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| {{#Wiki_filter:- - _ _ _ _ _ _ _ _ - _- _ _ _ __ | | {{#Wiki_filter:}} |
| ~ WESTINGHOUSE CLASS 3
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| -i 3
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| WCAP-12247 I
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| i
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| l 9
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| , EVALUATION OF THERMAL STRATIFICATION l
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| FOR THE COMANCHE PEAK ONIT 1
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| - PRESSURIZER SURGE LINE April, 1989 !
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| 1 i
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| R. L. Brice-Nash R. Krishnan P. L. Strauch i B. J. Coslow T. H. Liu S. A. Swamy l E. L. Cranford B. F. Maurer L. M. Valasek j J
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| K. R. Hsu J. F. Petsche J. C. Schmertz-E. R. Johnson D. H. Roarty C. Y. Yang Verified by: [/b4/[J K. C. Chhrfy '
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| . Verified by I /s . V M F./J. Witt l
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| i Approved by: M '
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| @ Approved by: MMbM -l R. B! Patel, Manager /5. 5/Parusamy, Manager 'l Systems Structural Analysis Structural Materials ';
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| and Development Engineering l
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| l l
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| l WESTINGHOUSE ELECTRIC CORPORATION i
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| Nuclear and Advanced Technology Division 4 P.O. Box 2728 Pittsburgh, Pennsylvania 15230-2728 3eS3e/04I880:10
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| '8905030071 890501 PDR ADOCK 05000445.
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| Q PDC.
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| TZ ' __- ._ _ . _ _ _ _
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| 1 1
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| i The authors are grateful for the assistance of Mr. William E. Klassen for his l valuable assistance in the preparation of this report. ~j e 1 i
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| mumnu,o ;;;
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| 1 l . . _ _ - - - _ _ . _ _ _ _ _ _ _ _ _ _ _ - _ _
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| : c. .
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| 1 1
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| TABLE OF CONTENTS
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| . 1 u
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| Section Title Page j
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| ==SUMMARY==
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| xvii 1
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| ==1.0 INTRODUCTION==
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| AND UPDATE OF DESIGN TRANSIENTS 1-1
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| < 1.1 Introduction 'l-1 j l
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| 1.1.1' System Description 1-1 1.1.2 -Thermal Stratification In The Surge Line 1 -?.
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| 1.1.3 Surge Line Stratification Program 1-3 1.2 Update of Design Transients 1-4 1.2.1 System Design Information 1-4 ;
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| 1.2,? Stratification Effects Criteria 1 1.2.3 Plant Monitoring 1-5 1.2.4 Heat Transfer and Stress Analyses 1-10 1.2.5. Stratification Profiles 1-10 1.2.6 Development of Conservative Nornal and 1-11
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| * ~
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| Upset Transients 1.2.7 Temperature Limitations During Heatup 1-12
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| ~
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| and Cooldown 1.2.8 Historical Data 1-13 1.2.9 Development of Heatup and Cooldown Design 1-13 Transients With Stratification 1.2.9.1 [ ]"'C'' Transients 1-14 1.2.9.2 [ Ja,c.e Transients 1-18
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| ~
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| 1.2.10 Striping T.ansients 1-18 1.3 Conclusions 1-19 2.0 STRESS ANALYSES 2-1 2.1 Piping System Structural Analysis 2-1 l- 2.1.1 Introduction 2-1 2.1.2 Discussion 2-2 3003s @ 1449 10 y
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| TABLE OF CONTENTS (cont.)
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| Section Title Page 2.1.3 Results 2-5 2.1.4 Additional Information on Linear Equivalent 2-5
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| - u Techniques q i
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| 2.1.4.1 Introduction 2-5
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| ~
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| 2.1.4.2 Theory 2-6 2.1.4.3 Application 2-9 1 2.1.4.4 Discussion 2-9 ,
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| 2.1.5 Conclusions 2-9 l' 2.2 Local Stress Due to Non-Linear Thermal Gradient 2-10 2.2.1 Explanation of Local Stress 2-10 l
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| 2.2.2 Superposition of Local and Structural Stresses 2-10 2.2.3 Finite Element Model of Pipe for Local Stress 2-10 2.2.4 Pipe Local Stress Results 2-11 2.2.5 Unit Structural Load Analyses For Pipe 2-12 2.2.6 RCL Hot Leg Nozzle Analysis 2-12 -
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| 2.2.7 Conservatism 2-12 2.3 Thermal Striping 2-13 -
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| 2.3.1 Background 2-13 2.3.2 Additional Background Information 2-13 f
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| 2.3.3 Thermal Striping Stresses 2-16 2.3.4 Summary of Striping Stress Considerations 2-16 2.3.5 Thermal Striping Total Fluctuations and Usage 2-18 Factor -
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| 2.3.6 Conservatism 2-19 3.0 ASME SECTION III FATIGUE USAGE FACTOR EVALUATION 3-1 3.1 Code and Criteria 3-1 j 3.2 Previous Design Methods 3-1 ..l 3.3 Analysis for Thermal Stratification 3-1 l 1
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| l 3.3.1 Stress Input 3-2 . i 3603s/041749 10 yj 1 1 L
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| TABLEOFCONTENTS(cont.)
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| . Section Title- Page, 3.3.2 Classification and Combination of Stresses 3-2
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| -; 3.3.3. Cumulative Fatigue Usage Factor Evaluation 3-3 3.3.4 Simplified Elastic-Plastic Analysis 3-4 a
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| . 3.3.5 Fatigue Usage Results. '3 3.4 Conservatism in Fatigue ~ Usage Calculation 3-5 3.5 References 3-5 l
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| 4.0 FATIGUE CRACK GROWTH '4-1 l 4.1 Introduction 4-1 4.2 Initial Flaw Size 4-2 4.3 Critical Locations for FCG 4-2 i
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| 4.4 Results of_FCG Analysis 4-3 )
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| \
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| 4.5 Refereness- 4-3 j l
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| 5.6 REASSESSMsNTOFLEAK-BEF' ORE-BREAK 5-1 5.1 Introduction 5-1 5.2 Material Properties 5-1 !
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| 5.3 Loading Conditions 5-2 5.4 Leak Rate Calculation 5-2 5.5 Reactor Coolant System (RCS) Cooldown Stratification- 5-3 Temperature Considerations l 5.5.1 Reactor Coolant System Temperatures 5-3 5.5.2 Pipe Versus System Temperature Difference 5-4 ,
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| 5.5.3 Conclusions 5-5 !
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| 5.6 Evaluation of Flux Welds 5-6 5.7 Results 5-6' ;
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| 5.8 References 5 -;
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| l
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| ==6.0 CONCLUSION==
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| S 6-1 APPENDIX A - LIST OF COMPUTER PROGRAMS A-1 1.
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| seawo417ss to yjj
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| L LIST OF TABLES
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| , Table Title Page-1-1 IMPORTANT DIMENSIONLESS GROUPS FOR SIMILITUDE IN 1-20 l HYDRODYNAMIC TESTING I
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| 1-2 STRATIFICATION POTENTIAL BASED ON RICHARDSON NUMBER 1-21
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| , 1-3 .SURGELINE TRANSIENTS WITH STRATIFICATION HEATUP (H) 1-22 -
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| AND C00LDOWN (C) - 200 CYCLES TOTAL 1-4' SURGE LINE TRANSIENTS WITH STRATIFICATION NORMAL AND 1-23 UPSET TRANSIENT LIST 1-5 STRATIFICATION PROFILES 1-25 !
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| 1-6 HEATUP - C00LDOWN TRANSIENTS 1-26 1-7 DESIGN TRANSIENTS WITH STRATIFICATION 1-27 1-8 OPERATIONS SURVEY 1-28 i 1-9 HEATUP DATA
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| ==SUMMARY==
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| (PZR - HOT LEG) TEMP. DIFFERENCE AND 1-29 TIME DURATION FOR EACH-PHASE j 1-10 C00LDOWN DATA
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| ==SUMMARY==
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| (PZR - HOT LEG) TEMP. DIFFERENCE AND 1-30 .
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| TIME DURATION FOR EACH PHASE
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| ~ ~
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| 1-11 TRANSIENT T'YPES 1-31 l 1-12
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| ==SUMMARY==
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| OF FATIGUE CYCLES FROM ( Ja,c.e 1-32 1-13
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| ==SUMMARY==
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| OF PLANT MONITORING HEATUP/COOLDOWN TRANSIENTS 1-33 WITH STRENGTH OF STRATIFICATION.(RSS) j 1-14
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| ==SUMMARY==
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| OF MONITORED TRANSIENT CYCLES (0NE HEATUP) 1 35 i i
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| 1-15
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| ==SUMMARY==
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| OF % TIMES AT MAXIMUM TEMPERATURE POTENTIAL RMTP g 1-36
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| . 1-16 SURGE LINE TRANSIENTS - STRIPING FOR HEATUP (H) AND 1-37 j -
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| C00LDOWN (C) 2-1 COMPARISON OF WECAN AND ANSYS RESULTS FOR LINEAR 2-20 l
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| STRATIFICATION - Case 2 2-2 COMPARISON OF WECAN ( ]a,c.e AND 2-21 ANSYS( )]a,c.e )ESULTS R FOR CASE 3 -l e
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| newwue to ;,
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| i
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| .._)
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| l LIST OF TABLES (cont.). -
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| Table' Title Page- .
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| j 3 TEMPERATURE DISTRIBUTIONS IN' PRESSURIZER SURGE LINE 2 2-4 2-23' THE EQUIVALENT LINEAR COEFFICIENTS JiK 2-5 COMANCHE PEAK UNIT 1 SURGjjlNE MAXIMUM LOCAL AXIAL STRESSES 2-24 j
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| )
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| AT ( .
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| i l
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| 2-6
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| ==SUMMARY==
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| OF LOCAL STRATIFICATION STRESSES IN THE SURGE LINE 2-25 l AT THE RCL N0ZZLE , .)
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| 2-7
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| ==SUMMARY==
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| OF PRESSURE AND BENDING' INDUCED STRESSES IN THE 2-26' l SURGE LINE RCL N0ZZLE FOR UNIT LOAD CASES 2-8 STRIPING FREQUENCY AT 2 MAXIMUM LOCATIONS FROM 15 TEST RUNS 2-27 q
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| 2 FLOW RATES AND RICHARDSON NUMBER FOR WATER MODEL FLOW TESTS 28 ;)
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| 2-10 RESULTS FROM TWO HIGHEST THERMOCOUPLE LOCATIONS 2-29 l l
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| ?,-30 2-11 ASME CODE STRESS
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| ==SUMMARY==
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| )
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| 4-4 l 4-1 FATIGUE CRACK GROWTH RESULTS FOR 10% WALL INITIAL FLAW SIZE 5-1 STEPS IN A LEAK-BEFORE-BREAK ANALYSIS 5-8 q 5-2 LBB CONSERVATISM 5-9 'l ;
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| 5-3 ROOM TLMRAWRE MECHANICAL PROPERTIES OF SURGE LINE 5-10 :<!
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| MATElHALT MD WELDS OF THE COMANCHE PEAK UNIT 1 PLANT l 5-4 TYPICAL TENSILE PROPERTIES OF SA376 TP316 AND WELDS OF' 5-11 l SUCH MATERIAL FOR REACTOR PRIMARY COOLANT SYSTEMS )
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| u 5-5 TENSILE,P@0gERTIES FOR THE SURGE LINE MATERIAL AT . 5-12
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| ( ) AND 653*F l 5-6 TYPES OF LOADINGS 5-13 -
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| l 5-7 NORMAL AND FAULTED LOADING CASES FOR LBB EVALUATIONS 5-14 ,
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| ~
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| 5-8
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| ==SUMMARY==
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| OF LOADS AND STRESSES AT THE CRITICAL' LOCATIONS 5-15 5-9 ASSOCIATED LOAD CASES FOR ANALYSES 5-16 LOAD CASES, LOCATION AND TEMPERATURES CONSIDERED FOR LEAK- 5-17 -
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| 5-10 BEFORE-BREAK EVALUATIONS 5-11 LEAKAGE FLAW SIZES, CRITICAL FLAW SIZES AND MARGINS 5-18 1
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| 5-12 SIGNIFICANT THERMAL TRANSIENTS '5-19 i
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| 3M3s/041He 10 g u-_____-____________
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| LIST OF FIGURES
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| . Figure Title Page j j
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| 1-1. Simplified Diagram of the RCS 1-38 j 1-2 Reactor Coolant System Flow Diagram (Typical Loop) 1-39 l l
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| 1-3 RCS Pressurizer 1-40 l
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| - 1 1-4 Estimate of Flow Stratification Pattern in Elbow Under 1-41 :)
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| Pressurizer I 1-5 Comanche Peak Unit 1 Pressurizer Surge Line Stratification 1-42 l ASME III and Qualification Program 1-6 Transient Development Flow Chart 1-43 ,
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| 1-7 [ ]a,c,e Pressurizer Surge Line Monitoring Locations 1-44 1-8 [ Ja,c.e Pressurizer Surge Line Monitoring Locations 1-45 1-9 [ ]a,c.e Pressurizer Surge Line Monitoring Locations 1-46 1-10 [ ]a,c.e Pressurizer Surge Line Monitoring Locations 1-47 i I
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| 1-11 Comanche Peak Pressurizer Surge Line Monitoring Locations 1-48 j 1-12 Reactor Coolant Pump Cut,-off Transient Location Approximately 1-49 10' From RCL Nozzle Safe-End 1-13 Reactor Coolant Pump Cut-off Transient RCLHL Nozzle Safe-End 1-50 l i
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| 1-14 Transient Typical of RC Pump Cut-off 1-51 i
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| \
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| 1-15 Temperature Profile (6.5-inch ID Pipe) 1-52 j 1-16 Dimensionless Temperature Profile (14.3-inch ID pipe) 1-53 l 1-17 Surge Line Stratification 1-54 1-18 Surge Line Hot-Cold Interface Locations 1-55
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| . 1-19 Typical ( ]a,c.e Temperature Profiles 1-56 l
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| an. wine io x3 j 1
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| I
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| j i
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| )
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| LIST OF FIGURES (cont.)
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| . Figure Title Page ,
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| 1-20 Inadvertent RCS Depressurization (AT = 260*F in Surge Line). 1-57
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| , ,?
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| 1-21 Steam Bubble Mode Heatup 1-58 1-22 Steam Bubble Mode Cooldown 1-59
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| ~
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| 1-23 Heatup [ .Ja,c e 1-60 1-24 Cooldown (- Ja,c.e 1-61 1-25 Heatup ( )"'C'' 1-62 1-26 Cooldown [ Ja,c.e 1-63 1-27 Heatup ( Ja,c,e 1-64 1-28 Cooldown [ ]a,c.e 1-65 1-29 Heatup ( Ja,c.e 1-66 1-30 Cooldown [ la,c,e 1-67 1-31 Heatup [ Ja,c e 1-68 1 Cooldown (- Ja,c.e .
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| 1-69 .
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| 1-33 (- Ja,c,e Location 1 - Heatup (7 Days) 1-70 I 1
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| 1-34 ( -Ja,c,e Location 1 - Heatup (4 Days) 1-71 1-35 [ la.c.e Location 1 Fatigue Cycles - Heatup (11 Days) 1-72 ,
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| 1-36 Thermal Cycle Distribution Assumed For One Heatup Cyc'e . 1-73
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| \
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| 1-37 ( 1-74 1
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| ],,c,,
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| ~
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| 1-38 Indication of Striping Thermal Cycles Assumed For One Heatup 1-75 i l
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| 1-39 Comparison of Design to Monitored Transients 1-76 ,
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| j 1-40 Comparison of Design to Monitored Transients 1-77 2-1 Determination of the Effects of Thermal Stratification 2-31
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| . I 2-2 Stress Analysis 2-32 j 2-3 Typical Pressurizer Surge Line Layout 2-33 l mwwm io xig
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| l LIST OF FIGURES (cont.) j l
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| Figure Title Page-2-4 Cases 1 to 4: Diametric Temperature Profiles 2-34 i 2-5 Case 5: Diametric anc Axial Temperature Profile 2-35 2-6 Finite Element Model of the Pressurizer Surge Line Piping 2-36 I
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| . General View 2-7 Finite Element Model of the Pressurizer Surge Line Piping 2-37 Hot Leg Nozzle Detail 2-8 Thermal Expansion of the Pressurizer Surge Line Under 2-38 !
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| 1 Uniform Temperature 2-9 Ccse 2 (linear) Temperature Profile at Hot Leg Nozzle 2-39 2-10 Case 2 (linear) Temperature Profile at Pressurizer Elbow 2-40 2-11 Thermal Expansion of Pressurizer Surge Line Under Linear 2-41 Temperature Grad'ent j 2-12 Bowing of Beams Subject to Top-to-Bottom Temperature Gradient 2-42 2-13 Case 3 (MidrPlane Step): Temperature Profile at Hot Leg Nozzle 2-43 l 1
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| 2-14. Case 3 (Mid-Plane Step):, Temperature Profile at Pressurizer 2-44 l Nozzle l l
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| 2-15 Case 4 (Top Half Step): Temperature Profile at Hot Leg Nozzle 2-45 2-16 Case 4 (Top Half Step): Temperature Profile at Pressurizer 2-46 l Elbow ]
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| 2-17 Case 5: Axial and Diametric Temperature Profile 2-47 2-18 Case 5: Axial and Diametric Temperature Profile at Hot Leg 2-48 Nozzle 2-19 Case 5: Axial and Diametric Temperature Profile at Pipe Bend 2-49
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| , 2-20 Case 5: Axial and Diametric Temperature Profile at 2-50 Pressurizer (1 bow 2-21 [ ]a,c.e Profile 2-51 2-22 Cemanche Peak Surge Line Model and Temperature Profile 2-52 2-23 Equivalent Linear Temperature 2-53 mwouns io xgjj
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| LIST OF FIGURES (cont.)
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| Figure Title Page 2-24 Local Stress in Piping Due to Thermal Stratification 2-54
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| ~
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| 2-25 Independence of Local and Structural Thermal Stratification 2-55 Stresses Permitting Combination by Superposition 2-26 Test Case for Superposition of Local and Structural Stresses 2-56 ,
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| 2-27 Local Stress - Finite Element Models/ Loading 2-57 2-28 Piping Local Stress Model and Thermal Boundary Conditions 2-58 2-29 Surge Line Temperature Distribution at [ Ja,c.e Axial 2-59 Locations 2-30 Surge Line Local Axial Struss Distribution at [ ]a,c.e 2-60 Axial Locations 2-31 Surge (ige
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| * Local Axial Stress on Inside Surface at 2-61
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| [ ] Axial Locations 2-32 Surge gige* Local Axial Stress on Outside Surface at 2-62
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| [ ] Axial Locations 2-33 Surge Line Temperature Distribution at Location ( Ja,c.e 2-63 2-34 SurgeLineLoga}'gxialS'tressDistributionat 2-64 Location [ ]
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| 2-35 Surge Line Temperature Distribution at Location [ ]a,c.e 2-65 2-36 SurgeLineLoga}'gxialStressDistributionat 2-66 Location [ ]
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| 2-37 Surge Line Temperature Distribution at Location [ ]a,c.e 2-67 2-38 Surge Line Loga},gxial Stress Distribution at 2-68 l
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| Location [ ] .
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| l 2-39 Surge Line Temperature Distribution at Location [ ]"'C'8 2-69 2-40 SurgeLineLoga}'gxialStressDistributionat 2-70 Location [ ] i l 2-41 Surge Line Temperature Distribution at Location [ ]a,c.e 2-71 1
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| 2-42 SurgeLineLoga}'gxialStressDistributionat 2-72 Location [ ]
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| 2-43 Surge Line RCL Nozzle 3-D WECAN Model #1 2-73 newww. m x$y l
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| l
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| l 1
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| LIST OF. FIGURES (cont.)
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| . Figure Title Page i i
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| 2-44 Surge Line RCL' Nozzle 3-D WECAN Model #2 2-74 j 2-45 Surge Line Nozzle Temperature Profile Due to Tnermal 2-75 Stratification -!
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| a
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| . 2-46 Surge Line Nozzle Stress Intensity Due to Thermal 2-76 !
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| Stratification >
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| , 2-47 Surge Line Nozzle Stress in Direction Axial to Surge Line Due 2-77 to Thermal Stratification 2-48 Surge Line Nozzle Stress Intensity Due to Pressure 2-78 2-49 Surge Line Nozzle Stress Intensity Due to Pressure 2-79 2-50 Surge Line Nozzle Stress Intensity Due to Bending 2-80 j 2-51 Surge Line Nozzle Stress in' Direction Axial to Surge Line 2-81 Due to Bending Showing Magnified Displacement-2-52 Surge Line Nozzle Stress Intensity Due to Bending Showing 2-82 Magnified Displacement 1 2-53 Surge Line Nozzle Stress Intensity Due to Bending 2-83 2-54 Thermal Striping Fluctuation 2-84 2-55 Stratification and Striping Test Models 2-85 2-56 Water Model of 'MFBR Primary Hot Leg 2-86 2-57 Attenuation of Thermal Striping Potential by Mglgeglar 2-87 l Conduction (Interface Wave Height of ( ) q 2-58 Thermal Striping Temperature Distribution 2-88
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| - 2-59 Striping Finite Element Model 2-89 4-1 Determination of the Effects of Thermal Stratification 4-5 ,
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| , on Fatigue Crack Growth ,
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| i 4-2 Fatigue Crack Growth Methodology 4-6 ;
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| 4-3 Fatigue Crack Growth Rate Curve for Austenitic Stainless Steel 4-7 4-4 Fatigue Crack Growth Equation for Austenitic Stainless Steel 4-8 4-5 Fatigue Crack Growth Critical Locations 4-9 mwume io !
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| xy l
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| LIST OF FIGURES (cont.)
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| Figure- Title Page ,
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| 4-6 Fatigue Crack Growth Controlling Positions at Each Location 4-9 5-1 Average True Stress-True Strain. Curve for the Surge Line 5-21 SA376 TP316 Stainless Steel at 653'F 5-2 Minimum True Stress-True Strain Curve for the Surge Line 5-22 .'
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| SA376 TP316 Stainless Steel'at 653'F 5-3 Average True Stress-True Strain' Curve 15'ihe Surge Line 5-23 '
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| SA376 TP316 Stainless Steel' at [ ]
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| 5 -Sketch of Analysis Model for Comanche Peak Unit 1 Pressurizer 5-24 Surge Line Showing Node Points, Critical Locations,. Weld Locations and Type of Welds 5-5 RCS Cooldown 5-25 5-6 -[ la,c,o Actual Cooldown Due To An RCS Leak 5-26 5-7 [ Ja,c.e Actual Cooldown 5-27 5-8 '( Ja,c.e Location 1 Cooldown System AT and Pipe 5-28 AT vs. Time-e i
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| )
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| l j
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| i i
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| i as w osteesio gyj t _
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| ==SUMMARY==
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| This report presents the methods, data, analysis and qualification r.esults for the Comanche Peak Unit 1 pressurizer-surge lines including thermal stratification. l
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| . l The report is divided into six sections. Appendix A is a list of computer
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| . codes.used in this work. The sections are presented in order, reflecting the logical progression of evaluations and analyses:
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| o Section 1.0 " Introduction and Update of Design Transients" presents the methods and data used to update the design thermal transients to incorporate'the effects of flow stratification in the surge line.
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| o Section 2.0 " Stress Analysis" describes the global and local stress effe. cts of stratification, including striping.
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| ~
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| o Section'3.'0 "ASME III Tatigue Usage Factor Evaluation" provides the evaluation results of the ASME III fatigue life.of the surge line subject to all design transients plus _ the effects of stratification.
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| o Section 4.0 " Fatigue Crack Growth" describes the methods and .
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| results of fatigue crack growth predictions in the surge line subject to stratification. .
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| o Section 5.0 " Leak Before Break" (LBB) is a reassessment of the LBB evaluations to account for the effects of stratification.
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| o Section 6.0 " Conclusions" summarizes the results of the )
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| ~
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| evaluations of the effects of stratification in the surge line.
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| . . . . , xyn J
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| o Appendix A " Computer Codes" is a list and description of computer.
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| codes used in this work.
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| The work presented in this report leads to the following conclusions:
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| (a) Based on plant monitoring results from (five]a,c,e Westinghouse PWR's and flow stratification test data, the thermal design ,
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| transients for the surge line have been updated to incorporate the-effects of stratification. .
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| (b) The global structural and local stresses and loads in the surge line piping and support system meet ASME III Code allowables. The maximum cumulative fatigue usage factor is [0.74]a,c.e for 40 year design life, compared to the Code allowable of 1.0.
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| (c) Fatigue crack growth (FCG) analyses show that a postulated 10%
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| initial crack will not propagate beyond 20% of the pipe wall in 40 years design life.
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| (d) Leak-Befor~e-Break (LBB) is confirmed for all loading combinations, -
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| including maximum postulated stratification, using the methods of
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| ~
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| SRP 3.6.3 and NUREG 1061, Vol. 3.
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| In summary, based on the current understanding of the thermal stratification phenomenon, it is concluded that thermal stratification has very limited impact on . cegrity of the pressurizer surge line of the Comanche Peak Unit i nuclear power plant. The forty year design life is not impacted.
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| e mwo im io xyjjj
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| SECTION
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| ==1.0 INTRODUCTION==
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| AND UPDATE OF DESIGN TRANSIENTS I
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| 1.1 Introduction _l 1.1.1 System Description V I
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| The primary function of the reactor coolant system (RCS) is to transport heat '
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| , from the reactor core to the steam generators for the production mf steam. j The Comanche Peak Unit 1 RCS consists of four similar heat transf, loops connected to the reactor vessel (figure'l-1). Each loop contains a reactor coolant pump (RCP) and a steam generator. The system also includes a pressurizer, connecting piping, pressurizer safety and relief valves, and a-relief tank.
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| The flow path for a typical reactor coolant loop is from the reactor vessel to the inlet plenum of the steam generator (figure 1-2). High temperature reactor coolant flows through the,U-tubes in the steam generator, transferring heat to the secondary water, out of the tubes into the outlet plenum to the suction of the reactor coolant pump. The reactor coolant pump increases the pressure head of the reactor coolant which flows back to the reactor vessel.
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| The pressurizer vessel (figure 1-3) contains steam and water at saturated conditions with the steam-water interface level between 25 and 60% of the ;
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| vc,1ume depending on the plant operating conditions. From the time the steam bubble is initially drawn during the heatup operation'to hot standby conditions, the level is maintained at approximately 25%. During power ascension, the level is increased to approximately 60%.
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| As illustrated in figure 1-2, the bottom of the pressurizer vessel is connected to the hot leg of one of the coolant loops by the surge line, a 14 inch schedule 160 stainless steel pipe with a 1 foot section of 14 inch- !
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| schedule 140 pipe.
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| i l
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| l 1
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| = = . . . 13
| |
| | |
| The simplified diagram shown in figure 1-2 indicates the auxiliary systems that interface with the RCS. Of particular significance to surge line
| |
| ')
| |
| stratification are the normal charging and letdown function provided by the Chemical and Volume Control System (CVCS), and the suction and return lines associated with the Residual Heat Removal System (RHRS). The former directly controls the RCS mass inventory and therefore affects flow in the surge line. ,
| |
| The RHRS is used to remove heat from the RCS and thereby influences coolant I temperature and consequently coolant volume through thermal expansion and ,
| |
| contraction.
| |
| Other systems which affect surge line flow conditions are main spray flow supplied to the pressurizer from one or two cold legs and the pressurizer electric heaters. Spray operation does not significantly alter the total RCS mass inventory, but does reduce system pressure by condensing some of the steam in the pressurizer. The pressurizer heaters when energized generate steam and as a result increase RCS pressure.
| |
| 1.1.2 Thermal Stratification In the Surge Line Thermal stratification in the pressurizer surge line is the direct result of-the difference in densities between the pr'essurizer water and the generally cooler het leg water. The lighter pressurizer water tends to float on the cooler heavier. hot leg water. The potential for stratification is increased as the difference in temperature between the pressurizer and the hot leg increases and as the insurge or outsurge flow rates decrease.
| |
| At power, when the difference in temperature between pressurizer and hot leg -
| |
| is relatively small (less than 50*F) the extent and effects of stratification have been observed to be small. However, during certain modes of plant heatup
| |
| * and cooldown, this difference in system temperature could be as large as 320'F, in which case the effects of stratification must be accounted for.
| |
| A common approach for assessing the potential for stratification is to evaluate the Richardson Number (tables 1-1 and 1-2) which is the ratio of the -
| |
| thermal density head diametrically across the pipe to the fluid flow dynamic !
| |
| head, or mmesee in 12 ;
| |
| | |
| Ri = 98Da72 U
| |
| where Ri = Richardson number g = gravitation constant U = hot fluid velocity AT = hot-to-cold fluid temperature difference D = pipe inside diameter 6 = coefficient of thermal expansion of water For a range of surge line flow rates from approximately 700 gpm down to a bypass flow of approximately 1 to 5 gpm and AT = 320"F, the Richardson number is greater than the value of 1 which is required to initiate st-ratification. Thus under this range of conditions, the flow has. the potential to be stratified due to the relatively large hot-to-cold fluid temperature difference combined with the low hot fluid velocity. To eliminate stratification (i.e., Ri smaller than 1) a flow velocity of over 2.4 fps (approximately 700 gpm) is needed (figure 1-4).
| |
| 1.1.3 Surge Line Stratification Program The surge line stratification program for Comanche Peak Unit 1 consists of three major parts:
| |
| (a) Jpdate of design transients (b) ASME III stress, fatigue cumulative usage factor (CUF), fatigue crack growth (FCG) and leak-before-break (LBB) analyses (c) Confirmatory monitoring.
| |
| I newom.. io i.3
| |
| | |
| Figure 1-5 shows the steps required to complete this program.
| |
| 1.2 Update of Design Transients ,
| |
| 9 The method used to update the desigg transients for stratification is illustrated'in figure 1-6 and is didussed in this section. .
| |
| 1.2.1 System Design Information ,
| |
| The thermal design transients for the Comanche Peak Unit 1 Reactor Coolant System, including the pressurizer surge line, are defined in Westinghouse Systems Standard Design Criteria (SSDC) documents SSDC 1.3.
| |
| The design transients for the surge line consist of two major categories:
| |
| I.
| |
| (a) Heatup and Cooldown transients (b) Normal and Upset operation transients. By definition, the emergency
| |
| ~
| |
| and faulted transients are not considered in the ASME III Section NB fatigue life assessment of components.
| |
| . In the evaluation of surge line. stratification, the FSAR chapter 3.9 definition of normal and upset design events and the number of occurrences of the design events remains unchanged.
| |
| l The total nu,nber of current heatup-cooldown cycles (200) remains unchanged. ~
| |
| However, sub-events and the associated number of occurrences (" Label", " Type" j and " Cycle" columns of tables 1-3 and 1-4) are defined to reflect monitoring. q data, as described later. '
| |
| j In all cases, the surge line fluid temperature distribution is modified from I the original uniform temperature to a stratified distribution with the maximum -
| |
| temperature differentials and the associated nominal temperatures (" MAX i AT strat " and " Nominal" columns on tables 1-3 and 1-4).
| |
| assumise in ;
| |
| 14
| |
| | |
| i 1.2.2 Stratification Effects Criteria To determine the normal and upset pipe top-to-bottom temperature difference, "aT strat " (tables 1-3 and 1-4), the following conservatism is introduced.
| |
| For a given event, the AT strat in the pipe will be the difference betwwen
| |
| ! the maximum pressurizer temperature and the minimum hot leg temperature, even though they do not occur simultaneously.
| |
| [
| |
| t
| |
| ),jac.e 1.2.3 Plant Monitoring Surge line stratification data have been obtained from ( Ja,c.e Westinghouse plants. Figures 1-7 through 1-10 show the instrumentation configuration for four of these plants. The data was obtained by continuous monitoring of the piping OD temperature, displacements and plant parameters.
| |
| . The pipe temperatures were obtained from RTD's located on the outside of surge line. Plant parameters were obtained from the plant computer. Figure 1-11 represents the Comanche Peak Unit 1 monitoring configuration, i
| |
| Temperature data from the other reference plant surge lines were reviewed.
| |
| . The data, in all cases, shows the presence of stratification in the surge lines. The stratification observed is assumed to behave under the influence of gravity and consequently will have an axial profile defined by the slope of 1-5
| |
| | |
| the pipe. The data interpretation herein is an attempt to classify and -
| |
| characterize observed thermal conditions.
| |
| There are two basic causes of thermal stratification. Thermal stratification '
| |
| can be initiated either by [ Ja.c.e or the [^
| |
| Ja c.e This is the condition which this report tddresses.
| |
| [
| |
| m D
| |
| b e
| |
| ga.C,e .
| |
| l w .,s m io 16 L
| |
| | |
| \
| |
| [
| |
| I ja.c.e The establishment of a highly stratified condition is best described by considering the following typical transient example. This transient is based on an observed reference plant transient which was caused by the cut-off of the RCP in the,same loop as-the surge line.
| |
| . Typical Transient
| |
| | |
| == Description:==
| |
| (RCP Cutoff figure 1-12)
| |
| [
| |
| 4 w
| |
| Ja,C,0 ma,m,ue so 1-7
| |
| | |
| l l
| |
| [
| |
| . n 1
| |
| (
| |
| i
| |
| .ja,c,e i
| |
| One-interpretation of the cause and affects of the transient just described is as follows:-
| |
| [
| |
| 9 i
| |
| e l
| |
| a
| |
| }a.C,e , .
| |
| I as musmie 1-8
| |
| | |
| i l
| |
| [
| |
| l 1
| |
| 1 1
| |
| I
| |
| - 1 I
| |
| l
| |
| , ja,c.e l
| |
| l The data are sufficient to characterize. stratification temperatures in the
| |
| . pipe during critical operating transients and heatup-cooldown operation.
| |
| Also, the data are sufficient to verify that the pipe movements are consistent I with analytical predictions, within an accuracy _normally expected from hot functional and/or power ascension tests, as discussed in section 2.1.
| |
| The monitoring of plant parameters is sufficient to correlate measured l ,
| |
| temperature fluctuations to changes in operation. In particular, it is I
| |
| l apparent that temperature fluctuations are due to flow insurge (into the l
| |
| pressurizer) and outsurge (out of the pressurizer) which in turn are due to differential pressure in the system. While a simple quantitative mechanistic relationship between plant operation and insurge and outsurge has not been found, the data indicate that a steady state stratified condition can be altered by any of the fo1~1owing events:
| |
| 1*
| |
| seaweisse ie 19 j L _-
| |
| | |
| l i
| |
| a) Expansion of the pressurizer bubble i 1
| |
| b) RCP trip in the surge line loop .
| |
| c) Safety injection d) Large charging - letdown mismatch ,
| |
| e) Large spray rates in light of these observations, the update of design transients is based on .
| |
| plant monitoring results, operational experience and plant operational procedures. Conservatism have been incorporated throughout the process in ,
| |
| the definition of transients (cycles, AT) and in the analysis, as described in the report. Confirmatory monitoring during the hot functional testing will be employed at Comanche Peak Unit 1 to verify the acceptability of the design transients used in this report.
| |
| 1 1.2.4 Heat Transfer and Stress Analyses l
| |
| l The correlation of measured pipe 00 temperature to ID temperature distribution 1
| |
| ' is achieved by heat transfer analysis as well as previous experience with flow l at large Richardson numbers (Ri 1) (figures i-15 and 1-16).
| |
| 1 l
| |
| These analyses and test data available to date show that a stratified flow condition,[ tr Ja,c.e is a proper and conservative depiction of the flow condition inside the pipe at large AT and low flow rates (Ri>1). f An additional conclusion from the heht transfer and stress analyses is that
| |
| (
| |
| i ja,c.e ,
| |
| i 1.2.5 Stratification Profiles Table 1-5 summarks the major stratification profile characteristics. The monitored data shows a consistent axial temperature profile along the -
| |
| horizontal portions of the [ ]a,c.e surge lines monitored.
| |
| an,* no io 1-10
| |
| | |
| The axial temperature profile is a function of the geometric characteristics
| |
| - of each line. Each line monitored showed a definite relationship between axial length of stratification and slope of the line. Figure 1-17 depicts a typical axial stratification profile. Note that the actual length of stratification is dependent on the volume of the insurge. Low volume insurges tend to stratify a shorter distance along the line. Similarly large volume insurges stratify longer distances provided the slope of the line is low enough. As the slope increases, smaller sections of the line will be affected
| |
| ~
| |
| by stratification. The slope also affects the type of stratification interface. As the slope is increased the flow characteristics of the interface are affected. There are two basic interface types; one which is i narrow and highly defined is characteristic of laminar flow. The other is l characteristically wide and a product of turbulent flow. The flow becomes turbulent at the interface when forced to a higher level than gravity would normally dictate. Flow velocity is also an integral part of this relationship. ,
| |
| Figure 1-18 shows a cross section of the pipe with the various hot and cold fluid interface levels created by a laminar flow or static steady state conditions. .
| |
| 1.2.6 Development of Conservative Normal and Upset Transients Transients in the surge line were characterized as either due to insurges or outsurges (1/0) from the pressurizer or fluctuations. Insurges and outsurges are the more severe transients and result in the greatest change in tempera- i ture in the top or bottom of the pipe. An insurge may cool the bottom of the i pipe significantly, to very close to the temperature of the RCS hot leg. !
| |
| l Conversely, an outsurge can sweep the line and heat the pipe to close to the j temperature of the pressurizer. The thermal transients are shown in figure 1-19.
| |
| Fluctuations, as opposed to the insurge-outsurge transients, are caused by j relatively insignificant surges and result in variations in the hot-cold l interface level. These variations in the interface level do not change the (
| |
| overall global displacement of the pipe and hence are modeled as changes in the depth of the interface zone.
| |
| an.w e in 1 11
| |
| | |
| s WesTINGH3USE PROMffARY class 2 The redefinition of the thermal fluid conditions experienced by the surge line -
| |
| during normal and upset transients was necessary in order to neglect the -
| |
| indirectly observed fluid temperature distributions. These redefined thermal I
| |
| . fluid conditions were developed based on the existing design transient system -
| |
| parameters assumed to exist at the time of the postulated transient and the knowledge gained from the monitoring programs. The redefined thermal fluid
| |
| ~
| |
| conditions conservatively account for the thermal stratification phenomena.
| |
| Several conservatism were introduced in the redefined normal and upset -
| |
| thermal transients (tables 1-3, 1-4, .-6 and 1-7).
| |
| [
| |
| 3a,c.e (b) Full stratification cycles are assumed for all transients, except for steady state fluctuations, unit loading and unloading, and redcred temperature return to power, where level fluctuations are sufficiently conservative based on flow rate and observations.
| |
| (c) The temperature of stratification was based on the minimum het leg temperature at any time during the transiont (for bottom of pipe) and -
| |
| the maximum pressurizer temperature (for top of pipe). Figure 1-20 shows a case where this resulted in a very conservative 260'F stratification transient although the maximon temperature difference at any point in time was about 50*F.
| |
| (d) The current number of design cycles of each event is unchanged. .
| |
| The normal and upset transients modified to account for the stratification ph.nomena are listed in tables 1-3 and 1-4. ;
| |
| i 1.2.7 Temperature Limitations During Heatup and Cooldown The maximum permitted temperature difference between the pressurizer and the ,
| |
| het leg for Comanche Peak Unit 1 is 320'F. Therefore the maximum possible top-to-bottom temperature stratification is 320*F.
| |
| ammwo 1 12
| |
| | |
| i e
| |
| With the RCL cold, the pressurizer pressure (and therefore temperature) is limited by the cold overpressure mitigation system (COMS). 4 Practically, plants operate to minimize downtime and heatup-cooldown time, when power is not being generated. The times at large AT are therefore i
| |
| reasonably limited, as discussed later.
| |
| 1.2.8 Historical Data S'ince not all heatup and cooldown parameters affecting stratification are l formally limited by Technical Specification or Administrative controls, it is necessary to reconsider plant operational procedures and heatup-cooldown practices to update the original heatup and cooldown design transient curves of SSDC 1.3 (figures 1-21 and 1-22).
| |
| To this end, a review of procedures, operational data, operator experience, and historical records was conducted for [ .)a,c.e Westinghouse PWR plants (table 1-8). Similarly, operations procedures for Comanche Peak Unit 1 were .
| |
| , reviewed and heatup cooldown curves developed (shown in figures 1-23 and 1-24).
| |
| . The heatup and cooldown operations information acquired from this review is summarized in tables 1-9 and 1-10, ( .-
| |
| ja.c.e I
| |
| The information is divided into heatup and cooldown tables and diagrams. The diagram presents the pressurizer water and het leg temperature profiles versus
| |
| , time. The various phases of the process are identified by letters along the diagrams' abscissa and in tables 1-9 and 1-10.
| |
| 1.2.9 Development of Heatup and Cooldown Design Transients With Stratification
| |
| ~
| |
| As described above, the database of information used to update the heatup and cooldown transients included the following:
| |
| a) Typical heatup and cooldown curves, as develcped from review of procedures, operational data and operators experience.
| |
| ==>..n nu w 1 33
| |
| | |
| i S
| |
| b) Transients as monitored at ( Ja.c.e plants c) Historical records of critical heatup and cooldown temperatures The heatup and cooldown transients are presented in the following sections as
| |
| [ la.c.e and in similar fashion to the normal and upset transients. Table 1-11 gives the. '
| |
| general characteristics of the two types of transients observed.
| |
| The heatup cooldown transient labels have the following logic:
| |
| : 1. Transients H1 through H12 correspond to insurge or outsurge transients' postulated during heatups (H).
| |
| : 2. Transients HFIA through HF3 correspond to fluctuation transients postulated during heatups (HF). .
| |
| : 3. Transients C1 through C9 correspond to insurge or outsurge transients postulated.during cooldown (C). .
| |
| : 4. Transient CF1 represents the fluctuation transients postulated for .
| |
| cooldowns(CF).
| |
| 1.2.9.1 [ Ja.c.e Transients A) Monitoring Transient Summary i For a given monitored location, plots of temperature difference versus time were generated (figures 1-33 and 1-34 are examples of relatively high ,
| |
| transientactivity). Two parameters were plotted, the pipe top to bottom temperature difference (labeled " surge line") and the pressurizer to hot leg temperature difference (labeled " system").
| |
| It is clear from figures 1-33 and 1-34 that for the observed heatups, ( '
| |
| ),a.c.e muwm io 1 14
| |
| | |
| (
| |
| for conservatism, the envelope from measured transients in all plants is '
| |
| . applied to define the transients.
| |
| B) F_atigue a Cycles The fatigue cycles were obtained using the technique illustrated on figure 1-35, (which reduces each transient to its. temperature range, rather than its absolute magnitude. Table 1-12 is a summary of the approximate magnitude of each cycle shown in figure 1-35. The heatup and cooldown design transients with stratification used in the piping qualification conservatively accounts for the mean stress effect. Figure 1-36 provides a single heatup interpretation of the design transients. Note the relative severity of'the design transients by comparing them with the actual thermal activity observed at the worst case plant (figure 1-33). [
| |
| j Ja,c.e Figure 1-37 illustrates the difference between the design transients and the transients observed at plant A.. .
| |
| C) _ Strength of Stratification Plant monitoring data indicate that for the various transients observed the AT in the pipe (top to bottom) is not as large as the AT in the system (pressurizer to hot leg). The ratio of AT in the pipe to AT in the system will be referred to as " strength of stratification". i 9
| |
| ja.c.e menwme io 1-15
| |
| | |
| i D) Number of Stratification Cycles (table 1-14) - -
| |
| P1' ant monitoring ~ data indicated the significant events which could occur -
| |
| during a given heatup.
| |
| [. !
| |
| ~
| |
| . Ja,c.e E) Maximum Temperature Potential The key factor in thermal stratification of the surge line is the temperature difference between the pressurizer and hot leg (section 1.2). This tempera- .
| |
| ture difference is clearly maximized during the heatup and cooldown, when the plant is in mode 5 cold shutdown (hot leg less than 200*F) and the pressurizer. ,
| |
| bubble has been drawn with the reactor coolant pump running (pressurizer temperature larger than 425'F). [
| |
| ja.c.e sen.wim in 1-16
| |
| | |
| 1 F) Final Cycles and Stratification Rances l
| |
| r 1
| |
| I l
| |
| 1 i
| |
| . 4 1
| |
| 3Ce 8
| |
| Meh m inein 1-17
| |
| | |
| 1 l
| |
| Example:
| |
| 1 I
| |
| i ja c.e G) Cooldown Transients
| |
| ~
| |
| {
| |
| 1 The procedure used in heatup is applied to develop transients for plant i
| |
| {
| |
| l cooldown. (
| |
| l i
| |
| ja,c.e 1.2.9.2 ( - Ja,c.e Transients .
| |
| [ -)
| |
| i 4
| |
| ja c.e ,
| |
| 1.2.10 Striping Transients ,
| |
| Maan stress effects are included in determining the usage factor contributed by thermal striping. Fatigue cycles like those shown in figure 1-35 were v.ot ,
| |
| used in the development of the striping design transients. [
| |
| gac.e asemmases,o -
| |
| 1-18 n___ ___ . . .
| |
| | |
| [ la.c.e It should be
| |
| . noted that each striping transient cycle is assumed to initiate a discrete hot '
| |
| to cold fluid interface that will be attenuated with time (see section 2.3 for-
| |
| . discussion). Figure 1-38 shows the relative magnitude and frequency of the i striping transients for one.heatup or cooldown with respect to the system AT j (PRZT - RCST). The highest pipe AT (pipe T Top pipe Tbot) observed "
| |
| during heatup never exceeded [ ).a.c.e However, the design striping transients consider [ Ja,c.e transients at pipe AT's greater than
| |
| [ 3a,c.e Striping transients use the labels HST and CST denoting striping transients (ST). [
| |
| 3a,c.e 1.3 Conclusions
| |
| ~
| |
| Design transients were updated to incorporate stratification. The transients were developed to conservatively represent the cyclic effects of l
| |
| , stratification. To illustrate the margin included in the development of I heatup transients, a simplified fatigue factor calculation is provided in figures 1-39 and'l-40. This comparison indicates that the design transients have a factor of conservatism of approximately [ j 3a,c.e ;
| |
| i i
| |
| u 3Mases130010 gg
| |
| | |
| I TABLE 1-1 1
| |
| IMPORTANT OIMENSIONLESS GROUPS FOR SIMiuTUDE IN HYDRODYNAMIC TESTING
| |
| ~
| |
| Peremeter Symeed Oettnemen Sognelsenes
| |
| ~
| |
| wegoace vcten / Da# 2ev't Pwe sorcecome torce tactor 2 Cavaanon numeer a 17. - 8, p a vd . Dessure eMerencerene
| |
| *me ,
| |
| 3 meynees numoer Ao avD a meme fortevecous *te 4 Sueuns numoer Sr .io V vones sneeeng t'ensey a eems forte 5 wooer nurnoer we oCV' a meme forte mence gesen mes a
| |
| 4 Peues numoer Fr V 90 ineres 8ertogreret 4:see j r menneesen numoer m. a,gorev8 Suoveey woeemanforce l iMeeses Fmuss nunger) 8
| |
| : 4. Ewer numoor Eu AP'aV Prosaure Wcererte force 9 Prenes rnareer Pr 4.* Memorewn apuenrayitnormal e#ues,ey 10 Peaes numeer Pe pVOC.* Conventare near venser-(Ao a Pr) conoucano nem transer 11 Graeful nwreer Gr da 8gdaT/m' Guoyancy testerveseus norte
| |
| : 12. Maytogn nutreer Rei datgAan Ar -
| |
| (Cr a Pr)
| |
| NOMOeCLATUME; C = sesenc nest g = aanseersten of gravey a = oeneer P = preneure a = surteen toneen 7, = seenflue preenwo e = inerms tenasevey 7, = Ause vapor prenewe A = voummers essenesen assacent 1.0 = caerectorees emeneens AT = flusi tomooresare enange V = m.e m '
| |
| r sonen Sneceng % a = veseesy O
| |
| 6 e
| |
| un..mus to 1-20
| |
| | |
| TABLE 1-2 STRATIFICATION POTENTIAL BASED ON RICHARDSON NUMBER
| |
| * Stratification potential exists if Ri > 1
| |
| . 8,C,e e
| |
| 4 1
| |
| 1 i
| |
| nu.-os un 'a
| |
| | |
| l j
| |
| TABLE 1-3 SURGELINE TRANS!ENTS WITH STRATIFICATION ,
| |
| HEATUP (H) AND C00LDOWN (C) - 200 PLANT CYCLES TOTAL l
| |
| ')
| |
| i 3.C e , (
| |
| 4 I
| |
| l l
| |
| l i
| |
| 1 i1 i
| |
| l l
| |
| l 1
| |
| l 1
| |
| .
| |
| * l N
| |
| 1
| |
| )
| |
| l
| |
| ' I l l l
| |
| l ;
| |
| l 1 m io 1-22 i
| |
| i l
| |
| n
| |
| | |
| i TABLE'1-4:
| |
| SURGE L2NE TRANS2ENTS W2TH STRAT2FICAT20N .
| |
| i -NORMAL AND UPSET TRANSIENT LIST a,C,e w'
| |
| e I
| |
| 9 e
| |
| 4 0
| |
| =
| |
| l..
| |
| l w.nuese in 1-23
| |
| | |
| i TABLE 1-4 (Cont'c.) I SURGE LINE TRANS2ENTS CTTH STRAT2FICAT80N )
| |
| NORMAL AND UPSET TRANSIENT L1f7 a,Cse O
| |
| O
| |
| ~ 0 4
| |
| O 1
| |
| e e
| |
| - 4 e
| |
| m.mnow in 1-24 1
| |
| .m
| |
| | |
| 1 l
| |
| l TABLE 1-5 STRATIFICATION PROFILES l
| |
| . l
| |
| [
| |
| 6 1
| |
| ja.c.e i
| |
| O we.mim to 1-25
| |
| | |
| I I
| |
| TABL E 1-6 HEATUP - CDOLDOWN TRANSIENTS o -. Transients Were Dev01oped Based On: a Typical Heatup Cooldown Curves Envelope (Plus Margin) of Events (Transients) Monitored
| |
| . Historical Data on Temperature Plateaus -
| |
| [
| |
| 1
| |
| -1 0 ,
| |
| o i
| |
| o 1
| |
| l 0
| |
| ya,C,8
| |
| ~
| |
| 9 i
| |
| e sens.mtess to
| |
| . -26 ,
| |
| 5
| |
| | |
| l TABLh. ', 7 DE3IGN-TRANSIENTS WITH STRATIFICATION
| |
| . o Heatup and Cooldown Combined With Other Events o Design Transient Criteria
| |
| [
| |
| l l
| |
| ja,c,e o Input for Local and Structural Analysis Defined - Plus Nozzle o 5triping Transients Defined to Consider Maximum Stratification Cycles Regardless of Range
| |
| * l 1
| |
| 1 L~
| |
| i so minein 1-27 l
| |
| | |
| l
| |
| -TA9LE 1-8 .i 1
| |
| OPERATIONS SURVEY l \
| |
| o- Summary of Plants Surveyed ,
| |
| NO. OF YEARS OF OPfRATION PLANT LOOPS (MAXIMUM) i -
| |
| l ja,c.e l- .
| |
| ; o Reviewed Typical Heatup Cooldown Process o Reviewed Administrative / Tech Spec Limitations
| |
| 'o Reviewed Historical Events and Time Durations i o Developed Heatup - Cooldown Profiles 4 l .. ,
| |
| 1 se maiseein 1-28
| |
| | |
| . . \
| |
| ! i
| |
| . ll1lll ii e.
| |
| c.
| |
| a t
| |
| G E
| |
| S A
| |
| H P
| |
| H C
| |
| A E
| |
| R O
| |
| F N
| |
| O I
| |
| T A
| |
| R U
| |
| D e Y
| |
| R ER A I 0
| |
| 9 U mT l
| |
| I
| |
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| TABLE 1-11 TRANSIENT TYPES
| |
| [
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| ja.c.o O
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| O 3064e-03154910 1-31
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| i 1
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| TABLE 1-12
| |
| | |
| ==SUMMARY==
| |
| OF FATIGUE .CYG.ES FROM [ Ja,c.e Cycle Delta Range (*F) Cycle Delta Range (*F).
| |
| l
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| -- a,c,e s
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| NOTE: The delta range represents the relative severity (AT) of each transient following the fatigue cycle approach.
| |
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| : w. ovies to 1-32 i l
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| 1 TA1LE 1-13
| |
| | |
| ==SUMMARY==
| |
| OF PLANT MONITORING HEATUP/COOLDOWN TRANSIENTS WITH STRENGTH OF STRATIFICATION (RSS)
| |
| [ 3a,c.e
| |
| [ 3a,c,e [ 3a,c.e Observed Observed Observed Cycles RSS(1) Cycles RSS (1) Cycles RSS (1) a,c,e l
| |
| {
| |
| i s
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| OBSERVED TRANSIENTS GROUPED BY STRENGTH OF STRATIFICATION (RSS) INTERVALS No. Observed % of RSS Cycles Total j l
| |
| _, a,c.e 1 I
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| l \
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| , Note: The No. of groups is reduced by combining the intervals .70 < x 8 and . < x < .70 % of total = 3.4% for the interval-9 seu.ms m o 1-33
| |
| | |
| -s I
| |
| TABLE 1-13 (cont.)- . .
| |
| | |
| ==SUMMARY==
| |
| OF PLANT MONITORING HEATUP/COOLDOWN TRANSIENTS 1 WITHSTRENGTHOFSTRATIFICATION(RSS) .
| |
| RSS J % of Transients . {
| |
| - a,c.e RELATIVE NUMBER OF CYCLES OF STRENGTH OF STRATIFICATION (RNSSj)
| |
| AFTER GROUPING RSSj
| |
| ..,RNSSj Strength of
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| % Transients (2) j Stratification (1) a,c.e Nomenclature: .
| |
| (1) Strength of Stratification.(RSS)
| |
| (2) Relative Number of Cycles of Strength of Stratification (RNSS) ,
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| 4 as s soas.ie 1-34
| |
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| 1 TABLE 1-14 1
| |
| | |
| ==SUMMARY==
| |
| OF MONITORED TRANSIENT CYCLES (ONE HEATUP) I i
| |
| Plant No. of Cycles _
| |
| -- a,c.e 1
| |
| i l
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| Avg. Monitored Cycles: 15.75 = x; Selected No. of Design Cycles: 36.5 (added 30% to observed-maximum number of cycles, planta) s DESIGN DISTRIBUTI'ON APPLIED TO MAX NUMBER OF TRANSIENTS EXCEPTED MULTIPLIED BY 200 HEATUP OR C00LDOWN CYCLES i
| |
| No. of Transients RSS
| |
| -- a,c.e >
| |
| 1 6
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| e Total l
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| mm.mmm,o 1-35 e
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| i TABLE 1-15
| |
| | |
| ==SUMMARY==
| |
| OF.% TIMES AT MAXIMUM TEMPERATURE POTENTIAL RMTP g HEATUP a,c.e ,
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| TABLE 1-16 SURGE LINE TRANSIENTS - STRIPING FOR HEATUP (H) and C00LDOWN (C) .
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| Figure 1-19. Typical [ ]8'C'' Temperature Profiles m4.miw in 1-56
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| l 4
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| l SECTION 2.0 i i
| |
| STRESS ANALYSES
| |
| ]
| |
| Flow diagram figure 2-1 describes the procedure to determine the effects of thermal stratification on the pressurizer surge line based on transients developed in section 1.0. [
| |
| O ja,c.e Section 2.1 Addresses the structural or global effect of stratification Section 2.2 Addresses the local stress effects due to the nonlinear portion of the temperature profile Section 2.3 Addresses the total stress effects due to the oscillation of, the hot-to-cold boundary layer (striping) plus the thermal stratification stress 2.1 Piping System Structural Analysis 2.1.1 Introduction The thermal stratification computer analysis of the piping system to determine the pipe displacement, support reaction loads as well as moment and force loads in the piping is referred to as the piping system structural analysis.
| |
| These loads are used as input to the leak-before-break, fatigue, and fatigue crack growth evaluations. The thermal stratification condition consists of both axial and diametric variations in the pipe metal temperature, as described in section 1.0. The model consists of straight pipe and elbow elements for the ANSYS computer code. [' .
| |
| Ja,c.e These studies verified ammine.io 2-1
| |
| | |
| :the suitability of the ANSYS computer code for the thermal stratification analysis. (The displacements, reaction loads and pipe moments produced by the .
| |
| ja,c.e 2.1.2 Discussion On Typical Surge Line Analysis The piping layout for a typical surgeline is shown-in figure 2-3. The rigid-l support, Ril, originally installed to reduce deadweight and seismic loads provides resistance to the displacements caused by thermal stratification.
| |
| [
| |
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| |
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| |
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| |
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| |
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| |
| Based on the above discussion, the ANSYS computer Code is suitable for the thermal stratification analysis. [ .
| |
| . Ja,C,e
| |
| [
| |
| . ja,c,e an.mim io 2-4
| |
| | |
| J I q
| |
| , ja,c.e 2.1.3 Results For Comanche Peak Unit 1 Surge.Line The calculated piping stress due to thermal stratification for Comanche Peak Unit 1 surge !ine is reviewed to ensure that the system will not collapse in a
| |
| " hinge-moment" mechanism. The primary plus secondary stress limit for this-piping stress is given by ASME III, Section NB 3600. Equation 12 as '3.0 Sm.
| |
| The maximum stress intensity range, which occurs at the RCL hot leg nozzle. is j 52.1 ksi, this is less than the Code allowable value of 57.9 ksi. This 1 corresponds to a bounding thermal stratification case with AT = 320'F. It should be noted that tt e stress index for the hot leg nozzle in equation 12 1 were developed from finite element saalysis of the RCL nozzle. A summary of maximum ASME code calcolated stresses are presented in table 2-11.
| |
| ~
| |
| l 2.1.4 Additional Information on L'inear Equivalent Techniques l 2.1.4.1 Introduction i
| |
| 1 .
| |
| i A review of the pressurizer surge line thermal stratification for several clants indicated that the actual stratification temperature profiles are i better described by nonlinear diametric (cross-octional) temperature l ,
| |
| distributions. These temperature profiles will have effects on the global structural behavior of the surge lines in terms of loads and displacements.
| |
| The use of isoparametric solid elements has made possible the study of nonlinear :ross-sectional temperature profiles, such as step change of l temperatures at mid plane. This study was performed using a model developed for the WECAN computer code. In order to achieve a less costly analytical .
| |
| l solution, an alternative model using pipe and elbow elements was developed fcr the ANSYS computer code. These elements can only be loaded with a constant cross-section temperature or a linear top-to-bottom cross-section.
| |
| temperature. It, therefore, becomes necessary to establish an equivalent
| |
| '"'''"""'' 2-5
| |
| | |
| linear temperature profile which will result in the same deflections and loads in the piping system, as would a nonlinear temperature profile. It should be ~
| |
| l 1.oted that there are differences in the WECAN and ANSYS models as described in 1 section 2.1.2. These modeling differences will contribute to minor
| |
| * differences when results obtained from the analyses are compared. The purpose of the study and the comparison with the measured displacements is to verify the suitability of the ANSYS code for the thermal stratification global ,
| |
| analysis. The theoretical basis for the equivalent linear temperature profile is based on a cantilever beam model and is summarized below. ,
| |
| 2.1.4.2 Theory I
| |
| The closed form solution is deter:nined for the free-end vertical and axial displacements of a cantilever cylindrical beam subject to two types of l
| |
| l stratification temperature profiles:
| |
| a) linear equivalent variation from top to bottom; b) step change at distance Y obelow the beam centerline.
| |
| The axis of the beani (x-axis) lies'in a horizontal plane. The solution is based on the following principles: ;
| |
| i i
| |
| I a
| |
| g ,c.e .
| |
| 1 e
| |
| i aswocue in 2-6
| |
| | |
| i i
| |
| [
| |
| (2.1-2)
| |
| (2.1-3) !
| |
| 1 i
| |
| i (2.1-4) !
| |
| ja,c.e ;
| |
| : 5. For a cantilever beam subject to thermal stratification, the axial-force (F)andbendingmoment(M)arezeroateachcrosssection(A),
| |
| , thus, F=IA e dA = 0 (2.1-5)
| |
| M=/A o y dA = 0 (2.1-6) i The above equations are solved in closed form with the following results:
| |
| [
| |
| (2.1-7)
| |
| (2.1-8)
| |
| )a,c.e -
| |
| t G
| |
| =7*'*''
| |
| 2-7
| |
| | |
| E i
| |
| 3 a,c.e -
| |
| The solution for the equivalent linear temperature in the form of coefficients J
| |
| ik is obtained by equating (2.1-7) with (2.1-9) and (2.1-8) with (2.1-10).
| |
| L e
| |
| 3 a,c.e
| |
| | |
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| |
| ~
| |
| : r. .
| |
| t ,
| |
| 2.1.4.3 Application .
| |
| The deflections and loads in the surge line for case 3 (step at mid plane) i have been calculated by WECAN. The same step change temperature profile is l converted to an equivalent linear temperature profile (case 3L) for ANSYS j
| |
| using the J ik coefficients with Y, = 0. Table 2-4 is an example for 14-inch schedule 140 pipe. The case 3 and case 3L temperature profiles used in the analyt.es are shown in figure 2-23c and 2-23d. The results are presented in table 2-2.
| |
| 2.1.4.4 Discussion The suitability of the ANSYS compu'ter code for the thermal stratification
| |
| - global analysis is demonstrated by the comparisons between case 3 and case 3L. WECAN and ANSYS pipe displacements on table 2-2 also confirm this. In addition, case 3L is representative of the eleven analysis cases which represent various step temperature profiles along the pipe axis.
| |
| 2.1.5 Conclusions -
| |
| Analytical studies with the ANSYS and WECAN computer codes have confirmed the validity of using an equivalent linear diametric temperature profile to represent the thermal stratification for displacement and loads. Comparison between the ANSYS results and the plant measured displacements will be performed when the plant monitoring data become available. Eleven cases of -
| |
| thermal stratification were analyzed using the ANSYS code for the Comanche Peak Unit 1 surgeline. Results for all other cases of stratification were obtained by interpolation. The resulting loads on the pressurizer and hot leg nozzles are acceptable. The surge line pipe stress satisfies the ASME III NB-3600 Code Equation 12 limits.
| |
| 3697s/041889 10 2-9
| |
| | |
| r 1
| |
| 2.2 Local Stress Due to Non-Linear Thermal Gradient
| |
| 'l 2.2.1 Explanation of Local Stress q I
| |
| Figure 2-24 depicts the local axial stress components in a beam with a sharply l
| |
| nonlinear metal temperature gradient. Local axial stresses develop due to the i restraint of axial expansion or contraction. This restraint is provided by
| |
| * the material in the adjacent beam cross section. For a linear top-to-bottom temperature gradient, the local axial stress would not exist. [ -
| |
| 1 3a,c.e 1
| |
| 2.2.2 Superposition of Local and Structural Stresses For the purpose of .this discussion, the stress resulting from the global structural analysis (section 2.1) will be referred to as " structural stress."
| |
| 1
| |
| [ .j Ja,c e Lnal and structural stresses may be superimposed to obtain !
| |
| the total stress. This is true because linear elastic analyses are performed ]
| |
| and the two stresses are independent of each other as summarized in figure 2-25.
| |
| Figure 2-26 presents the results of a test case that was performed to , ,
| |
| demonstrate the validity of superposition. As shown in the figure, the super-position of local and structural stress is valid. [' ,
| |
| ja,c.e 2.2.3 Finite Element Model of Pipe for Local Stress >
| |
| A short. description of the pipe finite element model is shown in figure 2-27.
| |
| The model with thermal boundary conditions is shown in figure 2-28. Due to w w .,i. m o 2-10
| |
| | |
| symmetry of the geometry and thermal loading, only half of the cross section was required for modeling and analysis. [
| |
| ja,c.e 2.2.4 Pipe Local Stress Results Figure 2-29 shows the' temperature distributions through the 14 in, schedule 160 pipe wall ( .
| |
| , ja,c,e uc.mi . i.
| |
| 2-11
| |
| | |
| 1 2.2.5 Unit Structural Load Analyses For Pipe In order to accurately superimpose local and global structural stresses, several additional stress analyses were performed using the 2-D pipe model. -
| |
| [ .
| |
| l ja,c.e l 2.2.6 RCL Hot Leg Nozzle Analysis Two RCl surge line nozzle models were developed to evaluate the effects of thermal stratification. These two models are shown in figures 2-43 and 2-44.
| |
| [
| |
| 3a,c,e Figures 2-45 thru 2-53 present color contour plots of temperature and stress distributions in the surge line RCL nozzle. A summary of local stresses in the RCL nozzle due to thermal stratification is given in table 2-6. A summary l of stresses for unit leading applied is shown in table 2-7.
| |
| l 2.2.7 Conservatism Conservatism in the local stress analysis are listed below:
| |
| : 1. The hot / cold fluid interface is assumed to have zero width. A more gradual change from hot to cold would significantly durease local ~
| |
| stresses. '
| |
| : 2. Stresses are based on linear elastic analysis even though stress levels exceed the material yield point.
| |
| mar..i.= in 2-12
| |
| | |
| 2.3 Thermal Striping 2.3.1 Background ,
| |
| At the time when the feedwater line cracking' problems in PWR's were first -
| |
| discovered,itwaspostulatedthatthermaloscillations(striping)may-significantly contribute to the fatigua cracking problems. These oscillations
| |
| . were thought to be due to either mixing of hot and cold fluid, or turbulence in the hot-to-cold stratification layer from strong buoyancy forces during low flow rate conditions. (See figure 2-54 which shows the thermal striping.
| |
| fluctuation in a pipe). Thermal striping was verified _to occur during subsequent flow model tests. Results of the flow model tests were used to establish boundary conditions for the stratification analysis and to provide striping oscillation data for evaluating high cycle fatigue.
| |
| Thermal striping was also examined during water model flow tests performed for the-Liquid Metal fast Breeder Reactor primary pipe loop. The stratified, flow was observed to have a dynamic interface region which esci11'ated in a wave
| |
| ~
| |
| pattern. (See figur's 2-55 for tort pipe sizes, thermocouple locations, and table 2-8 for typical frequency of-striping oscillations.) -These dynamic oscillations were shown to produce significant fatigue damage (primary crack initiation). The same interface oscillations were observed in experimental studies of thermal striping which were performed'in Japan by Mitsubishi Heavy Industries.
| |
| 2.3.2 Additional Background Information Thermal striping was examined during 1/5 scale water model flow tests performed for the Liquid Metal Fast Breeder Reactor primary pipe loop. These tests were performed by Westinghouse at the Waltz Mills test facility. In order to measure striping, thermocouple were positioned at 5 locations in the hot leg piping system (three in the small diameter pipe and two in the large diameterpipe.) The inside diameters of the large and small pipes were 6-1/2 and 4 inches, respectively. Figure 2-56 shows the test setup and locations of the thermocouple. (Figure 2-55 shows test pipe sizes with circumferential e
| |
| position of thermocouple.) Thermocouple locations were selected [ ]
| |
| up..an ie p .13
| |
| | |
| [ Ja,c.e The thermocouple extended [ Ja.c.e into the fluid. The flow rates and
| |
| ~
| |
| corresponding Richardson numbers for each pipe size are shown in table 2-9.
| |
| 8 'C''
| |
| A total of [ 3 tests were performed and evaluated. Three parameters were measured during the water tests which help define thermal , i striping: frequency of fluctuations, duration, and amplitude of delta fluid ,
| |
| temperature. The [ ,
| |
| ]''C# were recorded in the discussion of test results and are presented in table 2-10.
| |
| The frequencies of the temperature fluctuations from these test results were reported to be in the range of [ Ja,c.e As shown in table 2-10,the[-
| |
| ga c.e !
| |
| In order to use the water test data for the surge line striping analysis, the test data with a [
| |
| Ja.c.e was chosen to be used in the evaluation. From -
| |
| table 2-9, the [ Ja c.e inch I.D. pipe with flow rates of (
| |
| ja c.e ,
| |
| for the pressurizer surge line.
| |
| When all other factors are equal, it has been shown that the thermal striping .
| |
| stress is [ Ja,c.e wwwwe 2-14
| |
| | |
| , ~ A typical value of usage factor was calculated with the [
| |
| Ja,c,e as follows:
| |
| [
| |
| ja.c.e This distribution corresponded to [ 38 '*''
| |
| cbnsidered to occur at a stress level calculated with frequencies of (
| |
| ),a,c,e respectively. Calculations revealed that there was [
| |
| Ja,c.e in the usage factor when a [
| |
| Ja,c.e Therefore,[ Ja,c.e was assumed in all usage factor calculations.
| |
| For the Commanche Peak Unit 1 Pressurizer surge line, the frequency of [
| |
| ]a,c,e was used in the [
| |
| ja,c e As r %wn in table 2-10, the amplitude of AT varies from [
| |
| ]a,c.e of the full AT between the hot and cold fluid temperatures.
| |
| For the Commanche Peak Unit 1 Surge line, the amplitude was assumed to be at
| |
| [
| |
| Ja,c.e as shown by the curve in figure 2-57. This is conservative since a higher AT results in higher stress.
| |
| The maximum duration of thermal striping from table 2-10 shows that thermal striping occurred for [ 3a,c.e For the Commanche Peak Unit 1 pressurizer surge line, thermal striping was considered to occur [
| |
| ja.c.e ne,.wm to 2-15
| |
| | |
| 2.3.3 Thermal Striping Stresses Thermal striping stresses are a result of differences between the pipe inside .
| |
| surface wall and the average through wall temperatures which occur with time, due to the oscillation of the hot and cold stratified boundary. (Seefigure 2-58 which shows the typical temperature distribution through the pipe wall). - !
| |
| [
| |
| ja,c.e The peak stress range and stress intensity is calculated from a 2-D finite element analysis. (See figure 2-59 for a description of the model.)
| |
| (
| |
| Ja.c.e The methods used to determine alternating stress intensity are defined in the ASME code. Several locations were evaluated in order to determine the locatic'n where stress intensity was a maximum.
| |
| Stresses were intensified by K3 to account for the worst stress concentra-tion for all piping element in the surge line. The worst piping elements were the butt weld and the tapered transition.
| |
| [
| |
| )a,c.e 2.3.4 Summary of Striping Stress Considerations I .
| |
| 3a,:,e
| |
| -,.. . . j 2-a i
| |
| | |
| 'W-~- - - - - - - - - - _ . _ _ _ _ _ _ _ _ _ _ _ _
| |
| C .
| |
| 1 i
| |
| I
| |
| * b I
| |
| l 1
| |
| aa O
| |
| l l
| |
| i i
| |
| :I
| |
| -4 a
| |
| 1 i
| |
| l l
| |
| l ,
| |
| I i 1 <
| |
| i l
| |
| 1 I{
| |
| )
| |
| I a
| |
| e a
| |
| 3 a,c.e 3697s/041869 10 2-17
| |
| | |
| [
| |
| )a,c.e
| |
| * 2.3.5 Thermal Striping Total Fluctuations and Usage Factor .
| |
| Thermal striping transients are shown at a AT level and number of cycles. ,
| |
| e i
| |
| e I
| |
| i 9
| |
| 6 e
| |
| ja,c.e I
| |
| =>..im to 2-18
| |
| | |
| WESTINGHOUSE PQoPRIETAQY CLASS 8 2.3.6 Conservatism The conservatism in the striping analysis are: striping occurs at one location; surface film coefficients assume high values with constant flow; and conservative design transients are used. The major conservatism involves the -
| |
| ~
| |
| combination of maximum striping usage factor with fatigue usage factor from all other stratification considerations. The [
| |
| ja,c.e s
| |
| 4 3007s/0did8910 p}g
| |
| | |
| i TABLE 2-1 COMPARISON OF WECAN AND ANSYS RESULTS FOR LINEAR STRATIFICATION - Case 2 '
| |
| (Displacements in Inches)
| |
| ANSYS/WECAN (JOBANSF)WECAN (AGJAQLM) ANSYS (PERCENTAGE) a,C,9 e
| |
| 9 e
| |
| O e -
| |
| . 8 k
| |
| . i i
| |
| - l l
| |
| mim:io 2-20 L
| |
| | |
| 1 TADLE 2-2 COMPARISON OF WECAN [. Ja,c.e AND .
| |
| ANSYS (l Ja,c.e~RESULTS FOR CASE 3 Case 3L/ Case 3 Location Direction WECAN Case 3 ANSYS Case 3L (Percentage) a,c,a l
| |
| l 1
| |
| 1
| |
| ]
| |
| 'l i
| |
| . Case 3L ANSYS: DCISKXY, 11/12/88 4
| |
| y
| |
| | |
| i i
| |
| TABLE 2-3 J TEMPERATURE DISTRIBUTIONS IN COMANCHE PEAK UNIT 1 PRESSURIZER SURGE LINE ,
| |
| i i
| |
| . a,c.e
| |
| ~
| |
| Case PZR Temp ('F) RCLTemp(*F) TTop(F) TBot ( F) AT(*F) 1 1 455 140 455 140 315 2 455 140 455 273 182 -
| |
| 3 455 140 455 385 70 4 540 320 540 380 160 5 653 550 653 550 103 6 530 350 530 450 80 7 653 617 653 617 36 8 653 617 653 653 0 9 653 583 653 583 70 10 540 320 540 440 100 11 530 350 530 333 144 m
| |
| l l
| |
| -l
| |
| ~
| |
| e l
| |
| ...i. 2-22 l
| |
| | |
| TABLE 2-4.
| |
| ~
| |
| THE EQUIVALENT LINEAR COEFFICIENTS J ik (14 inch-Schedule 140 Pipe)
| |
| Y, J O J J hh hc ch ce a,c.e 4
| |
| a e
| |
| e e
| |
| M e
| |
| 9 0
| |
| 4 4 O m m.4 i m ia.
| |
| 2-23 t
| |
| 'l
| |
| | |
| ' TABLE 2-5 CDMANCHE PEAK UNIT 1 SURGE LINE -
| |
| MAXIMUM LOCAL AXIAL STRESSES AT [ Ja,c.e Local Axial Stress (psi) location Surface Maximum Tensile Maxinium Compressive
| |
| ~
| |
| a,c.e ~
| |
| l I
| |
| Note: Local thermal stresses shown are for a AT = 260*F.
| |
| )
| |
| 1 l
| |
| . i e
| |
| w . w io 2-24
| |
| | |
| TABLE 2-6
| |
| | |
| ==SUMMARY==
| |
| OF LOCAL' STRATIFICATION STRESSES
| |
| -IN THE COMANCHE PEAK UNIT 1 SURGE'LINE:AT THE RCL N0ZZLE All Stress'in psi
| |
| .. Linearized Stress Peak Stress.
| |
| Intensity Range- Intensity Range Diametral
| |
| . Location Location Inside Outside Inside Outside a,c.e-
| |
| - Nozzle-Hot Leg Bottom 10036 10019 7684 18606
| |
| - Crotch Region Nozzle-Hot Leg 45' 4569 5665 1562 10614 Crotch Region Nozzle-Hot Leg Side 8365 8808 10468 10954 Crotch Region Nozzle-Hot Leg 135' 11695 .13635 12980 14623 Crotch Region Nozzle-Hot Leg- 150' 14104 14012 14685 '20556
| |
| - Crotch Region ,
| |
| Nozzle-Hot Leg 165* 9633 16038 9143 26877 Crotch Region Nozzle-Hot Leg Top 7657 16622 6610 28966
| |
| . Crotch Region-Safe End Top 32411 55538 34956 60123 Safe End Side 10361 19249 11326 20812 e
| |
| i O
| |
| S 9
| |
| -U w.mimaa 2-25
| |
| | |
| TABLE 2-7
| |
| | |
| ==SUMMARY==
| |
| OF PRESSURE AND BENDING INDUCED STRESSES '
| |
| IN THE COMANCHE PEAK UNIT 1 SURGE LINE RCL N0Z2LE FOR UNIT LOAD CASES All Stress in psi i
| |
| Linearized Stress Peak Stress-Intensity Range Intensity Range I Diametral Unit Loading Location Location Condition Inside Outside Inside Outside
| |
| . J a,c.e Nozzle-Hot Leg Bottom 1000 psi 2186 5385 1109 7099 Crotch Region Pressure Nozzle-Hot Leg 45' 1000 psi 9703 4602 9342 3633 Crotch Region Pressure Nozzle-Hot leg Side 1000 psi 15441 1154 18514 1692 Crotch Region Pressure Nozzle-Hot Leg 135' 1000 psi 9703 4602 9342 3633 Crotch Region Pressure Nozzle-Hot Leg Top 1000 psi 2186 5385 1109 7099 Crotch Region Pressure Nozzle-Hot Leg Bottom 1.0E6 in-lb 3476 4536 421 9142 Crotch Region Bending About RCS Longitudinal Axis Nozzle-Hot Leg 45' 1.0E6 in-lb 2513 1848 1674 4464 Crotch Region Bending About RCS Longitudinal Axis Nozzle-Hot Leg Side 1.0E6 in-lb 654 1228 338 1004 Crotch Region Bending About RCS Longitudinal Axis Nozzle-Hot Leg 135' 1.0E6 in-lb 2513 1848 1674 4464 Crotch Region Bending About RCS Longitudinal Axis .
| |
| Nozzle-Hot Leg Top 1.0E6 in-lb 3476 4536 421 9142 Crotch Region Bending About RCS Longitudinal Axis ,
| |
| Nozzle-Hot Leg Bottom / 1.0E6 in-lb 1670 1471 720 1050 Crotch Region Top Bending About RCS Vertical Axis (Out of Plane)
| |
| Nozzle-Hot Leg 45'/135' I.0E6 in-lb 1280 1870 675 3362 .
| |
| Crotch Region Bending About RCS Vertical Axis (Out of Plane)
| |
| Nozzle-Hot Leg Side 1.0E6 in-lb 1074 1457 1006 3404 Crotch Region Bending About RCS Vertical Axis (Out of Plane) 2"'"$' "" " 2-26
| |
| | |
| \
| |
| 9 TABLE 2-8 STRIPING FREQUENCY AT 2 NAXINUN LOCATIONS FRON 15 TEST RUNS
| |
| .l J,
| |
| ". l M
| |
| ' - 8,C,9 9
| |
| 1 l
| |
| 1 \
| |
| l 1
| |
| 4 l
| |
| )
| |
| 1
| |
| .)
| |
| ~
| |
| 'O M
| |
| i 6
| |
| . i 1
| |
| i
| |
| ~j sies. m mesao 2-27 ,
| |
| 1
| |
| | |
| i l I TABLE 2-9 FLOW RATES AND RICHARDSON NUMBER I FOR WATER MODEL FLOW TESTS Cold Water Flow Rate -
| |
| Pipe Section (GPM) Ri 4.0 inch I.D. a,c.e
| |
| =-
| |
| .?
| |
| 1' 6.5 inch I.D.
| |
| 1 l
| |
| Y 4
| |
| un.mim:io 2-28
| |
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| S E
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| L -
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| %C Y
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| ) G L V A A I
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| N S E E T L 3 O %C P Y C
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| O / C L N S E O E N S .)
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| _ P A A Y H E 2
| |
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| |
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| _ C T D # L T 2 O
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| - M ) e 0 R N u r
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| _ F O et
| |
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| |
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| |
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| |
| - g ni id d
| |
| a
| |
| *
| |
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| |
| p r
| |
| i el
| |
| _ P l
| |
| _ p
| |
| _
| |
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| |
| r l eh
| |
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| |
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| |
| * Y!8 1
| |
| (
| |
| | |
| TABLE 2-11 ASME CODE STRESS
| |
| | |
| ==SUMMARY==
| |
| - a,c.e d
| |
| *ASME Boiler and Pressure Vessel Code, Section III,1979 . Edition.
| |
| l l
| |
| l i
| |
| ,4 I
| |
| t a
| |
| wr.mine in 2-30 1
| |
| 'l
| |
| : I
| |
| | |
| DETERMINATION OF THE EFFECTS OF THERMAL STRATIFICATION a,c,e i
| |
| i i,
| |
| f I
| |
| l l
| |
| 1 1
| |
| e l .
| |
| Figure 2-1. Determination of the Effects of Thermal Stratification m,.ns,ue ,a g_37
| |
| | |
| I
| |
| /
| |
| i a,c.e 6
| |
| 1
| |
| )
| |
| j Figure 2-2. Stress Analysis un.mim is 2-32
| |
| | |
| 2 -
| |
| ^
| |
| X PRESSURIZER r %'
| |
| i Y
| |
| j fS2 P 6 FT 3 FT 27 FT
| |
| , SNUBBER SNUBSER Q, RiI 21 FT j RCL HOT LEG l
| |
| d Figure 2-3. Typical Press.irizer Surge Line Layout I
| |
| 2-33 wi.mim io f
| |
| | |
| a,c,e I .
| |
| i i
| |
| .i -
| |
| 1 i
| |
| )
| |
| i
| |
| - Figure 2-4. Cases 1 to 4: Diametric Temperature Profiles usi.nsine io 2-34
| |
| | |
| a,c,e s
| |
| 1 i
| |
| I 1
| |
| i i
| |
| 1 l
| |
| l 3
| |
| 1 I
| |
| l 1
| |
| Figure 2-5. Case 5: Diametric and Axial Temperature Profile ue,. w ses ,o g_35
| |
| | |
| E I i
| |
| 4 a,c.e l
| |
| j l
| |
| 1 1
| |
| I i
| |
| Figure 2-6. Finite Element Model of the Pressurizer. Surge Line Piping ,
| |
| General View- !
| |
| sasi.mim i. ''
| |
| '2-36 ,
| |
| I
| |
| | |
| a,c.e i
| |
| d I
| |
| i .
| |
| 1 >
| |
| l Figure 2-7. Finite Element Model of the Pressurizer Surge Line Piping Hot Leg Nozzle Detail mi.mises io t
| |
| 2-37
| |
| | |
| 1 a .c .e
| |
| )
| |
| * l i
| |
| ~
| |
| l
| |
| * i i
| |
| {
| |
| l
| |
| ~ - *-l tigure 2-8. Thermal Expansion of the Pressurizer Surge Line Under Uniform Temperature . 1 l
| |
| J
| |
| =,.mi ..
| |
| i 2-38 !
| |
| F
| |
| | |
| (
| |
| a,c.e l
| |
| l l
| |
| 1 .
| |
| 1 Figure 2-9. Case 2 (linear) Temperature Profile. at Hot leg Nozzle m in n ie 2-39
| |
| | |
| 9 a,c,r i .
| |
| I 1
| |
| 1 1
| |
| - l l
| |
| l 1
| |
| i s
| |
| 1
| |
| . Figure 2-10. Case 2 (linear) Temperature Profile at Pressurizer Elbow .
| |
| N83e4139At le 2-40 0
| |
| | |
| a,c.e !
| |
| = 1 i
| |
| 1
| |
| .I I
| |
| l .
| |
| /
| |
| Fihre2-11. Thermal Expansion of Pressurizer Surge Line Under Linear i Temperature Gradient .l 1
| |
| ues m,su se 2-41 i'
| |
| | |
| y l
| |
| a,c.e .
| |
| l l
| |
| f l .
| |
| - ?
| |
| l i
| |
| Figure 2-12. Bowing of Beams Subject to Top-to-Bottom Temperature Gradient l
| |
| t M1o/12108410 2-42
| |
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| &,C,t N
| |
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| = 0 a
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| ec=
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| Figure 2-54. Thermal Striping Fluctuation .
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| mi. we 2-84
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| l Figure 2-56. Water Model of LMFBR Primary Hot leg
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| l c net. mins in 2-86 i
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| l Figure 2-57. Attenuation of Thermal Striping Potential by Molecular l Conduction (Interface Wave Height of [ ')]ac.e j
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| . w.mim io 2-87 !
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| 4
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| Figure 2-58. Thermal Striping Temperature Distribution ,,
| |
| l- m,..im in 2-88
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| 7_. _ - _ _ _ , - . - - - - - - . - - - - . - - - - - - - - - - - - - - - - - - - - - - ' - - - - - - - - " - - - - - " ' " - " -
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| | |
| SECTION 3.0
| |
| . ASME SECTION IIi FATIGUE USAGE FACTOR EVALUATION 3.1 Code and Criteria
| |
| ~
| |
| Fatigue usage factors for the Comanche Peak Unit i surge line were evaluated based on the requirements of the ASME B & PV Code, Section III (reference 3-1), Subsection NB-3600, for pipir.g components. The more detailed techniques
| |
| ~
| |
| of NB-3200 were employed, as allowed by NB-3611.2. ASME III fatigue usage factors were calculated for [ Ja,c.e points in the surge line piping using program WECEVAL (reference 3-2).
| |
| 3.2 Previous Design Methods Previous method of surge line piping fatigue evaluation used the NB-3653 techniques but with thermal transients defined by W SSDC 1.3 F(3-3] and 1.3.X (3-4], assuming the fluid surges to sweep the surge line piping with an axisymmetric temperature loading on the pipe inside wall. These evaluations produced typical usage factors of approximately [
| |
| la,c.e at girth butt welds,[ Ja,c.e 3t elbows and bends, and [ 3a,c.e at the RCL hot leg nozzle crotch region. Effects of stratification were not included in previous design analyses.
| |
| It must be notea that these usage factors are conservative since, in the design process, calculations are carried to the point where results meet code requirements, and are not further refined to reduce the usage factor.
| |
| 3.3 Analysis for Thermal Stratification With thermal transients redefined to account for thermal stratification as described in section 1.0, the stresses in the piping components were established (section 2.0) and new fatigue usage factors were calculated. Due -
| |
| to the non-axisymmetric nature of the stratification loading, stresses due to all loadings were obtained from finite element analysis and then combined on a
| |
| ~
| |
| stress component basis.
| |
| wuuwne io 3-1
| |
| | |
| L 3.3.1 Stress Input' 5 tresses in the pipe wall due to internal pressure, moments and thermal-stratification' loadings were obtained from the WECAN 2-D analyses of 14 inch, .-
| |
| schedule 160, and a short spool piece of' schedule 140 pipes. ( _
| |
| )a,c.e-(
| |
| .)a.c.e 3.3.2 Classification'and Combination of Stresses
| |
| ~
| |
| As described in 3.3.1 the total stress in the pipe wall was determined for each transient load case. Two types of stress were calculated - Sn (Eq 10) . ,
| |
| to determine elastic plastic penalty factors, K,, and Sp (Eq 11) peak-stress. For most components in the surge line (girth butt welds, elbows, bonds) no gross structural discontinuities are present. As a result the code-defined "Q" stress (NB-3200), or C3ElsaT, abi lin Eq _(10)-
| |
| b of NB-3600 is zero. Therefore, for these components, the Eq. (10) stresses are due to pressure and moment. .
| |
| For the RCL hot leg nozzle, the results of the 3-D finite element WECAN ~
| |
| -analysis of the nozzle were used to determine "Q" stress for transients with j
| |
| stratification in the nozzle. Note also that the Eq. (10) stresses included
| |
| ^
| |
| appropriate stress intensification using the secondary stress indices from-NB-3681. The pressurizer nozzle load was also specially qualified by finite element analysis technique.
| |
| ' I l
| |
| . i me .ime ie 3-2 i
| |
| | |
| i Peak stresset, including the total surface stress from all loadings -
| |
| pressure, moment, stratification - were then calculated for each transient.
| |
| [
| |
| )a,c.e 3.3.3 fumulative Fatigue Usage Factor Evaluation Program WECEVAL uses the n S and p S stresses calculated for each transient to determine usage factors at selected locations in the pipe cross section.
| |
| Using a standard ASME method, the cumulative damage calculation is performed accordingtoNB-3222.4(e)(5). The inside and outside pipe wall usage factors were evaluated at [ J''C'' through the pipe wall'of the 2-D WECAN model.
| |
| This includes:
| |
| : 1) Calculating.the Sn and Sp ranges, K , and Salt I # 'V''Y possible combination of the [ Ja, ,e transient load sets.
| |
| . 2) For each value of Salt, use the design fatigue curve to determine the maximum number of cycles which would be allowable if this type of cycle were the only one acting. These values, N1 , N 2 ...N n '
| |
| were determined from Code figures 1-9.2.1 and I-9.2.2, curve C, for I austenitic stainless steels.
| |
| . 3) Using the actual cycles of each transient loadset supplied to WECEVAL, n y ,n2'***"n, calculate the usage factors U ,
| |
| 1 U . . .U 2 n IT
| |
| * Ui * "i/N i. This is done for all possible combinations. If Ng is greater than 10 11 cycles, the value of U$ is taken as zero.
| |
| l l
| |
| (
| |
| 3a,c.e mm,wmo 33
| |
| | |
| 1
| |
| [
| |
| ja.c e
| |
| : 4) The cumulative usage factor, l' cum, is calculated as U cum *U1*
| |
| U2 * ***
| |
| * Un. The code allowable value is 1.0. ,
| |
| 3.3.4 Simplified Elastic-Plastic Analysis When code Eq. (10),nS , exceeded the 3Sm limit, a simplified elastic plastic l
| |
| analysis was performed per NB-3653.6. This requires separate checks of expansion stress, Eq. (12), and Primary Plus Secondary Excluding Thermal Bending Stress, Eq. (13), and Thermal Stress Ratchet, and calculation of the
| |
| ; elastic plastic penalty factor, Ke, which affects the alternating stress by Salt
| |
| * EeS p/2. The K, values for all combinations were automatically calculated by WECEVAL. Thermal stress ratchet is also checked by WECEVAL.
| |
| Eq. (13) is not affected by thermal stratification in ^he pipe where no gross structural discontinuities exist, but required to be verified at the nozzle.
| |
| l Eq. (12) was evaluated in the Global ANSYS analysis by checking the worst -
| |
| l possible range of stress due to the expansion bending moments (section 2).
| |
| 3.3.5 Fatigue Usage Results The maximum Usage factors were [ ja c.e at the RCL nozzle safe end (node 1010, figure 1-7) and [ la.c.e at the 5-D bend located underneath the pressurizer, (figure 1-7) which are less than the code allowable of 1.0.
| |
| The above usage factors included the effects of striping. The nature of striping damage is at a much higher frequency, varies in location due to fluid ,
| |
| level changes and is maximized at a different location than the ASME usage factor.
| |
| e ues,* ue ,o 34
| |
| | |
| i 3.4 Conservatism in Fatigue usage Calculation I
| |
| The above calculated ASME usage factors contain the inherent conservatism i known to be in the ASME Code methods. These include the conservatism in the elastic plastic penalty factor, K,, the method of combining loadsets based -
| |
| . on descending Salt, and the factor of 2 on stress and 20 on cycles in the j design fatigue curve.
| |
| Also, due to input limitations in program WECEVAL, the maximum value of peak stress intensification for all loading types was used. This was conservative l J
| |
| at girth butt welds, since K1 = 1.2, K2 = 1.8, K3 = 1.7 in NB-3681 and K=1.8 was used in WECEVAL for all stresses.
| |
| t 3.5 References i
| |
| 3-1. ASME Boiler and Pressure Vessel Code, Section III,1986 Edition.
| |
| 3-2. WCAP-9376', WECEVdt, A Computer Code to Perform ASME BPVC Evaluations ;
| |
| Using Finite Element Model Generated Stress States, April,1985. l l [ Proprietary) )
| |
| 3-3. W Systems Standard 1.3.F. Rev. O. (Proprietary) 3-4. W Systems standard 1.3.X, Rev. 0 (Proprietary) i D
| |
| l mm. un io 3-5
| |
| | |
| WESTINGHOUSE PQopQlETARY class 2 SECTION 4.0 FATIGUE CRACK GROWTH 4.1 Introduction To determine the sensitivity of the pressurizer surge line to the presence of
| |
| . small cracks when subjected to the transients discussed in section 1, fatigue crack growth analyses were performed. This section summarizes the analyses and results.
| |
| Figure 4-1 presents a general flow diagram of the overall process. The methodology consists of seven basic steps as shown in figure 4-2. Steps 1 thru 4 a'ra discussed in sections 1 and 2 of this report. Steps 5 thru 7 are specific to fatigue crack growth and are discussed in this section.
| |
| There is presently no fatigue crack growth rate curve in the ASME Code for austenitic stainless steels in a water environment. However, a great deal of work has been done recently which. supports the development of such a curve.
| |
| An extensive study was performed by the Materials Property Council Working Group on Reference Fatigue Crack Growth concerning the crack growth behavior of these steels in air environments, published in reference 4-1. A reference curve for stainless steels in air environments, based on this work, will appear in the 1988 Addenda of Section XI. This curve is shown in figure 4-3.
| |
| A compilation of data for austenitic stainless steels in a PWR water environment was made by Bamford (reference 4-2), and it was found that the effect of the environment on the crack growth rate was very small. For this l -
| |
| reason it was estimated that the environmental factor should be set at 1.0 in the crack growth rate equation from reference 4-1. Based on these works (references 4-1 and 4-2) the fatigue crack growth law used in the analyses is ,
| |
| , as shown in figure 4-4.
| |
| e
| |
| : m. im. ie 43
| |
| | |
| l 4.2 Initial Flaw Size Various initial surface flaws were assumed to exist. The flaws were assumed to be semi elliptical with a six-to-one aspect ratio. The smallest flaw size ,
| |
| assumed was one having one-fourth the depth of a surface flaw found acceptable -
| |
| by paragraph IWB 3514.3, Allowable Flaw Standards for Austenitic Piping-of the ASME Code. The largest initial flaw assumed to exist was one with a depth equal to 10% of the wall thickness.
| |
| 4.3 Critical Locations for FCG All [ ]a,c.e locations (as shown in figure 1-18), representing all cross sections of the surge line where thermal stratification could occur, were checked for fatigue crack growth. Figure 4-5 identifies-(
| |
| Ja,c e locations (locations [ la,c.e) as sections along the length of the surge line. Figure 4-6 identifies the positions at each location where fatigue crack growth was checked. These positions ([
| |
| ]a,c.e) are controll.ing pcsitions because the global structural bending stress is maximum at. positions (
| |
| la c.e while the local axial stress on -
| |
| ).a,c.e the inside surface is maximum at positions [-
| |
| Location ( ]a,c.e (as shown in figure 1-18) is not shown on figures '4-5 and 4-6. The location [ Ja c.e stratification profile exists at the surge line RCL nozzle when the RCP pump is not running and, therefore, turbulent mixing caused by flow in the main RCL piping is not occurring. This effect was observed in the surge line monitoring programs.
| |
| I J I
| |
| me. maim io 4-2 1 I
| |
| | |
| 4.4 Results of FCG Analysis Results of the fatigue crack growth analysis are presented in table 4-1 for a 10% wall initial' flaw.
| |
| Conservatism existing in the fatigue crack growth analysis are listed below.
| |
| : 1. Plant operational transient data has shown that the conventional design transients contain significant conservatism
| |
| [
| |
| ja.c.e
| |
| : 4. Fatigue crack growth calculations based conservatively on elastic stresses -
| |
| : 5. FCG neglects fatigue life prior to initiation
| |
| ~
| |
| 4.5 References 4-1. James, L. A. and Jones, D. P., " Fatigue Crack Growth Correlations for Austenitic Stainless Steel in Air," in Predictive Capabilities in Environmentally Assisted Cracking, ASME publication PVP-99, December 1985.
| |
| 4-2. Bamford, W. H., " Fatigue Crack Growth of Stainless Steel Reactor Coolant Piping in a Pressurized Water Reactor Enviornment," ASME Trans. Journal of Pressure Vessel Technology, Feb. 1979.
| |
| neswim ie 43
| |
| | |
| TABLE 4-1 FATIGUE CRAcX GROWTH RESULTS FOR 10% WALL INITIAL FLAW SIZE Initial Initial Final (40 yr) Final Flaw .
| |
| Location ~ Position Size (in) (% Wall) Size (in) (% Wall) a,c.e e
| |
| m
| |
| . ' j f.
| |
| D e
| |
| 9 e
| |
| m e.-a m m is 44 l
| |
| | |
| =
| |
| 1 i
| |
| .i a,c.e.
| |
| i
| |
| -i
| |
| ')
| |
| 1 t
| |
| 1 l
| |
| i 1
| |
| 1 l
| |
| l i
| |
| 1 4
| |
| Figure 4-1. Determinination of the Effects of Thermal Stratification on Fatigue Crack Growth DeGae-121980 it .
| |
| l 4-5
| |
| | |
| . a,c.e i
| |
| ~ -
| |
| Figurn 4-2. Fatigue Crack Growth Methodology ses..,sswa -
| |
| 46 )
| |
| i 4
| |
| | |
| WSSTINGH;USE PROPRIETARY Ct. ASS 2 s.e d t
| |
| /.i /rp+.. ..
| |
| f SOUD UNC5 FOR 70T p l/ l i )*
| |
| DASMCD U'C5 90R S509 . . -
| |
| [ _ _ ), . ;. I ', s..
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| ..+. .... , . .
| |
| .. g . .. : . ./. ..j ..
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| ../.. ; /.. /.. . ....
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| . . . . .. .,f.. ..,.7..,,f
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| . .. . . . . . .. a ..,.. .. .
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| .. . . - . . . L1 v
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| 73 p ga 's( al!
| |
| f . .. .. .
| |
| f....5.,...,....,.
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| |
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| |
| v . .
| |
| [R s 0.79 4.. r-
| |
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| . . . . .. . ./ . i x ef e
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| : c. A l l[ t
| |
| . ..,..g._,.f......f.f........,......,....,.,.....,....
| |
| .. .g ._9 7 7 ..
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| 9.,. ,
| |
| : p. . p q 9...g
| |
| .. . .y h..g..
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| /,. . ..
| |
| 7 7,. ./../.a...y,f..+. 7. a =.l. o.o7F ;7 , .r.a..
| |
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| |
| .I..,...i, .. /.I"7 ../..!
| |
| l i #
| |
| /, /.i/'4/.. /. /. .. .. .
| |
| 8 I
| |
| /* i/ ' / '
| |
| d,2.0. sd 101
| |
| \
| |
| sa' . '
| |
| 6K (ksi /in.) .
| |
| 1
| |
| \
| |
| l I
| |
| 1 l
| |
| Figure 4-3. Fatigue Crack Growth Rate Curve for Austenitic Stainless Steel
| |
| . l seassesisee to 4,,7 i
| |
| | |
| 1
| |
| . 1 3
| |
| C F S E AK .30 d
| |
| where
| |
| = Crack Growth Rate in micro-inches / cycle d
| |
| -20 C= 2.42 x 10 F = frequency factor (F = 1.0 for temperature below 800*F)
| |
| S= R ratio correction (S = 1.0 for R = 0; S = 1 + 1.8R for ,
| |
| O < R < .8; and 8 = -43.35 + 57.97R for R >
| |
| 0.8) l =
| |
| E Environmental Factor (E = 1.0 for PWR)
| |
| AK = range of stress intensity factor, in psi /in and R is the ratio of the minimum Kg (bmin) to the .
| |
| maximum Kg (Kg ).
| |
| 1 Figure 4-4. Fatigue Crack Growth Equation for Austenitic Stainless Steel nu..euw. i.
| |
| 4-8 i
| |
| | |
| a e.
| |
| c, a
| |
| ie l I I l
| |
| | |
| SECTION 5.0 l REASSESSMENT OF LEAK-BEFORE-BREAK 5.1 Introduction Leak-before-break evaluations were performed for the pressurizer surge line of the Comanche Peak Unit I nuclear power plant in April, 1988 (reference 5-1).
| |
| The report was then submitted to the Nuclear Regulatory Commission (NRC). The NRC subsequently raised concerns about the surge line thermal stratification I issue. This section addresses leak-before-break as related to this issue. ,
| |
| The ingredients of the leak-before-break methodology are reviewed in table 5-1. I'tems 2, 3 and 8 are addressed in sections 2.0, 1.0 and 4.0 of this report, respectively. This section addresses the remaining items. The conservatism used in this section are listed in table 5-2.
| |
| 5.2 Material Properties Applicable material' properties were developed from those in the Certified Materials Test Report as given in table 5-3. The ASME code minimum properties ,
| |
| I are also given in table 5-3. It is seen that the measured properties well exceed those of the code. As seen later properties at [ ]a,c.e and 653'F are required for the leak rate and stability analyses.
| |
| I Industry data at 650*F were used as a basis for determining tensile properties at 653*F. Data for SA376 TP316 stainless steel pipe and welds are given in table 5-4 taken from reference 5-2. Data for Plant 3 in table 5-4 are quite similar to the Comanche Peak Unit 1 piping data in table 5-3. By maintaining ,
| |
| a constant ratio of properties at room temperature and 650'F, the 653'F (i.e.
| |
| 650'F) properties for the surge line material were estimated. The properties at ( la.c.e were obtained by maintaining the same ratio as those given . 1 in the ASME Code (reference 5-3). The modulus of elasticity at [ Ja,c.e were obtained from reference 5-3. The stress strain curves at 653*F were l obtained from the Nuclear Sy. stem Materials Handbook (reference 5-4). The l modulus so determined was used for consistency. All the tensile properties nu.+ine in 5-1 l
| |
| )
| |
| | |
| 4 are given in table 5-5. Both the required average and minimum properties are given. The average tensile properties were obtained from eighteen SA376 TP 316 materials including the Comanche Peak Surge line pipe. -
| |
| The average and minimum true stress-true strains curves at 653*F are given in figures 5-1 and 5-2. The average stress strain curve at [ Ja,c.e was -
| |
| obtained from multiple tensile tests and based on the average properties given .
| |
| in table 5-5. The curve is given in figure 5-3.
| |
| 5.3 Loading Conditions Because thermal stratification can cause large stresses at heatup an'd cooldown temperatures in the range of 455'F, a review of stresses was used to identify the wor'st situations for LBB applications. The loading states so identified are given in table 5-6. The procedure for determining the worst loading situations is explained in section 5.6. One location, node 1010, as shown in figure 5-4, was found to be the most critical for the LBB ev~aluation. At node 1010 there is a GTAW field weld. The weld types at other locations are also indicated in figure 5-4.
| |
| Five loading cases were identified for LBB evaluation as given in table 5-7.
| |
| Cases A and B are cases for leak rate calculations with the remaining cases the corresponding faulted situations for stability evaluations. The loads at the critical location for the five cases are given in table 5-8. The four case combinations for the leakage and stability evaluations are listed in table 5-9. Details of the evaluations for the four cases are given in table 5-;0.
| |
| 5.4 Leak Rate Calculation .
| |
| Leakage through postulated cracks were calculated using a procedure similar to ,
| |
| that discussed in reference 5-1. The resulting leakage flaws (i.e., the flaw sizes giving 10 gpm) are given in table 5-11. The norr.a1 loads for cases A ,
| |
| and B were used noting that the temperature was [ la,c.e Actually the smaller leakage flaw associated with normal operation is more relevant since the normal operating stratification surges would lead to leakage detection for -
| |
| the smaller leakage flaw (i.e., [ la.c.e
| |
| - "** 5-2 L _ ___ _ ----- _ --- _ -- _ _--_ _ ___-------_--- -
| |
| | |
| 5.5 Reactor Coolant System (RCS) Cooldown Stratification Temperature
| |
| , Considerations
| |
| . This section provides additional information for the postulated leak-before-break (LBB) cases [ Ja,c.e These are the cases in which [
| |
| Ja,c.e . The specific area of interest is the maximum pipe AT (Pipe Top-to-Bottom Temperature) that is postulated to occur during the cooldown.
| |
| Two factors are used to predict the pipe AT; the system AT (pressurizer temperature minus hot leg temperature), and the ratio of pipe AT to system AT (measured during testing).
| |
| 5.5.1 Reactor Coolant System Temperatures The postulated leak rate for [ ]a,ce would not be expected to cause the activation of any safety injection systems. Therefore, the generic Westinghouse procedure " Plant Shutdown from Minimum Load to Cold
| |
| ^
| |
| Shutdown" would apply. This procedure and information from Comanche Peak Unit 1 procedures an~d operator int'erviews will be the basis for the discussions which follow.
| |
| A plot of the pressurizer water temperature and the reactor coolant water temperature during a typical normal cooldown is given in figure 5-5. Note that figure 5-5 shows both the water solid and steam bubble modes of cooldown. For the steam bubble mode cooldown, the pressurizer initially contains a steam bubble with the water level at 25% of level spian (no load conditions). Early during the cooldown procedure, all but one of the reactor coolant pumps are stopped. The operating pump is in the loop to which the pressurizer surge line and a spray line are connected.
| |
| In general, the process can be divided into three phases; first the -
| |
| pressurizer water and reactor coolant system water are cooled down together with the reactor coolant temperature maintained approximately 50 to 100*F below the pressurizer saturation temperature. This temperature difference is mandated by subcooling requirements of the RCS. When the reactor coolant nu,w ae io 5-3
| |
| | |
| pressure and temperature have decreased to approximately 325 psig and 350'F, the residual heat removal system is placed in operation. The pressurizer ,
| |
| pressure remains constant at the pressure required to operate the reactor
| |
| . coolant pumps while the reactor coolant temperature continues to decrease to .
| |
| 160*F. When the reactor coolant system temperature has deceased to 160*F, the ,
| |
| operating reactor coolant pump is stopped. The system is depressurized using auxiliary spray. The steam bubble in the pressurizer is collapsed when the pressure has been decreased .to 25 psig.
| |
| Based on the above discussion, the maximum temperature difference between pressurizer water and hot leg water is 265'F, that is, the difference between the saturation temperature of 425'F corresponding to 325 psig and th's reactor coolant system temperature of 160'F. This temperature difference is the maximum potential temperature difference and is considered (when corrected to pipe versus system AT) in the stress and fatigue analysis.
| |
| The case under question is (the postulated scenario of a 1 gpm leak being detected at full power and the subsequent cooldown to locate and repair the leak ( ,
| |
| Ja,c e For this case a more realistic postulation of RCS conditions is assumed.
| |
| Per Comanche Peak Unit 1 procedures, the leak will be located while the plant is in mode 3, or time 0-4 hours on figure 5-5. If a leak in the RCS was detected the first priority would be to depressurize the system. This depressurization would most likely occur when the RCS is at approximately 180*F. Therefore, the maximum expected system temperature difference would be between ( Ja c.e . For the analyses of ( ,
| |
| la,c.e the system temperature difference is assumed to be [ Ja c.e 5.5.2 Pipe Versus System Temperature Difference The maximum expected pipe AT (Pipe T Top - Pipe TBottom) is a function of the system AT.
| |
| 1 Thermal hydraulic considerations and actual monitoring data indicate that, for the phase of the cooldown considered herein, the pipe AT will always be less 1 _ . _ . .
| |
| 54 l
| |
| l
| |
| | |
| I i
| |
| than the system AT. The phase of cooldown under investigation occurs after the reactor has achieved cold shutdown status. Data from the entire plant cycle (as available) were considered even though there was significantly.less theimal activity observed during plant cooldowns. This investigation is based on monitored thermal transient information from ( Ja,c,e plants. The number of thermal transients considered significant in this investigation was I
| |
| [ la,c.e These transients are provided in table 5-12.
| |
| The mean (x) cf the ratio of pipe AT to system AT was determined to be
| |
| [ la,c e The maximum range of the data showed that'the ratio of pipe f.T to system AT varied from (
| |
| la,c e It should be noted that a significant number of thermal l transients were observed with pipe AT to system AT ratios much lower than I
| |
| ( la,c.e These transients were excluded from consideration for conservatism.
| |
| 5.5.3 Conclusions j 1
| |
| l The pipe AT is determined from the system AT in section 5.5.1 and the ratio of the pipe to ~ system AT in section 5.5.2. The maximum system AT is
| |
| [ la,c.e based on the arguments of section 5.5.1. The ratios of pipe to
| |
| ~
| |
| system AT were between [ la.c.e Therefore, the magnitude of pipe AT's expected for [
| |
| ]a,c.e The pipe AT used in the analyses of ( la.c.e
| |
| ~
| |
| Additional data obtained to date from actual plant cooldown supports the above position. Figure 5-6 (roference [ la,c.e) shows the pressurizer and RCS hot leg temperatures vs. time for two consecutive cooldowns. The maximum system AT was [ la,c.e; the pipe AT was not monitored during this !
| |
| cooldown. The first cooldown was due to a weld leak on the seal injection .
| |
| line and is typical of the type of response expected for a small leak that did not require a safety injection. The second cooldown was for repairs on the !
| |
| secondary side of a steam generator manway gasket (also detected as a leak in i
| |
| containment).
| |
| l nu. m ue is 5-5 :
| |
| | |
| (
| |
| Figure 5-7 (reference ( )]a,c.e illustrates the system temperatures from another actual plant cooldown. The maximum system AT obtained was ,
| |
| [ Ja,c.e Pipe AT's were also obtained-for this period and are shown in figure 5-8 along with the system AT. The maximum pipe AT's for this- ,
| |
| cooldown was less than'( Ja,c,e _
| |
| 5.6 Evaluation of Flux Welds Stability evaluations were performed at node 1010 using the ASME Section XI IWB 3640 procedure (reference 5-5). This procedure uses the limit load methodology with a correction factor related to the type of weld. This-procedure was applied at all weld locations in establishing the crit'ical locations to be node 1010. The flaw sizes for instability are also given in table 5-11.
| |
| 5.7 Results The leakage flaws and instability flaws are given in table 5-11. The margins (ratio of instability flaw to leakage flaw) are also given. The margins for
| |
| ~
| |
| all cases well exceed the factor o'f 2.
| |
| ~
| |
| The conclusions of the evaluations of this section are:
| |
| : 1. LBB exists at operating temperature without stratification
| |
| : 2. LBB exists at operating temperature with stratification
| |
| : 3. LBB exists for forced cooldown due to leakage In summary, LBB is reconfirmed for the surge line subjected to thermal stratification.
| |
| 5.8 References '
| |
| 5-1 CPSES-1 WHIPJET Program Report, Comanche Peak Steam Electric Station, -
| |
| Unit Number 1, Robert L. Cloud and Associates, Inc. April 1988.
| |
| O numim io 5-6
| |
| | |
| 5-2 Witt, F. J. et. al., Integrity of the Primary Piping Systems of -
| |
| Westinghouse Nuclear Power Plants During Postulated Seismic Events, WCAP-9283, Westinghouse Electric Corp., March 1978, p 3-3. ;
| |
| 5-3 1986 ASME Boiler and Pressure Vessel Code, Nuclear Power Plant Components ~
| |
| Division 1 Appendices.
| |
| , 5-4 Nuclear Systems Materials Handbook, Part I - Structural Materials, Group I 1 - High Alloy Steels, Section 2, ERDA Report TID 26666, November 1975 Revision.
| |
| 5-5 ASME Code Section XI, Winter 1985 Addendum, Article IWB-3640.
| |
| l l
| |
| l l
| |
| G 9
| |
| m e
| |
| m .m m ,o 5.y -
| |
| | |
| TABLE 5-1 STEPS IN A LEAK-BEFORE-BREAK ANALYSIS '
| |
| (1)
| |
| (2) Establish material properties including fracture toughness value!.
| |
| Perform stress analysis of the structure {
| |
| (3) Review operating history of the plant (4)
| |
| Select locations for postulating flaws .
| |
| (5)
| |
| Determine a flaw size giving a detectable leak rate (6)
| |
| Establish stability of the selected flaw -
| |
| {
| |
| (7) {
| |
| Establish adequate margins in terms of leak rate detection, flaw size and load. k (8)
| |
| Show that a flaw indication acceptable by inspection remains s throughout service life, 4
| |
| 4 e
| |
| j 9
| |
| J e
| |
| MN12ee 10 ,
| |
| 5-8
| |
| | |
| 1 TABLE 5-2 LBB CONSERVATISM o
| |
| o Factor of 10 on Leak Rate Factor of 2 on Leakage Flaw -
| |
| o .
| |
| o Algebraic Sum of Loads for Leakage -
| |
| ~ o Absolute Sum of Loads for Stability o Average Material Properties for Leakage V Minimum Material Properties for Stability I
| |
| l I
| |
| i i
| |
| I
| |
| . i 1
| |
| 1 1
| |
| 1 1
| |
| e s h
| |
| a 4
| |
| 0 0
| |
| 5-9
| |
| | |
| e.
| |
| c, a
| |
| F O
| |
| E SH ET I
| |
| TF t ROT n ,
| |
| E N e PSA m ODL e RLP r PE i W1 u L q ADT e 3CNI R
| |
| - I AN 0 0 0 5N U .
| |
| m 0 0 0 AS EHLK u
| |
| m 0, 0, 0, LCAA i 5 0 0 BEIE n 7 8 8 AMRP i T E M ETE RAH e UMC d T N o AEA C 0 0 RNM 0 0 EIO PLC E
| |
| M 0, -
| |
| 0, M S 0 5 EE A 3 3 TG R
| |
| MU N OS 6 O 1 6 R 3 1 P 3 T F 6 2 7 8 8 3 0 1 A 3 A S E S l
| |
| e e d z p l z i e o P W N
| |
| | |
| TABLE 5-4 TYPICAL TENSILE PROPERTIES OF SA376 TP316 AND WELOS OF SUCH MATERIAL FOR REACTOR PRIMARY COOLANT SYSTEMS Test Temperature Average Tensile Properties Plant Material
| |
| ('F)
| |
| ~
| |
| Yield (psi) Ultimate (psi) 1 SA376 TP316 70 40,900 (48)" 83,200 (48)
| |
| ~
| |
| , 650 23,500 (19) 67,900(19)
| |
| E 308 Weld 70 63,900(3) 87,600(3) 2 SA376 TP316 70 47,100(40) 88,300(40) 650 26,900 (22) 69,100(25)
| |
| E 308 Weld 70 59,900(8) 87.200(8) 650 31,500(1) 68,800(1) 3- SA376 TP316 70- 46,600(36) 87,300(36) 650 24,200 (18) 66,800(19)
| |
| E 308 Wald 70 61,900(4) 85,400 (4)
| |
| : a. (_)indicatesthenumberoftestresultsaveragedobtainedfrom
| |
| - f Certified Materials Test Report of the primary coolant system of a plant.
| |
| 1 .
| |
| t 3986%1MG 10 5-11
| |
| _ - - _ - - - _ _ _ _ _ - _ _ - - - - - - - - - - - - - - - - - - - - - - - - - - - --- - ----- ------ - - -- --- --- ~ - - - ^
| |
| | |
| I TABLE 5-5 !
| |
| i TENSILE PROPERTIES FOR THE SURGE LINE M'ATERIAL AT ( Ja,c,e and 653'F .
| |
| Yield Stress Ultimate Strength Modulus of _
| |
| Temperature (psi). (psi) Elasticity 0
| |
| (*F) Average. Minimum ~ Average Minimum (psix10 )
| |
| 70a ,b- 44300 43300 87200 86600 ja.c.e
| |
| [
| |
| 653 [ .Ja.c.e a
| |
| Average room temperature results from 18 SA376 TP316 materials including surge line materiais.
| |
| b Minimum values from table 5-3.
| |
| b e
| |
| j e
| |
| m..,o.im io 5-12 L .
| |
| | |
| TABLE 5-6
| |
| - TYPES OF LOADINGS ,
| |
| Pressure (P)
| |
| Dead Weight (DW)
| |
| Normal Operating Thermal Expansion (TH)
| |
| Safe Shutdown Earthquake and Seismic Anchor Motion (SSE)a 3a,c.e
| |
| [
| |
| 3a,c.e
| |
| [
| |
| . ~
| |
| aSSE is used to refer'to the absolote suta of these loadings.
| |
| 6 i
| |
| i 9
| |
| 1 e
| |
| ...ime io 5-13
| |
| | |
| TABLE 5-7 NORMAL AND FAULTEL LOADING CASES FOR LBB EVALUATIONS CASE A: This is the normal operating case at 653'F consisting of the algebraic sum of the loading components due to P, DW and TH. _
| |
| (
| |
| ^
| |
| 3a,c.e CASE D: This is the faulted operating case at 653*F consisting of the absolute sum (overy component load is taken as positive) of P, DW, TH and SSE. l
| |
| [
| |
| ja.c.e
| |
| (
| |
| 3a,c.e .
| |
| e O
| |
| m...m. in 5-M
| |
| | |
| TABLE 5-8
| |
| | |
| ==SUMMARY==
| |
| OF LOADS AND STRESSES AT THE CRITICAL LOCATION Force Stress Moment Stress Total Node Case F (1bs) F (psi) M (in-lbs) og (psi) Stress (psi)
| |
| ~
| |
| 8 7293 11751 1010 A 222878 4458 1069639 1010 [ Ja,c.e
| |
| , 1010 D 240526 4811 3436696 23433 28244 1010 [ Ja,c.e 1010 [ Ja,c.e aDimensions: 0.D. = 14 in., minmum wall thickness = 1.249 in.
| |
| b Ja,c.e Stratification AT is [
| |
| 1 i
| |
| e 4
| |
| b i
| |
| f
| |
| : 2. .. u.. , .
| |
| 5-15
| |
| | |
| f
| |
| ' TABLE 5-9
| |
| . ASSOCIATED LOAD CASES FOR ANALYSES- . ,
| |
| A/D' 'This is here-to-fore standard leak-before-break evaluation. ..
| |
| [
| |
| ja C,9
| |
| [
| |
| 3a,c.e
| |
| [
| |
| 3a.c.e e
| |
| E O
| |
| 9 m i i.
| |
| 5 ,
| |
| | |
| TABLE 5-10
| |
| . LOAD CASES, LOCATION AND TEMPERATURES CONSIDERED FOR LEAK-BEFORE-BREAK EVALUATIONS
| |
| ~
| |
| Temperature ('F) _
| |
| Case Node Leak Rate Stability A/D 1010 6bs 653
| |
| [ 3a,c.e 1010 [ 3a,c.e
| |
| [ 3a,c.e 1010 [ Ja,c.e
| |
| [ jac.e 1010 [ Ja c e 9
| |
| 6 e
| |
| S e
| |
| 1 m..in. io 5-17
| |
| | |
| TABLE 5-11 LEAKAGE FLAW SIZES, CRITICAL FLAW SIZES AND MARGINS Location of Smallest Critical Flaw _
| |
| ,- Lead Critical Flaw Size Size Based On
| |
| ~
| |
| a
| |
| ( Case Based on IWB-3640 Calc. IWB-3640 Leakage Flaw Margin
| |
| ~
| |
| A/D 1010 12.4 1.60 2.7
| |
| [ ja,c.e 1010 [ 3***''
| |
| ( 3a,c.e 1010 [ ]
| |
| [ 3a,c.e 1010 [ 3'*''
| |
| 8 "GTAW Weld e
| |
| W e
| |
| J B
| |
| w e
| |
| n ..im in 5-18
| |
| | |
| TABLE 5-12 SIGNIFICANT THERMAL TRANSIENTS
| |
| ' ~ a ,c.e
| |
| [ -1
| |
| . Pipe AT System AT %
| |
| a,c.e i
| |
| [ 3a,c.e l
| |
| Pipe AT System AT %:
| |
| a,c.e l
| |
| .l 1
| |
| I
| |
| . 4
| |
| =
| |
| 1 no...ime io 5-19
| |
| | |
| WESTINGHOUSE PROPRIETARY CLASS 2 TABLE 5-12 (cont.)
| |
| SIGNIFICANT THERMAL TRANSIENTS ,
| |
| 3 c.
| |
| [ 3a c.e Pipe aT System AT %
| |
| J
| |
| ~~ -
| |
| a,c.e -
| |
| 1 1
| |
| [ 3a c.e Pipe AT System AT %
| |
| ~
| |
| a,C,9 l
| |
| ~
| |
| 1 1
| |
| 0 b
| |
| a
| |
| : m. ine in 5-20
| |
| | |
| a,c.e 4
| |
| 1
| |
| -]
| |
| l i l j
| |
| 1 Figure 5-1. Average True Stress-True Strain Curve for the surge.line ,
| |
| SA376 TP316 Stainless Steel at 653*F l
| |
| n .. = o 5-21 1
| |
| . _ . . _ . _ _ . _ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ . . _ _ _ _ _ _ . _ _ _ _ _ _ _ . _ . _ _ _ _ _ _ . _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . . _ . _ _ _ _ _ _ _ . _ _ _ . _ _ _ _ _ . _ . _ . _ _ _ _ _ . . _ _ _ _ ]
| |
| | |
| B,Cse ~
| |
| da 1
| |
| . l 1
| |
| l l
| |
| I l
| |
| ~
| |
| Figure 5-2. Minimum True Stress-True Strain Curve for the surge line SA376 TP316 Stainless Steel at 653*F
| |
| .. min .io 5-22
| |
| | |
| d
| |
| - 1
| |
| 'I a,c.e -f
| |
| ~*
| |
| ~
| |
| .{
| |
| i l
| |
| l l
| |
| i 1
| |
| Figure 5-3. Average True Stress-True Strain Curve for the surge line..
| |
| SA376 TP316 Stainless Steel at (205'F]a,c.e
| |
| . .j m .mineaa
| |
| : j. 5-23 .!
| |
| i l
| |
| | |
| l 1
| |
| e : Field Welds (GTAW) !
| |
| ~
| |
| o : Shop Weld (SMAW) -
| |
| "" I "
| |
| ==. : Critical Location O
| |
| i I
| |
| l 3
| |
| 9 !
| |
| l
| |
| * i u.
| |
| 2010
| |
| ~
| |
| 1170 ggg 1140 l
| |
| Il 2 ,
| |
| NORTH .
| |
| CIRCLED NOS. ARE ANSYS NOCE PQ1NTS i
| |
| Figure 5-4. Sketch of Analysis Model for Comanche Peak Unit 1 Pressurizer l I
| |
| Surge Line Showing Node Points, Critical Locations, Weld Locations and Types of Welds m .. u u is 5-24
| |
| | |
| -~ - - - - - - - - - , , - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ ____ ___ __
| |
| 1 l
| |
| O I _
| |
| 4 i
| |
| .e 9
| |
| l l 1 1
| |
| c 'k 8
| |
| m 8
| |
| u 1 0
| |
| - l l
| |
| e )
| |
| O 4
| |
| L !'
| |
| 3
| |
| .O.
| |
| b 4
| |
| J l 2 I
| |
| a 9
| |
| 5-25
| |
| | |
| l!lllll11lIll.
| |
| i11 8,
| |
| C, 8 ,
| |
| ( .
| |
| ~
| |
| 3 k a
| |
| e .
| |
| L S
| |
| C R
| |
| n a
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| t o
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| e u
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| D y n w
| |
| o d
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| l o
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| o C
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| l a
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| u t .
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| c A
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| e.
| |
| : c. .
| |
| a l
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| [
| |
| 6 5 .
| |
| e r
| |
| ' u g
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| i F .
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| e e
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| e ~ e,
| |
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| I I
| |
| u l o
| |
| N
| |
| +
| |
| =
| |
| i i
| |
| 1 i
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| l i
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| 8 u
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| 3 ee U
| |
| *IC
| |
| @ l U.
| |
| N m
| |
| e a e
| |
| Q b
| |
| 4 40 W
| |
| L.
| |
| 3 er- ,
| |
| 1 l
| |
| u i
| |
| l I
| |
| O 5-27
| |
| | |
| e U.
| |
| i
| |
| * 1 i
| |
| I 1
| |
| W , 1 0
| |
| GO C
| |
| k i
| |
| \
| |
| Q- !
| |
| l
| |
| 'O C
| |
| 9 i 4
| |
| 40 h
| |
| M i
| |
| g v
| |
| .-4 C e O
| |
| .e-atmil 4
| |
| 8 a
| |
| G.
| |
| U.
| |
| N r==s e
| |
| W e
| |
| G e
| |
| ll)
| |
| W '
| |
| t D
| |
| D 1 4
| |
| La.
| |
| t 1 g
| |
| : s. - )
| |
| i 5-28
| |
| | |
| ~
| |
| SECTION 6.0 CONCLUSIONS
| |
| , Based on the monitoring and analysis results presented.in.the report the following conclusions are reached:
| |
| (a) The global structural and local stresses in the surge line piping and support system meet ASME III Code allowables. The maximum cumulative fatigue usage factor is [ Ja,c.e for 40 year design life, compared.
| |
| to the Coos allowable of 1.0.
| |
| (b) Fatigue crack growth (FCG) analyses show'that a postulated 10%' initial
| |
| . crack will not propagate beyond 20% of the pipe wall in'the 40 years d/signlife.
| |
| (c) Leak-Before-Break (LBB) is confirmed for all loading combinations, including maximum postulated stratification.
| |
| In summary, based on the current understanding of the thermal stratification phenomenon, it is con'cluded that thermal stratification does not~ affect the integrity (nor the ebility to fulfill." Leak-Before-Break" criteria)'of the O
| |
| pressurizer surge line of the Comanche Peak Unit i nuclear power plant. The design life (forty years) and ASME III Code compliance are not affected.
| |
| S o
| |
| h e
| |
| 9 mu. unao 6-1
| |
| | |
| APPENDIX A LIST OF COMPUTER PROGRAMS This appendix. lists and summarizes the computer codes used in the analysis of stratification in the Comanche Peak Unit 1 pressurized surge line. The codes are:
| |
| t 1. WECAN 1
| |
| : 2. WECEVAL
| |
| : 3. STRFAT2
| |
| : 4. ANSYS
| |
| : 5. FCG A.1 WECAN A.1.1 Description WECAN is a Westinghouse-developed, general purpose finite element program. It contains universally accepted two-dimensional and three-dimensional iseparametric elements that can be used in many different types of finite element analyses. Quadrilateral and triangular structural elements are ut,ed for plane strain, plane stress, and axisymmetric analyses. Brick and wedge structural elements are used for three-dimensional analyses. Companion heat conduction elements are used for steady state heat conduction analyses and transient heat conduction analyses.
| |
| A.1.2 Feature Used The temperatures obtained from a static heat conduction analysis, or at a specific time in a transient heat conduction analysis, can be automatically input to a static structural analysis where the heat conduction elements are replaced by corresponding structural elements. Pressure and external loads ,
| |
| . can also be inclu e in the WECAN structural analysis. Such coupled thermal-stress antlyses are a standard application used extensively on an ,
| |
| . industry wide basis.
| |
| : m. .* ""
| |
| * A-1
| |
| | |
| A.1.3 Program Verification Both the WECAN program and input for the WECAN verification problems, currently numbering over four hundred, are maintained under configuration control. Verification problems include coupled thermal-stress analyses for the quadrilateral, triangular, brick, and wedge isoparametric elements. These problems are an integral part of the WECAN quality assurance procedures. When ,
| |
| a change is made to WECAN, as part of the reverification process, the configured inputs for the coupled thermal-stress verification problems are .
| |
| used to reverify WECAN for coupled thermal-stress analyses.
| |
| A.2 WECEVAL A.2.1 Description WECEVAL is a multi purpose program which processes stress input to calculate ASME Section III, Subsection NB equations and usage factors. Specifically, the program performs primary stress evaluations, primary plus secondary stress intensity range analysis, and fatigue analysis for finite element models generated and run using the WECAN computer program. Input to WECEVAL consists -
| |
| of card image data and data evtracted from the output TAPE 12's generated by WECAN's stress elements. The program reads the input data, performs the necessary calculations, and produces summary sheets of the results.
| |
| The required stresses are read from the WECAN TAPE 12's and placed onto intermediate or restart files. The user may then catalog these files for use in later evaluations. The stress state for a particular loading condition is obtained by a ratio-superposition technique. This optimal stress state is formed by manipulating the signs of the applied loads to generate the largest possible stress magnitude. .
| |
| A.2.2 Feature Used
| |
| \
| |
| WECEVAL has many options and features which enhance its versatility. Among I those used for this evaluation were: ,
| |
| nu,m me "
| |
| A-2
| |
| : 1. The ability to perform simplified elastic plastic analysis per NB-3228.5, including the automatic calculation of Ke facters and removal of thermsl bending stresses from the maximum range of stress intensity evaluations.
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| : 2. Built-in ASME fatigue curves plus provisions for accepting user-defined fatigue curves. ~'
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| : 3. Equivalent moment linearization technique, along with the ability to correct for the radius effects in cylindrical and spherical geometries, i
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| : 4. The ability to limit the interactions among load conditions during the fatigue analysis, d
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| : 5. Generating input for the fatigue crack growth program FCG.
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| l A.2.3 Program Verification WECEVAL is verified to Westinghouse procedures by independent calculations of l ASME !!! NB Code equations and comparison to WECEVAL results, i A.3 STRFAT2 A.3.1 Description 4
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| i STRFAT2 is a program which computes the alternating peak stress on the inside surface of a flat plate and the usage factor due to striping on the surface.
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| The program is applicable to be used for striping on the inside surface of a pipe if the program assumptions are considered to apply for the particular j pipe being evaluated.
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| For striping the fluid temperature is a sinusoidal variation with numerous cycles. .
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| 3 The frequency, convection film coefficient, and pipe material properties are
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| . input.
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| l m ..i ie A-3
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| .The program computes maximum alternating stress based on the maximum difference between inside surface skin temperature and-the average through , ,
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| wall temperature.
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| 4 A.3.2 Feature used _
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| The program is used to calculate striping usage factor based on a ratio of' actual cycles of stress for a specified length of time divided by allowable ,
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| cycles of stress at maximum the alternating stress level. Design fatigue curves for several materials are contained into the program. However, the user has the option to input any other fatigue design curve, by designating that the fatigue curve is to be user defir.ed.
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| A.3.3 Program Verification STRFAT2 is verified to Westinghouse procedures by independent review of the stress equations and calculations.
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| A.4 ANSYS
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| * A.4.1 Description ANSYS is a public domain, general purpose finite element code.
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| A.4.2 Feature Used The ANSYS elements used for the analysis of stratification effects in the .
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| surge line are STIF 20 (straight pipe), STIF 60 (elbow and bends) and STIF14 (spring-damper for supports). .
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| A.4.3 Program Verification As described in section 2.1, the application of ANSYS for stratification has been independently verified by comparison to WESTDYN (Westinghouse piping]
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| w.wm io A-4 1
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| i 1
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| l analysis code) and WECAN (finite element code, section 8.1). The results from
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| . ANSYS are also verified against closed form solutions for simple beam !
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| configurations.
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| I A.5 FCG l l
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| , i' A.5.1 Description The FCG computer program models fatigue crack growth using linear elastic fracture mechanics methods. In order to provide a realistic model of crack growth the design transients which are input are automatically scheduled evenly over the life of the system or component. 1 l
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| A 5.2 Features Used I i
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| The program options enable calculation of crack tip stress intensity factors I (Kg ) for surface flaws and embedded flaws in a large number of coometries, under any leading condition. Crack growth results are determined for each year of operation, and summarized in tabular form at the end of the output, at
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| ~
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| 10' year intervals.
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| l
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| * \
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| A.S.3 Program Verification ;
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| ! The program has been verified by performing alternate calculations and placed under Westinghouse configuration control. The calculations using this program were presented and approved by the NRC staff in connection with several applications.
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| i m..wm ,o g.5 L}}
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