was applied near the base of the ring girder to eliminate any signi ficant overturning tendency.
For example, see PUAR Appendix G Plates 18 and 21.
Additionally, the suppor t sys tem for S/RV discharge lines and quenchers includes a 15-inch x 15-inch box-beam which forms a continuous ring inside the torus and prevents any possibility of overturning each ring girder in the region of a t tachment.
Th e nm in suppo r t membe r s fo r t he ca twalk per form a similar func tion a t each r ing girder.
See PUAR Plates 12, 13, 26, and 27.
Local acceleration response spectra for each dynamic load were defined at each ring girder attachment point.
The a t tached piping sys tems and s tructures were analyzed for these input spectra and associated displacements.
I l-FRC 3-IR2 PUAR.04
Reactions were calculated from the piping analyses per Sections 7 and 8 of the PUAR, and for the catwalk per Section 9.
The local reinforcement was designed for these reactions.
The localized stresses transmitted to the ring girder were limited to 3 ksi.
When combined with the general ring girder stresses, no allowables were exceeded.
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ITEM 4:
With respect to the 22-1/20 torus model mentioned in Section 5.4.1.1 of the PUAR (5), the boundary conditions are based on the assumption that all loads are applied equally to each of the 16 segments.
However, the safety-relief valve and chugging loads are asymmetrical.
Jus t i fy the use of a 22-1/20 model to evaluate the torus for S/RV and chugging instead of the 1800 model required by the criteria (1).
RESPONSE
The following points pertain:
1.
Symmetric loading, for S/RV and chugging, is certainly bounding for cradle loads just from the point of view of the magnitude of the net applied load.
2.
Shell responses are primarily a localized phenomenon.
They can be affected by ring girder ovalling, but this too is related to the net load, and so would be more severe for symmetric loads.
3.
The BFN torus support system inhibits asymmetric response.
The development of asymmetric modes would require longitudinal motion of the torus, which is prevented by the seismic lugs.
It would also require radial movement of the ring girders which is prevented by the torus snubbers (PUAR Pla te 1).
Friction at the cradle support pads would also inhibit any tendency for asymmetric response.
4.
If significant asymmetric response could develop, it would have been present in the single-and multi-valve S/RV tests.
None was evident, and analysis results based on symmetric loading boundary conditions were shown to be conservative (PUAR Appendix C).
Finally, it is important to recognize that Section 6 of the PUAAG (Reference (1) of the questions) is not a criteria.
It is a guideline for analysis methods.
I l-FRC 4-IR2 PUAR.04
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ITEM 5:
Figure 5-6 in the PUAR (5), which depicts the 1800 model I
o f the torus, shows only the lower half of the torus shell.
Indicate whether the model includes the torus supports.
I
RESPONSE
Figure 5-6 of the PUAR depicts only the lower half of the I
1800 model for clarity.
The actual model includes the upper and lower halves.
All supports are included (i.e.,
the torus snubbers, seismic lugs, and the support pad-tiedown system).
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ITEM 6:
Since NRC Regulatory Guide 1.61 (4) deals with damping values for the setsmic design of structures, explain how this Regula tory Guide validates the use of 4 percent damping for the 0.0 AP pool swell analysis of the torus (Section 5.4.2.7 (5)).
RESPONSE
The use of NRC Regulatory Guide 1.61 damping was accepted for analysis by Section 4.4.2 of NUREG 0661, and the 0.0 oP pool swell case was designated as a Service Level D condition by Section 4.3.3.1 of NUREG 0661.
Regulatory Guide 1.61 speci fies 4 percent damping for welded s teel structures under SSE loading which is normally associa ted with Ser. ice Levels C and D conditions.
The torus and vent system are welded steel structures.
Two percent damping was conservatively used for BFN vent system 0.0 AP pool swell analysis.
Four percent damping was used - for the torus.
This assumption in combination with the overall analysis method produced a reasonably conservative dynamic response prediction.
BNL Items I through 4 provide additional infornation on the BFN pool swell analysis method.
Finally, it is noteworthy that two percent damping was conservatively assumed for all Service Level C load combination torus and vent system analyses, to reduce the number of cases for analysis.
Four percent damping is considered justifiable for the Service Level C and D load combinations on the basis of Regula tory Guide 1.61.
I I-FRC 6-lR2 PUAR.04 I
ITEM 71 With respect to Section 5.4.2.11 of the PUAR (5), provide the technical basis and justification for considering the forcing functions from 0 to 30 Hz instead of the full 0 to 50 Hz for post-chug analysis of the torus.
RESPONSE
The intent of the discussion in Section 5.4.2.11 of the PUAR was to emphasize the following mejor points:
1.
There are signi ficant consstrva tisms in the PIN post-chug analysis method which offsot the effects of not considering the harmonics in the 30 to 50 Hz range.
2.
The 30 to 50 Hz harmonics were considered in the drag load analyses.
(See Section D.I.2.4.1 of the PUAR.)
3.
Pre-chug generally controls over post-chug for torus analysis.
The DFN pre-chug analysis was performed in complete accordance with NUREG 0661 and the LDR, I
including additional conservatisms inherent in the method.
4.
Any remaining concern with the response of high I
frequency modes for torus attachments is offset by the high frequency content in the pool swell analyses.
Finally, a strong empirical indication (not a rigorous analytical proof) of the conservatism of the BFN chugging analyses (both pre-and post-chug) is seen in the attached tab 3e.
BFN shell surface stress and support reaction I
forces are presented, factored as nearly as possible to an FSTF-cquivalent basis, and compared to measured and calculated NEP values for the FSTF.
The conservatism of the BFN responses to post-chug analysis results is primarily due to the absolute summation of all 30 harmonics in the O to 30 Hz range.
(Note that the dominant BFN torus modes are in the O to 30 Hz range. ) PUAR Reference 20 reconmends absolute I
sunmation of 5 harmonics plus SRSS of remaining harmonics to achieve an 84 percent NEP.
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I-FRC 7-lR2 PUAR.04 I
TABLE FRC-7-1 COMPARISON OF FSTF AND BROHNS FERRY CHUGGING RESPONSES.
i NAx ph, 00E TO POST-CNUS M Y BFN RESPONES PER F
, PGt FS RESULTS p
RESPONSE
PRE-POST-PM-POST-T 5-1 TABLE 4-1 TAEK 4-1 i
l BDC SURFACE STRESS I ENSITY 1.10 1.12 1.31 1.33 0.90 0.98 0.98 INSI RE CTION 102 100 57 58 31.2 17.0 19.2 OUTSI gCTION 113 110 64 S2 32.3 17.7 20.2 (1) FSTF EQUIVALENT SHELL STRESS INTENSITY = 8
" 1 19 b SFN gg BFN MHERE, R = NINOR RADIUS OF T E TORUS, T = SNELL THICKNESS, AND K = STRESS INTENSITY
^
(2) FSTF EQUIV. SUPPORT REACTIONS =(BFN REACTION)
=0.562 (BFN REACTION)
ITEM 88 I
Items 2 and 3 in Section 5.4.2.11 of the PUAR (5) suggest that the pre-chug load bounds the post-chug load in the analysis of the torus; however, Item 5 in Section 5.4.2.11 indicates a higher surface stress for post-chug.
Explain I
this apparent inconsistency and indicate whether pre-chug or post-chug was considered in the controlling load combinations for the torus.
RESPONSE
The discussion of Item 5 in Section 5.4.2.11 of the PUAR demonstrates the inherent conservatism of the BFN post-chug analysis, relative to the FSTF data.
The discussion,of I
items 2 and 3 of 5.4.2.11 show that the LDR prescribed pre-chug analysis method bounds the actual measured FSTF chugging responses due to both the pre-and post-chug phases.
This is not to say that pre-chug will always bound post-chug.
It only says that consideration of the pre-chug phase alone is sufficient to demonstrate the conservatism of torus results based on the LDR method as compared to actual measured FSTF results for the combined pre-and post-chug phases.
Also see the response to FRC Item 7, including Table FRC-7-1.
For all load combinations involving chugging, the maximum stress due to either pre-or post-chug was used (i.e., an envelope of pre-and post-chug responses).
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I I-FRC 8-1R2 PUAR.04 I
I ITEM 9:
With respect to the fatigue analysis of the torus presented in Section 5.4.6 of the PUAR (5), speci fy the elasticity methods used to calculate stress in tens i fica t ion factors a t the penetrations.
RESPONSE
The stress intensi fication factors presented in Section 5.4.6 of the PUAR were calcula ted using formulas presented in the following text:
Formulas for Stress and Strain, 5th edition, by R. J. Roark and W.
C. Young, McGraw-fli 11.
The insert pad to shell junction intensification factor was g
based on Case 14, page 598.
For the insert pad to nozzle g
junction, Case 5, page 593 was used.
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I I-FRC 9-IR2 PUAR.04
I ITEM 10:
l Provide and justify the bounding technique used to determine the controlling load cases presented in the PUAR (5) in the foilowing sections:
5.5.1, page 5-21 6.3.2 (and Table 6-5), page 6-6 6.4.2 (and Table 6-7), page 6-7 I
1 6.5.2 (and Table 6-9), page 6-8 6.7.2 (and Table 6-12), page 6-12 6.8.2 (and Tables,6-15 and 6-16), page 6-14 6.9.2 (and Tables 6-17, 6-18, and 6-19), page 6-15 i
I 7.3.1 (and Table 7-1), page 7-7
)
7.4.1 (and Tables 7-2 to 7-4), page 7-12 8.2.22 (and Table 8-2), page 8-3 I
9.1, page 9-2 9.2, page 9-2
RESPONSE
PUAR Seetion 5.5.I No bounding techniques were used for determining controlling load cases for the torus analysis.
All load combinations were analyzed, including the calculation of stress I
intensities and reaction loads for the cradle and torus snubbers.
The referenced section was stating which of the combinations produced the highest stress intensities.
PUAR Sections 6.3.2 through 6.9.2 PUAR Tables 6-5, 6-7, 6-9, 6-12, 6-15, 6-17, 6-18, and 6-19 I
give the event, combination, and service level of the vartous locations of inspectlon.
Table 3-1 of the PUAR shows this information in a different form.
The bounding technique for the tables in PUAR Section 6 was such that I
Service Level C event combina t ions would be quali fied using Service Level B allowables when possible.
When this was not possible, actual service level combinations coincided with I
the. assigned service level stress limits in PUAR Table 3-1.
for each A logie description of the bounding justification
,I tabic follows:
g g
1-FRC 10-IR2 PUAR.04 y
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PUAR Table 6-5 Logic o
Load Combination (LC) 15 to Service Level (SL) B allowables envelops LC 1 through LC 14.
o LC 27 to SL B allowables envelops LC 17, LC 20, LC 21, LC 23, and LC26.
o LC 25 to SL B allowables envelops LC 16, LC 18, LC 19, LC 22, and LC24.
o IBA is not indicated in LC 15 because:
(1)
SBA chugging is no less severe than IRA chugging.
(2)
IBA CO is enveloped by DBA CO in LC 27.
PUAR Table 6-7 Logic o
LC 18 to SL B allowables envelops LC 16.
o For the vacuum breaker to main vent cap inter-section, dynamic loading due to pool swell vent response far exceeds any chugging, CO, or S/RV g
effect.
Therefore, evaluation of LC 1 through I
LC 15 plus LC 17, LCs 20 through 23, and LCs 26 through 27 is not necessary.
o LC 18 does not envelop LC 19, LC 24, and LC 25.
However, the load contribution from SSE versus OBE-and the S/RV contribution are small and have been g
neglected considering the increased allowable for g
o Note that the pool swell load considers pool swell l
vent response and direct impact of the swell on the 3
vacuum breaker shell.
(The vacuum breaker shell is actually partially shielded by the vacuum breaker access platform.)
PUAR Table 6-9 Logic o
This logic is similar to PUAR Table 6-5 except LC 27 could not be satisfied for SL B primary plus secondary stresses (PL*Q<3Sme).
Therefore, l
g LC 21 was evaluated for primary plus secondary g
instead.
(LC 21 is the same as LC 27 with no S/RV.
PUAR Table 6-9 incorrectly. Indicated LC 27 instead of LC 21.
This correction was made in PUAR revision 2.)
I-FRC 10-2R2 PUAR.04
I PUAR Table 6-12 Logic o
Again this logic is similar to PUAR Table 6-5 except LC 27 would not meet SL B allowables.
Therefore, LC 27 was evaluated to SL C allowables and LC 21 was evaluated to SL B allowables for bo th pr ima ry and I
primary plus secondary stresses.
(PUAR Table 6-12 incorrectly indicated IE 27 instead of LC 21 for the Service Level B combinations.
This correction was I
made in PUAR revision 2.)
PUAR Tables 6-15 and 6-16 Logic o
Again this logic is similar to PUAR Table 6-5.
The noted loads are the worst possible combination of g
any accident condition, including thernal effects, and the buckling evaluation is performed on this
^g basis.
PUAR Tables 6-17, 6-18, and 6-19 Logic o
Again, IE 15, IE 25, and LC 27 are the worst case load combinations for the system.
The indicated I
stresses are maximum surface including secondary effects with significant margin against 1.5 Sme-Therefore, the primary plus secondary stress range evaluation is automatically assured.
PUAR Section 7.3.I Section 7.3.1 provides a general description of the bounding technique that was used to determine controlling load cases for S/RV piping in the drywell a t Browns Ferry.
Speci fically,. the controlling load cases for drywell S/RV I
piping were determined by the following process:
(1)
Survey all defined normal, seismic, and LOCA load I
definitions to determine which of these have significant effect on drywell S/RV piping.
Note that separate models of the piping systems were I
developed to analyze the drywell and wetwell portions o f the sys tem.
The wetwell models were developed in significant detail to study torus hydrodynamic gg phenomenon closely.
These models extended a significant distance into main vent to account for a t tenua t ion.
It was found that the S/RV piping in the main vent is isolated from most of the hydrodynamic ef fects of S/RV I-
' discharge or LOCA excitation of the suppression pool.
(An exception to this is the containment vent response I
I-FRC 10-3R2 PUAR.04 I
I induced by DBA LOCA pool swell.)
The drywell piping is sufficiently removed from the suppression pool to discount effects from water clearing transients in wetwell portions of the S/RV lines.
The load sources that were determined to have a significant ef fect on drywell S/RV piping are:
a.
Deadweight b.
Al1 S/RV blowdowns d.
Pool swell vent response e.
Thermal expansion f.
Pressure (2)
Perform an inspection of Table 5-2 in the PUAAG (PUAR Reference 13) considering the resultant S/RV load sources noted in step 1 above.
A summary of the findings with respect to PUAR Table 7-1 follows.
Case 1:
Sa t is fies LC 1.
Case 2:
Sa t is fies LC 3 which envelops LC 2.
Case 3:
Sa tis fies LC 15 which envelops LC 4 thrcugh LC 14 except as indicated by note 6 on PUAR Table 7-1.
Case 4:
Sa tis fies LC 27 which envelops LC 16 through LC 26.
PUAR Seetion 7.4.I Section 7.4.1 provides a general discussion of the bounding technique for wetwell S/RV piping.
All load sources are treated.
A suninary of the enveloping logic for PUAR g
Tables 7-2, 7-3, and 7-4 with respect to Table 5-2 of the g
-PUAAG (PUAR Reference 13) is listed below.
P_UAR Table 7-2 Case I and Case 2:
Sa t i s fi es LC 1.
l Ca se 3 and Ca se 4 :
Sa tis fies LC 3 which envelops LC 2.
I I-FRC 10-4R2 PUAR.04
I PUAR Table 7-3 Case 1 and Case 2:
Sa t is fies LC 11, row 11, which envelops LC 4 through If 10 for row 11.
Case 3 and Case 4:
Sa t is fies LC 15, row 11, which envelops LC 12 through LC 14 for row II.
I Case 5 and Case 6:
Sa t is fies LC 15, row 10, which envelops LC 4 through If 14 for row 10.
PUAR Table 7-4 Case 1:
Sa t is fies LC 27 which envelops LC 17, LC 20, LC 21, LC 23, and LC 26.
(Note that CO and chugging do not occur simultaneously and chugging is addressed more conservatively in I
PUAR Table 7-3 for SBA/ IBA even ts. )
Case 2:
Sa t is fies If 25 which envelops LC 16, LC 18, I
Case 3:
Sa t is fies LC 16 for the 0.0 AP case.
(Due to the low probability of occurrence, S/RV blowdown I
and earthquake are not assumed to be concurrent with the 0.0 AP pool swell.
This is in accordance with the PUAAG.)
-I.
PUAR Seetion 8.2.2.2
.3 The boundary technique used to reduce the number of load
' 5 case combinations shown in PUAAG Table 5-2 (PUAR Reference r
- 13) to those shown in PUAR Table 8-2 was based on using the most conservative combination of load cases associa ted with each of the service levels and ASME Section III, NC-3600 (PUAR Reference 68) equation 9 stress limits.
(Note that for dif ferent local combinations, the stress lirai t s could be 1.2 S, 1.8 S, or 2.4 S, corresponding to Service Levels B,
, I C, and D, respectively. ) The following are two examples of how this bounding technique was applied:
(1)
When two series of load combinations listed in the PUAAG were the same except that one included OBE and the other SSE and both sets of load combina tians had the same stress limits, the OBE and SSE losd cases I
were enveloped.
In this way the two sets of load ecmbinations could thus be reduced to one.
1-FRC 10-5R2 PUAR 04 I
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(2)
Another example would be when one set of load 3
combina t ions consis ted of all the load cases found in g
another set of load combinations plus at least one more load case.
If both sets of load combinations had to meet the same stress limits then only the combination with the greater number of load cases was evaluated.
The controlling load combinations for each service level g
equation 9 stress limit were found in this way.
In additlon, 3
NC-3600 equation 10 or 11 was sa tis fied for each of the controlling load combinations.
PUAR Section 9.1 The new catwalk finite element model was analyzed for all 3
applicable load events.
The results showed that the largest g
stresses, by far, were due to pool swell impact and drag.
Therefore, the most severe condition involving this load was limiting.
For Load Combination 25, the ef fects of pool swell impact-drag, pool swell and vent header motions, S/RV motions, deadweight, and SSE were added absolutely.
PUAR SeetIon 9.2 In the same manner as the catwalk, the vacuum breaker valve pla t form is most severely affected by pool swell impact-drag loads, and the 1imiting load combination was determined in the same way.
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I-FRC 10-6R2 PUAR.04 I
I ll ITEM 11:
Provide the stress results from the analysis of the torus shell and supports.
I
RESPONSE
Stress intensi t ies were calcula ted by postprocessor computer I-codes for all load combinations and for all elements in the finite element model.
The results were screened to locate predicted overstresses.
Following modi fica t ions, all stresses were below allowables.
The most highly stressed torus support locations are in the cradle, adjacent to the scab plates described in Section I
5.2.4.3 of the PUAR.
The addition of the scab plates reduced the local cradle stresses for Load Combina t ion 14 from 24.4 ksi to 19.7 ksi.
Relative to the Service Level B allowable of 21.6 ksi (see Section 5.3.2 of the PUAR), the I
stress factor was reduced from 1.13 to 0.91.
The maximum stress for any Service Level C or D load case was in the same region of the cradle and was due to Load Conbina tion I
25.
Even without the scab plates, however, the maximum stress intensity was 26.8 ksi, compared to the 28.8 ksi allowable.
These cradle stress results conservatively I
neglect the S/RV load reduction factor for torus supports defined in PUAR Section C.9.1.
The shell and ring girder stress allowables presented in I
Table 5-1 of the PUAR were never approached except in the vicinity of large piping penetrations.
A number of the piping modifications described in Section 8, as well as the nozzle reinforcements and local shell reinforcement around I
the ECCS header penetrations were required in order to meet the containment vessel (ASME Section III, Subsection MC) allowables.
At some of these locations, the calculated I
stresses are greater than 90 percent of Service Level B allowables for the primary plus secondary stress intensity range.
An indication of the general shell stress state away from the influence of penetration loads is given by the following t a b l e..
Membrane and surface stress intensities are presented I
for mid-bay bottom-dead-center, one of the more highly stressed locations, for the most critical load combinations.
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I-FRC ll-lR2 PUAR.04
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Stress Intensity (ksi)
% of Allowable Load Service Membrane +
Membrane +
Comb.
Level Membrane Bending Membrane Bending 14 B
15.1 15.4 78 53 18 B
5.1 5.4 26 19 20 B
9.4 9.9 49 34 g
25 C
8.9 9.3 23 16 g
27 C
13.3 13.9 35 24 16B D
6.2 6.8 15 11 I
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I-FRC ll-2R2 PUAR.04 l
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ITEM 12:
Regarding the analysis of the main vent /drywell inter-section, clarify whether the seismic and thermal response of the drywell was considered (Sections 6.2.1.2.i and 6.2.1.2.9 (5)).
RESPONSE
Thermal growth of the containment shell was considered in the analysis of the main vent /drywell intersection.
Thermal I
displacements were calculated for the drywell based upon maximum alr temperatures occurring during the DBA, IBA, and SBA events.
These displacements were input at the nodes representing the drywell/ main vent intersectlon.
The thermal analysis was then completed considering expansion and restraint-of free end displacement of the vent cystem.
I The BFN LTP seismic analysis was based upon the methods employed in the original plant design.
Seismic response of the drywell/ main vent intersection was analyzed using equivalent static loads determined f rom appropr ia te I
acceleration levels of the vent system.
This is consistent with the general guidelines of NUREG 0661, Sect ion 4.4.1 as welI as the PUAAG (PUAR Reference 13).
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i-rRc 12-m2 PU
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ITEM 13:
Provide a sunmary of the analysis of the vacuum breaker valves; indicate whether they are considered Class 2 l
components as required by the criteria (1).
m
RESPONSE
Apparently this request relates to analysis of the drywell/
wetwell vacuum breakers for cyclic loads occuring during chugging events.
Since this concern is not part of the Mark I Containment Long-Term program, a separate response was sent to the NRC on November 5, 1984 (NEB 841113615).
If this request relates to the new 10-inch S/RV vacuum "o r e a k e r s, these valves have been analyzed and subsequently mod i f ' ed to sa t is fy ASME Sec t ion III Class 2 stress limits l
for all postulated conditions including opening impacts.
e The modified S/RV vacuum breakers are shown on TVA drawing 47W401-9.
l ADDENDUM Additional information on qualification testing of the modi fied S/RV vacuum breakers for opening impact. loads was requested during the September 13, 1984 meeting with NRC g
and FRC representatives.
That information follows:
E Preliminary forcing functions for design of opening impact modifications were based on conservative predictions B
extrapolated from Monticello test results.
Modification 5
designs were made and preliminary tests for short-term adequacy were conducted on that basis.
The final forcing function for the S/RV line E vacuum breaker was determined from discharge line pressure measurements taken during the April 1983 BFN S/RV tests E-(PUAR Reference 41).
This forcing function was analytically 5
extrapolated for all BFN S/RV lines and all long-term program load condi t ions.
Then a prototype vacuum breaker g
was tested at Wyle Laboratory, Huntsville, Alabama, to g
demonstrate operability for all forcing functions and the full.40-year plant life.
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I I-FRC 13-IR2 PUAR.04 I
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ITEM 14:
The PUAR (5) indicates that the calculated stress values I
at the following locations are very close to the respect ive allowables:
o downcomer/ vent header intersection (Section 6.5.4.1) o downcomer/tiebar intersection (Section 6.7.4.1)
Indicate conservatisms in the analysis to show that these I
calculated values would not be exceeded if a di f ferent analytical approach were to be used.
RESPONSE
Although the stress values at the intersections mentioned
. I above were close to the respective allowable stresses, this should not represent a signi ficant concern.
Design modifications were made such that the stresses resulting I
from the new configurations were just below the acceptable values.
This would normally be ant icipated.
There are conservatisms which could be removed to obtain a I
greater difference in the allowable and actual calculated stress values for the above intersections.
For example:
(1)
Absolute summation was used in the combination of loads.
(2)
The downcomer/tiebar was conservatively analyzed using I-Service Level B allowables for Service Level C loads.
(SSE seismic loads were included in the actual loading in place of the prescribed OBE seismic load.)
(3)
There are other conservatisms associated with the DBA CO la teral load analysis method as described in the response to FRC Item 1.
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l-FRC 14-lR2 PUAR.04 I
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ITEM l_5:
Stress in tensi fica t ion fac tors for the miter bends in the vent system are not found in Table 6-17 as stated in Section 6.9.1 of the PUAR (5).
Provide these factors.
RESPONSE
Main vent miter bend SIF
- 3.85 Vent header miter bend SIF - 8.2 Downcomer miter bend SIF
- 3.82 (See Table 6-21 of revision 2 of the PUAR. )
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.I I-FRC 15-lR2 PUAR.04 I
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ITEM 16:
l Regarding the torus bellows analysis in Section 6.10.1.1 of the PUAR (5), provide the method and technical basis for calculating the spring values that represent the bellows I
flexibility in the computer models of the vent system (Figures 6-2 and 6-3 (5)).
RESPONSE
The spring values that represent the bellows flexibility
~g were calculated using the " Standards of the Expansion Joint 3
Manu fac turers Associa t ion, Inc." (PUAR Reference 24).
Appropriate data for the BFN bellows was input including convolution depth, thickness, number of convolutions, and I
modulus of elasticity.
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l-rRc le-m2 eU
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ITEM 17:
1 Provide and justify the approach for the fa tigue evaluat ion of the bellows mentioned in Section 6.10.3 of the PUAR (5).
I
RESPONSE
The fa tigue evaluation was carried out using " Standards of the Expansion Joint Manufacturers Association, Inc." (PUAR Reference 24) and'the Mark I Containment Program Augmented Class 2/3 Fatigue Evaluation Method and Results for Typical g
Torus Attached and S/RV Piping Systems (PUAR Reference 21).
g Deflections for the torus and bellows were obtained for each load event using computer analysis results and hand calcu-lations.
These deflections resulted in bellows stresses which were then combined in accordance with the fa t igue evaluation method noted above.
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I I-FRC 17-IR2 PUAR.04 I
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ITEM 18:
According to Section 7.3.3.1 of the PUAR (5), the safety-relief valve line penetration of the main vent was modeled using cylindrical shell flexibility characteristics.
Indicate the method for determining these characteristics.
RESPONSE
For the S/RV line penetration of the main vent, a six degree of freedom " support" was modeled.
For three degrees of I
freedom (pipe torsion and the two translational shear directions) full fixity was assumed.
For the ci rcumferent ial and longitudinal bending directions, Bijlaard's methods (PUAR Reference 64) were utilized to determine rotatlonal spring I
rates.
For the main vent radial direction, a translational spring rate was determined per the R. J. Roark text, Formulas for Stress and Strain.
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1-FRC 18-lR2 PUAR.04
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ITEM 19:
Provide the technical basis for obtaining the stress in tensi fica t ion fac tors used in the analysis of the safe ty-relief valve discharge piping system (Sections 7.3.3.1 and 7.4.3.1 (5)).
RESPONSE
In general, the 1977 ASME Code Section III (Figure ND-3673.2(b)-1) is the-technical basis for the stress 3
in tensi fica t ion factors utilized in the S/RV discharge 5
piping analysis with the following except ions:
Component Basis Weld-O-Let Bonny Forge stress intensification g
factors and stress indices for 5
weld-o-lets.
Sweep-O-Let Stress intensification factors and stress indices for the Bonny Forge sweep-o-lets.
Quencher Near Stress intensification factor based on Collar Support effeetIve section of quencher.
Assumes hole zone of quencher provides no a
structural contribution.
E i = Section modulus of 12 inch Sch 80 Pipe EffeetIve seetion modulus of quencher I
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I I-FRC 19-lR2 PUAR.04 I
'I ITEM 20:
Provide the stress results from the wetwell and drywell I
safety-relief valve discharge piping analysis (Sections 7.3.4.1 and 7.4.4.1 (5)).
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RESPONSE
Tables FRC-20-1 through FRC-20-4 provide a sunmary of the I
maximum equation 9 stresses from the S/RV discharge piping analysis.
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TABLE FRC-20-1 E
DRYWELL LOAD COMBINATIONS - MAXIMUM STRESS u
I LOAD CASE III NODE STRESS (KSI)
STRESS RA O
CASE 1 59 17.8 0.991 LINE CASE 2 59 20.5 0.760 C (2)
CASE 3 47 22.1 0.613 I
CASE 4 59 20.6 0.573 I
CASE 1 157 16.2 0.898 I
LINE.
CASE 2 157 23.8 0.883 E (2)
CASE 3 157 23.8 0.662 CASE 4 157 23.9 0.664 I
- 1. LOAD CASE PER TABLE 7-1 0F PUAR.
- 2. LINE C IS A REPRESENTATT/E SHORT LINE.
l LINE E IS A REPRESENTATIVE LONG LINE.
- 3. STRESS RATIO = ALLOHABLE STRESS I
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TABLE FRC-20-2 I
NOC - SERVICE LEVEL B AND C LOAD COMBINATIONS MAXIMUM STRESSES - HETHELL EVALUATION I
STRES8 LOAD CASE III NODE STRESS (KSI)
RA 0
CASE 1 ENVELOPED BY CASE 2 (2)
LINE CASE 2 2
10.3 0.S73 L
CASE 3 ENVELOPED BY CASE 4 (2)
CASE 4 2
10.9 0.404 I
CASE 1 ENVELOPED BY CASE 2 (2)
LINE CASE 2 8
13.3 0.7S8 H (3)
CASE 3 ENVELOPED BY CASE 4 (3)
CASE 4 8
14.1 0.S22
- 1. LOAD CASES PER TABLE 7-2 0F PUAR.
- 2. THE ENVELOPING OCCURS BECAUSE A HORST CASE NOC - BLOHOOHN I
(SCREENED BETHEEN FIRST AND SECOND ACTURTION) IS USED FOR THE STRESS ANALYSIS.
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- 3. LINE L - TYPICAL SHORT LINE.
LINE H - TYPICAL LONG LINE.
- 4. STRESS RATIO = ALLOHABLE STRESS I
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I TABLE FRC-20-3 g
SBA/IBA - SERVICE LEVELS C AND D LOAD COMBINATIONS MAXIMUM STRESSES - HETHELL EVALUATION I
LORD CASE (1)
NODE STRESS (KSI)
STRES8 RA 0
CASE 1 12 21.6 0.887 CASE 2 2
24.9 0.922 LINE CASE 3 ENVELOPED BY CASE 5 (2)
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CASE 4 ENVELOPED BY CASE 6 CASE 5 12 21.8 0.674 CASE 6 2
26.8 0.745 CASE 1 8
26.5 0.982 CASE 2 8
26.0 0.963 LINE CASE 3 ENVELOPED BY CASE 5 H
CASE 4 ENVELOPED BY CASE 6 CASE 5 10 31.7 0.881 CASE 6 8
32.2 0.994
- 1. LOAD CASES PER TABLE 7-3 0F PUAR.
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- 3. STRESS RATIO = ALLOHABLE STRESS I
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TABLE FRC-20-4 I'
DBA - SERVICE LEVEL D LOAD COMBINATIONS MAXIMUM STRESSES - HETHELL EVALUATION I
STRES8 LOAD CASE (1)
NODE STRESS (KSI)
RA 0
CASE 1 235 22.4 0.692 LINE CASE 2 1
33.4 0.927 I
L CASE 3 1
31.7 0.880 CASE 1 8
28.1 0.781
'I LINE CASE 2 8
28.3 0.787 CASE 3 48 27.4 0.760 lI
- 1. LOAD CASES PER TABLE 7-4 0F PUAR.
CALCULATED STRESS
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ALLOHABLE STRESS LI
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ITEM 21:
I Provide and justify the allowable safety-relief valve nozzle loads which were referred to in Section 7.3.4.2 of the PUAR (5).
RESPONSE
I As summarized in the PUAR, relief valve nozzle loads calcula ted in the drywell S/RV piping analysis were compared to a set of allowable nozzle loads used in the original S/RV l
analysis performed by Teledyne Engineering Services and 3
documented by PUAR Reference 45.
These allowable loads were provided by the relief valve vendor, Target Rock, and are g
incorporated in the design report for this component.
The g
allowables and worst case calculated loads are:
Allowable Resultant Valve Bending Moment from Flange Worst Load
- Dynamic Loads Inlet 320,961 IN-LB 400,000 IN-LB Outlet 287,568 IN-LB 300,000 IN-LB
- SRSS combination of dynamic loads.
In addition to the vendor allowable nozzle load check, the connecting' flanges for the relief valve installation were g
evaluated against static and static plus dynamic load
.3 allowables as calculated per the procedure of Paragraph NB-3658.1 in the 1977 ASME Boiler and Pressure Vessel Code (PUAR Reference 68).
The following table provides a summary of that evaluation:
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I I-FRC 21-lR2 PUAR.04 I
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l Valve Worst Load Allowable Flange Condition (IN-LB)
(IN-LB)
Inlet S
296843 437321 Outlet S
306205 372971 Inlet S + D (B) 390466 874642 Outlet S + D (B) 353947 745942 Inlet S + D (C,D)*
637582 1375829 Outlet S + D (C,D)*
S15009 1096778
- Direct addition of dynamic load components.
As can be seen, all loads are acceptable.
Condition notes:
S
= Static D = Dynamic I
(B) = Service Level B (C) = Service Level C i
l (D) = Service Level D l
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l-,ac 21-2B2 euAa.04 I
I ITEM 22:
With respect to Section 7.4.3.2.1 of the PUAR (5), provide and jus t ify all dynamic ampli fica t ion fac tors used in the calcula tion of safety-relief valve discharge-induced fluid drag forces on the safety-relief valve system.
RESPONSE
Sa fe ty-relief valve (S/RV) discharge-induced fluid drag forces were applied pseudo-statically to the S/RV system.
Th e 'IQFORBF comp u t e r code wa s u s ed wi t h S/RV l i n e i np u t properties that would produce the highest force ampli tudes possible from any line for any S/RV discharge case.
Maximum g
amplitude force-t ime histories were thus determined.
From g-these time histories, the peak amplitudes of the distributed forces were taken for equivalent s ta tic application to the system.
The equivalent pseudo-static force distribution was determined by conservatively assuming the distribution of a
peak forces to act as a perfectly steady-sta te sinusoidal g
forcing function.
A modal analysis of the S/RV system indicated that there were no potentially responsive modes of the system within the broadened 4.2-14.7 Hz range (see l
PUAR Section 5.5.2) of possible S/RV forcing frequencies.
5 Therefore, the lowest potentially responsive natural frequency of the system was assumed to be driven by the a-highest possible broadened forcing frequency of each load g
case considered.
This pulls the two frequencies (i.e.,
the forcing frequency and the system frequency) as close together as they can ever possibly be, thereby conserva-E tively maximizing the dynamic load factor (DLF).
The DLF 5
was computed for each S/RV load case considered using the following expression for a harmonic forcing function (see PUAR Equation D.1.2-24):
1 il-Q2 fg )2 + 4 ({Q/ w )2 2
where
(=-dampingratio (2% was used)
O = forcing function frequency w = system natural frequency I
I-FRC 22-lR2 PUAR.04 I
I The load cases considered and corresponding DLFs were as follows:
Case O
UJ No.
Deseription (Hz)
(Hz)
DLF 1
NOC,DBA-Ist Actuation 8.34 18.25 1.26
?
NOC,DBA-2nd Actuation 10.29 18.25 1.46 3
SBA,lBA-Ist Actuation 12.31 18.25 1.84 4
SBA,IBA-2nd Actuation 14.69 18.25 2.83 Finally, BFN S/RV test results showed significant conservatism of S/RV discharge line and support stresses rela tive to analytically predicted values.
See PUAR Sections C.7.1 and C.7.2.
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I ITEM 23:
With respect to Section 8.2.2.3 of the PUAR (5), provide and justify the reasons for not considering the contributions of higher modes above 20 Hz for seismic analysis of torus attached piping systems.
RESPONSE
Seismic analysis of torus-a t tached piping sys tems was performed using the original analysis methodology as g
permi t ted by Sec t ion 4.4.1 of NUREG 0661.
Original seismic g
piping analysis methodology of BFN documented in FSAR Appendix C.3.2.1.a, includes use of 20 Hz as the " cut-of f" frequency.
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I-FRC 23-lR2 PUAR.04 l
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I ITEM 24:
With respect to Section 8.2.5.2 of the PUAR (5), provide justification for considering branch lines having peak spectral accelerations below 5.0 g at the point of attachment to the process line to be qualified without I
further evaluation.
RESPONSE
TVA's eriterla for excluding branch 1ines from additicnal analysis may have been misinterpreted.
The exclusion limit I
is not the acceleration input to the branch line from the process 1ine.
It is the ampIifled motion of the process
- line, i.e.,
the exclusion limit is based on the dynamic I
reponse spectra for the branch line.
The 5-g limit was originally selected based on TVA's experience with seismic qualification of small lines with I
typical BFN configurations.
Experience with DFN LTP analysis of branch lines which exceeded the 5-g limit provided further verification of the acceptability of this limit for BFN branch lines.
ADDENDUM During the September 13, 1984 meeting, FRC representatives indicated some concern with this response and during a
- I telecon on September 24, 1984, additional information was requested.
That information follows:
^
g All branch lines which connect.to the torus attached piping g
process lines were evaluated for dynamic response of the branch line, thermal and dynamic displacement of the attached process line, and sustained loads (deadweight and pressure).
Referring to paragraph NC3650 of the 1977 ASME Seetion III Code (PUAR Reference 23), branch 1ine dynamic response stresses plus sustained load stresses are included g
in code equation 9, whereas branch line stresses due to g
process line displacements and sustained loads are included in equation 11.
The 5-g limit for branch line dynamic response analysis was speci fied on the basis of experience as described above.
The adequacy of..that limit and the fact that equation 11 I
I l-PRC 24-m2 rear.04
I stresses are typically more critical for BFN branch lines than equation 9 stresses is shown by Table FRC-24-1.
That tabula tion gives results for ten BFN Unit 3 branch lines which had response spectra exceeding the 5-g limit.
3 Stresses are presented as a percentage of the code equation g
9 and 11 allowables.
The peak of response spectra accelerations and branch line identifiers are also given.
Recognizing that the equation 11 stresses were more critical, all branch lines.were analyzed for these conditions.
It was not necessary or cost ef fective to g
rigorously analyze all branch lines for equation 9 g
stresses--hence the 5-g limit.
Finally, the small compact valves which are located in BFN I
branch lines are structurally adequate for accelerations much greater than 5-g's.
Therefore, the 5-g limit is also appropr ia te from a component operabili ty standpoint.
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TABLE H C-24-1 TYPICAL HWOI LINE ANALYSIS RESL."-
Process Line Process Line Fraction of Allowable Stress Peak Spectral Penetration Node Point Equation 11 Equation 9 Acceleration, g's X212-37
.3
.3 9.5 30
.9
.1 6.1 40
.1
.1 9.3 46
.1
.1 6.3 107_
.4
.3 9.9 X214 55
.4
.1 9.5 I
X223B E30
.2
.1 10.8 X231 55
.6
.4 9.9 75X
.2
.1 20.1 75Z
.5
.2 20.1 ll 6
ll PUAR.04 ll
ITEM 25:
With respect to Section 8.2.5.5 of the PUAR (5), provide justification for considering the valves with accelerations less than 3-g horizontal and 2-g vertical and having no operator supports to be quali fied without further evaluation.
RESPONSE
The 3-g horizontal and 2-g vert ical accelera tion limits on valve accelerations are justified by our experience with seismic quali fication of similar valves on four TVA nuclear plants (Browns Ferry, Sequoyah, W'a t ts Bar, and Belle fon te).
In addition, none of the numerous valves which were evaluated for the BFN LTP had any problem with sa tis fying the requirements of PUAR Section 4.3.3 with applied accelerations in excess of the 3 g/2-g limits.
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I-FRC 25-lR2 PUAR.04
ITEM 26:
Provide a schedule for the completion of pipe support I
modifications for Units 2 and 3.
RESPONSE
The BFN Unit I and Unit 3 pipe support modifications are complete, and all Unit 2 Internal pipe support modifications are complete.
An integrated nodification schedule was submitted in August 1984 for NRC review and approval, indicating I
compl e t i on o f Un i t 2 external pipe support modifications during the Cycle 6 refueling outage.
That schedule was subsequently revised to show completion of all Unit 2 I
modifications during the Cycle 5 refueling outage.
Therefore, all BFN LTP pipe support modifications will be installed before restart for Cycle 6 operations in I
accordance with NRC's orders (PUAR Reference 12).
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I-FRC 26-1R2 PUAR.04
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