ML20199L620
| ML20199L620 | |
| Person / Time | |
|---|---|
| Site: | Arkansas Nuclear, 07201007  |
| Issue date: | 12/31/1973 |
| From: | Salvatori R WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML20199L511 | List: |
| References | |
| WCAP-8185, WCAP-8185-A01, WCAP-8185-A1, NUDOCS 9901280076 | |
| Download: ML20199L620 (250) | |
Text
{{#Wiki_filter:. . - . . . . . o EEEEEENCE (ORE REPORT IFmf7 VOttlHE 1 o Westinghouse Nudear Energy Systems O , p=:88a:;8apae
l l0 EEEEEENCE (GEE REFON 11x11 VOKEME1 i i l O . I Westinghetase Nudear Energy l Systems 1O - l ' l $3R'*8838184,8ag
TECHNICAL LIBRARY I PORTLAND GENERAL ELECTRIC COMPARI Westinghouse Non-Proprietary Class 3 E liii 5$ 5
$"E REFERENCE CORE REPORT t --..
17 X 17 O December,1973 APPROVED M R. Salvatori, Manager Nuclear Safety l l a WESTINGHOUSE ELECTRIC CORPORATION Nuclear Energy Systems O v P. O. Box 355 Pittsburgh, Pennsylvania 15230
I TABLE OF CONTENTS l J Section Title Page FOREWORD 1 l l l 1.5 REQUIREMENTS FOR FURTHER TECHNICAL INFORMATION 1.5-1 1.5.1 Verification Test (17 x 17) 1.5-2 1.5.1.1 Rod Cluster Control Spider Test 1.5-3 1.5.1.2 Grid Test 1.5-4 , 1.5.1.3 Fuel Assembly Structural Tests 1.5-4 1.5.1.4 Guide Tube Test 1.5-6 1.5.1.5 Prototype Assembly Test 1.5-8 1.5.1.6 Departure from Nucleate Boiling (DNB) 1.5-10 1.5.1.7 Incore Flow Mixing 1.5-11 1.5.2 Inpile Fuel Densification 1.5-12 1.5.3 LOCA Heat Transfer Test (17 x 17) 1.5-13 c 1. 5. 3.1 Verificati5n Facility Testing 1.5-14 2 1.5.3.2 D NB Test 1.5-16 1.5.3.3 Single Rod Burst Test (SRBT) 1.5-19 1.5.4 Westinghouse Test Engineering Laboratory Facility 1.5-21 1.5.4.1 Introduction 1.5-21 j 1.5.4.2 Test Loops and Equipment 1.5-23 ; 1.5.5 References 1.5-32 I O 1-1
TABLE OF CONTENTS Section Title Pg 4.0 REACTOR 4.1' SIM4ARY DESCRIPTION 4.1-1 4.2 MECHANICAL DESIGN 4.2-1 4.2.1 Fuel 4.2-2 4.2.1.1 Design Bases 4.2-2 { 4.2.1.2 Design Description 4.2-5 4.2.1.3 Design Evaluation 4.2-10 4.2.1.4 Tests and Inspections 4.2-24 4.2.2 Reactor Vessel. Internals 4.2-29 4.2.2.1 Design Basis 4.2-39 4.2.2.2 Description and Drawings 4.2-30
- 4.2.2.3 Design Loading Conditions 4.2-36 4.2.2.4 Design Loading Categories 4.2-37 4.2.2.5 Design Criteria Basis 4.2-39 4.2.3 ' Reactivity Control System 4.2-39 4.2.3.1 Design Bases 4.2-39 j 4.2.3.2 Design Description 4.2-45 4.2.3.3 Design Evaluation 4.2-64 4.2.3.4 Tests. Verification and Inspections 4.2-78 4.2.3.5 l Instrumentation Applications 4.2-82
~
4.2.4 References 4.2-84 l E 4.3 NUCLEAR DESIGN 4.3-1 4.3.1 Design Bases 4.3-1 4.3.1.1 Fuel Burnup 4.3-2 i 4.3.1.2 Negative Reactivity Feedbacks (Reactivity Coefficient) 4.3-3 4.3.1.3 Control of Power Distribution 4.3-4 4.3.1.4 Maximum Controlled Reactivity Insertion Rate 4.3-5 4-1 y v e ~.z -. .y- - -y , - - - , -w -
M TABLE OF CONTENTS (Continued) d
.. Section Title Page 4.3.1.5 Shutdown Margins 4.3-6 4.3.1.6 Stability 4.3-8 4.3.1.7 Anticipated Transients Without Trip 4.3-9 4.3.2 Description 4.3-9 4.3.2.1 Nuclear Design Description 4.3-9 4.3.2.2 Power Disteibutions 4.3-12 4.3.2.3 Reactivity Coefficients 4.3-26 4.3.2.4 Control Requirements 4.3-31 4.3.2.5 Control 4.3-34 4.3.2.6 Control Rod Patterns and Reactivity Worth 4.3-38 4.3.2.7 Criticality of Fuel Assemblies 4.3-40 4.3.2.8 Stability 4.3-41 4.3.2.9 Vessel Irradiation 4.3-48
^"''' "" " ~
C') 4.3.3.1 Fuel Temperature Calculations 4.3-50 4.3.3.2 Macroscopic Group Constants 4.3-51 4.3.3.3 Spatial Few-Group Diffusion Calculations 4.3-53 4.3.4 Changes 4.3-54 4.3.5 References 4.3-55 4.4 THERMAL AND HYDRAULIC DESIGN 4.4-1 4.4.1 ' Design Bases 4.4-1 4.4.1.1 Departure from Nucleate Boiling Design Basis 4.4-2 4.4.1.2 Fuel Temperature Design Basis 4.4-2 4.4.1.3 Core Flow Design Basis 4.4-3 4.4.1.4 Hydrodynamic Stability Design Bases 4.4-4 4.4.1.5 Other Considerations 4.4-4 4.4.2 Description 4.4-5
- 4. 4. 2.1 Summary Comparison 4.4-5 O
4-11
1 TABLEOFCONTENTS(Continued) Section Title Page 4.4.2.2 Fuel Cladding Temperatures 4.4-6 4.4.2.3 Critical Heat Flux Ratio or Departure from Nucleate Boiling Ratio and Mixing Technology 4.4-12 ; 4.4.2.4 Flux Tilt Considerations 4.4-23 l 4.4.2.5 Void Fraction Distribution 4.4-23 ! 4.4.2.6 Core Coolant Flow Distribution 4.4-24
]
4.4.2.7 Core Pressure Drops and Hydraulic Loads 4.4-24 4.4.2.8 Correlation and Physical Data 4.4-26 4.4.2.9 Thermal Effects of Operational Transients 4.4-29 4.4.2.10 Uncertainti.es in Estimates 4.4-30 4.4.2.11 Plant Configuration Data 4.4-33 4.4.3 Evaluation 4.4-34 4.4.3.1 . Core Hydraulics 4.4-34 4.4.3.2 Influence of Power Distribution 4.4-36 v 4.4.3.3 Core Thermal Response 4.4-38 4.4.3.4 Analytical Techniques 4.4-39 4.4.3.5 Hydrodynamic Instability 4.4-46 l 4.4.3.6 Temperature Transient Effects Analysis 4.4-48 4.4.3.7 Potentially Damaging Temperature Effects During Transients 4.4-49 I 4.4.3.8 Energy Release During Fuel Element Burnout 4.4-50 l 4.4.3.9 Energy Release or Rupture of Waterlogged Fuel Elements 4.4-51 4.4.3.10 Fuel Rods Behavior Effects from Coolant i Flow Blockage 4.4-51 4.4.4 Testing and Verification 4.4-53 l 4.4.4.1 Tests Prior to Initial Criticality 4.4-53 4.4.4.2 Initial Power and Plant Operation 4.4-54 l 4.4.4.3 Component and Fuel Inspection 4.4-54 lO 4-111
TABLEOFCONTENTS(Continued) Section Title a
,P,aSe_
t 4.4.2.2 Fuel Cladding Temperatures 4.4-6 4.4.2.3 Critical Heat Flux Ratio or Departure from Nucleate Boiling Ratio and Mixing Technology 4.4-12 , l 4.4.2.4 Flux Tilt Considerations 4.4-23 4.4.2.5 Void Fraction Distribution 4.4-23 , l 4.4.2.6 Core Coolant Flow Distribution 4.4-24 4.4.2.7 Core Pressure Drops and Hydraulic Loads 4.4-24 4.4.2.8 Correlation and Physical Data 4.4-26 4.4.2.9 Thermal Effects of Operational Transients 4.4-29 4.4.2.10 Uncertainties in Estimates 4.4-30 4.4.2.11 Plant Configuration Data 4.4-33 4.4.3 Evaluation 4.4-34 4.4.3.1 Core Hydraulics 4.4-34 4.4.3.2 Influence of Power Distribution 4.4-36 4.4.3.3 Core Themal Response 4.4-38 4.4.3.4 Analytical Techniques 4.4-39 4.4.3.5 Hydrodynamic Instability 4.4-46 4.4.3.6 Temperature Transient Effects Analysis 4.4-48 4.4.3.7 Potentially Damaging Temperature Effects During Transients 4.4-49 4.4.3.8 Energy Release During Fuel Element Burnout 4.4-50 { 4.4.3.9 Energy Release or Rupture of Waterlogged i Fuel Elements 4.4-51 l 4.4.3.10 Fuel Rods Behavior Effects from Coolant l Flow Blockage 4.4-51 4.4.4 Testing and Verification 4.4-53 l 4.4.4.1 Tests Prior to Initial Criticality 4.4-53 l l 4.4.4.2 Initial Power and Plant Operation 4.4-54 l i-4.4.4.3 Component and Fuel Inspection 4.4-54 .o 4-iii
- , . , , _ - , - . - _ . . -. - . - - -. 4
TABLEOFCONTENTS(Continued) Section Title M l l 4.4.5 Instrumentation Application 4.4-54 l l 4.4.5.1 Incore Instrumentation 4.4-54 1 4.4.5.2 Overtemperature and Overpower AT l . Instrumentation 4.4-55 4.4.5.3 Instrumentation to Limit Maximum Power Output 4.4-56 4.4.6 References 4.4-58 L l l l i I I O l 1 l l i l I l l I l l t 'O 4-iv
i l l LIST OF TABLES
.O Table Title l I'
4.1-1 A Reactor Design Comparison Table (FourLoopPlant) 4.1 -1 B Reactor Design Comparison Table (Three Loop Plant) 4.1.2 Analytic Techniques in Core Design 4.1.3 Design _ Loading Conditions For Reactor Core Components 4.2-1 Maximum Deflections Allowed For Reactor Internal Support Structures 4.3-1 Reactor Core Description 4.3-2 Nuclear Design Limits ) l 4.3-3 Reactivity Requirements for Rod Cluster Control Assemblies 4.3-4 Axial Stability Index PWR Core with a j 12 Ft Height i i 4.3-5 Typical Neutron Flux Levels at Full Power 4.3-6 Comparison of Measured and Calculated i Doppler Effects 4.3-7 Benchmark Critical Experiments O 4-v
, - . ~ ,
l l 7 LISTOFTABLES(Continued)
-( i Table Title L 4.3-8 Saxton Core II Isotopics
- Rod MY+, Axial Zone 6 t.
4.3-9 Critical Boron Concentrations 4.3-10 Comparison of Measured and Calculated Rod Worth 4.3-11 Comparison of Measured and Calculated Moderator Coefficients at HZP, BOL l 4.4-1 Reactor Design Comparison I 4.4-2 Thermal-Hydraulic Design Parameters For i One of Four Coolant Loop, Out of Service 4.4-3 Void Fractions at Nominal Reactor Conditions with Design Hot Channel Factors 4.4-4 Comparison of THINC-IV and THINC-I Predictions l with Data from Representative Westinghouse l Two and Three Loop Reactors I l 1 4.4-5 Comparison of HYDNA with Experimental Data l O ; [ 4-vi I j 3
. .. . .. ~. .- ~. . . _ - - - - -. - . ..
d l 1 LIST OF FIGURES l Figure Title 4.2-1 Fuel Assembly Cross Section (17 x 17) 4.2-2 ruel Assembly Outline (17 x 17) 4.2-3 Fuel Rod Schematic 4.2-4 Typical Clad and Pellet Dimensions as a Function of Exposure l
^
4.2-5 Representative Fuel Rod Internal Pressure
. and Linear Power Density for the Lead Burnup Rod as a Function of Time 4.2-6A Three Loop Lower Core Support Assembly Ox (CoreBarrelAssembly) 4.2-6B Four Loop Lower Core Support Assembly (CoreBarrelAssenbly) 4.2-7A Three Loop Upper Core Support Structure 4.2-7B Four Loop Upper Core Support Structure 4.2-8A Three Loop Plan View of Upper Core Support Structure 4.2 8B Four Loop Plan View of Upper Core Support Structure 4.2-9 Full Length Rod Cluster Control and Drive Rod fm Assembly with Interfacing Components i
4-vii
. . = - - . . . . . . - - . . - - . _ . - . -_ . - _ - .
LISTOFFIGURES(Continued) Figure Title 4.2-10 Full Length Rod Cluster Control Assembly Outline 4.2-11 Full Length Absorter Rod 4.2-12 Part Length Rod Cluster Control Assembly Outline 4.2-13 Burnable Poison Assembly (Conceptual) 4.2-14 Burnable Poisor. Rod - Cross Section 1 4.2-15A ThreeLoopSourceAssembly(Conceptual) 4.2-15B /w r Loop Source Assembly (Conceptual) i 4.2-15C Four Loop Secondary Source Assembly (Conceptual) ! 4.2-16 ThimblePlugAssembly(Conceptual) j 4.2-17 Full Length Control Rod Drive Mechanism j 4.2-18 Full Length Control Rod Drive Mechanism Schematic 4.2-19 Part Longth Control Rod Drive Mechanism 4.2-20 Nominal Latch Clearance at Minimum and Maximum Temperature O - 4-viii
LIST OF FIGURES (Continued) O , Fieure- Tit 1e 4.2-21 Control Rod Drive Mechanism Latch Clearance Themal Effect 4.3-1 Fuel Loading Arrangement l 4.3-2 Production and Consumption of Higher Isotopes l' 4.3-3 Boron Concentration VS First Cycle Burnup With-and Without Burnable Poison Rods l 4.3-4 Burnable Poison Rod Arrangement Within l An Assembly 1 l-4.3-5 Burnable Poison Loading Pattern 4.3-6 Nomalized Power Density Distribution Near Beginning of Life, Unrodded Core, Hot Full Power, No Xenon 4.3-7 Noma 11 red Power Density Distribution Near Beginning of Life Unrodded Core, Hot Full Power, Equilibrium Xenon 4.3-8 Normalized Power Density Distribution har l Beginning of Life, Group D 75% Inserted, Hot I. Full Power, Equilibrium Xenon l l 4.3-9 Nomalized Power Density Distribution Near i Beginning of Life, Group D 35% Inserted, Plus l PL Inserted Hot Full Power, Equilibrium Xenon ir o 4-ix
q LISTOFFIGURES(Continued) NJ Figure Title 4.3-10 Nomalized Power Density Distribution Near Middle of Life, Unrodded Core, Hot Full Power, Equilibrium Xenon 4.3-11 Nomalized Power Density Distribution Near End of Life, Unrodded Core, Hot Full Power. Equilibrium Xenon 4.3-12 Rodwise Power Distribution in a Typical Assembly (Assembly G-9) Near Beginning of Life, Hot Full Power, Equilibrium Xenon, Unrodded Core 4.3-13 Rodwise Power Distribution in a Typical Assembly (G-9) Near End of Life, Hot Full Power, Equilibrium O Xenon, Unrodded Core 4.3-14 Typical Axial Power Shapes Occurring at Start of Life 4.3-15 Typical Axial Power Shapes Occurring at Middle of Life 4.3-16 Typical Axial Power Shapes Occurring at End of Life 4.3-17 Comparison of Assembly Axial Power Distribution with Core Average Axial Distribution, Bank Slightly Inserted, Part length at Mid-Plane O 4-x
.. .-. . - . _ - .. .- - - - = . . . __ . _ - -
LISTOFFIGURES(Continued) O Figure Title 4.3-18 Fg Total Versus Axial Offset for Four Loop Plant for Condition II Events Which Involve Control Bank Motion 4.3-19 Fq Total Versus Axial Offset at Beginning of Life for Normal Operation 4.3-20 Fg Total Versus Flux Difference at End of Life For Normal Operation f 4.3-21 Comparison Between Calculated and Measured Relative Fuel Assembly Power Distribution 4.3-22 Comparison of Calculated and Measured Axial Shape 4.3-23 Measured Values of F g for Full Power Rod Configurations 4.3-24 Doppler Temperature Coefficient at Beginning of Life and at End of Life, Cycle 1 4.3-25 Doppler-Only Power Coefficient, Beginning of Life and at End of Life, Cycle 1 4.3-26 Doppler-0nly Power Defect at Beginning of Life and at End of Life, Cycle 1 4.3-27 Moderator Temperature Coefficient at Beginning - of Life, Cycle 1, Ne Rods O 4-xi
p LISTOFFIGURES(Continued) N Figure Title l R.3-28 Moderator Temperature Coefficient at End of Life, Cycle 1 4.3-29 Moderator Temperature Coefficient as a Function of Boron Concentration at Beginning of Life, Cycle 1 No Rods 4.3-30 Hot Full Power Moderator Temperature Coefficient During Cycle 1 for the Critical Boron Concentration 4.3-31 Total Power Coefficient at Beginning of Life and at End of Life, Cycle 1 4.3-32 Total Defect at Beginning of Life and End of Life, Cycle 1 4.3-33 Rod Cluster Control Assembly Pattern 4.3-34 Accidental Simultaneous Withdrawal of Two Control Banks EOL, HZP Banks B and D Moving in the Same Plane, PL at 140 Steps 4.3-35 Design - Trip Curve 4.3-36 Nomalized Rod Worth Versus Percent Insertion, All Rods But One j 4.3-37 Axial Offset versus Time, PWR Core with a i 12-Ft. Core Height and 121 Assemblies O 4-xii
I l (q/ LISTOFFIGURES(Continued) Figure Title 4.3-38 XY Xenon Test Thermocouple Response l Quadrant Tilt Difference versus Time l 4.3-39 Calculated and Measured Doppler Defect and Coefficients at Beginning of Life, Two-Loop Plant,121 Assen611es,12 Foot Core 4.3-40 Comparison of Calculated and Measured Boron Concentration for 2-Loop Plant, l 121 Assemblies, 12 Foot Core l 4.3-41 Comparison of Calculated and Measured CB 2-Loop Plant,121 Assemblies,12 Foot Core 4.3-42 Comparison of Calculated and Measured CB in 3-Loop Plant,157 Assemblies,12 Foot Core i i 4.4-1 Fuel Temperature as a Function of Linear Power at Beginning of Core Life 4.4-2 Fuel Temperature as a Function of Linear Power at End of Core Life 4.4-3 Thennal Conductivity of UO2 (Data Corrected to 95% Theoretical Density) 4.4-4 Axial Variation of Average Clad Temperature for Rod Operating at 5.43 KW/FT 4.4-5 Comparison of All "R" Grid Data for Typical Cell , 4-xiii
LISTOFFIGURES(Continued) Figure Title 1 4.4-6 Comparison of All "R" Grid Data for Thimble Cells 4.4-7 Probability Distribution Curve 4.4-8 TDC versus Reynolds Number for 26 inch Grid Spacing 4.4-9 Normalized Radial Flow and Enthalpy Distribution at 4-Ft Elevation 4.4-10 Nomalized Radial Flow and Enthalpy Distributions at 8-Ft Elevation j 4.4-11 Normalized Radial Flow and Enthalpy Distribution at 12-Ft Elevation - Core Exit 4.4-12 Void Fraction versus Thermodynamic Quality H -H SAT
/H g -H SAT 4.4-13 PWR Natural Circulation Test
~
4.4-14 Comparison of a Representative Westinghouse 2-Loop Reactor Incore lhermocouple Measurements ! with THINC-IV Predictions i 4.4-15 Comparison of a Representative Westinghouse j 3-Loop Reactor Incore Thermocouple Measure-ments with THINC-IV Predictions l lO ' 4-xiv
_ _ _ . . . _ _ _ . . . . . ___ _ _ _ _ . . _ _ _ _ _ _ . _ _ _ _ _ _ _ . . . _ _ . ~ . _ . _ _ . _ i LIST OF FIGURES (Continued) l O Fiaure Tit 1e 4.4-16 Hanford Subchannel Temperature Data Comparison with THINC-IV f 4.4-17 Hanford Subcritical Temperature Data
. Comparison with THINC-IV i
4.4-18 HYDNA Results Indicating Power Level at which Flow Oscillations Occur 4.4-19 Hydrodynamic Flow Instability Study, Nomal Power = 3250 W 4.4-20 Parallel Channel Test Station 4.4-21 Experimental Flow Stability Data 4.4-22 Distribution of In-Core Instrumentation l LO l 4-xy > l
,n. ._ __ _ - ,_
FOREWORD This Topical Report is to serve as a Design Report on the Westinghouse , ! Pressurized Water Reactors which employ a new fuel assembly design. l The new fuel assembly design is characterized by a 17 x 17 array
- of fuel rods and guide thimbles which make up the fuel assembly.
Prior to the change-over to this design, which was announced early in 1973, Westinghouse three- and four-loop PWR's employed fuel assemblies characterized as 15 x 15. l In April of 1973, Westinghouse published Amendment 3 to RESAR Revision 3 (RESAR being the Westinghouse Reference Safety Analysis Report), which provided PSAR data on the 17 x 17 design. The RESAR Amendment 3 included Section 1.5 on the R & D and Testing Program associated with the new fuel design; a revised Chapter 4 which provided preliminary drawings and parameters for the mechanical, nuclear and thermal-hydraulic design features of the 17 x 17 core; and a revised Chapter 15 which described preliminary analysis of the transient behavior. The information contained , therein has been updated in Amendment 5, which responds to questions raised in review. It is appropriate for Westinghouse reference four-loop plants rated at 3411 MWt core power. l Certain plants currently under construction plan to employ l 17 x 17 fuel. The FSAR's for these plants will be amended when the final design and transient analyses are available. Currently, these FSAR's are based on the 15 x 15 fuel design. In general, FSAR amendments for 17 x 17 fuel are to be filed in the spring of 1974. Therefore, to aid in the regulatory review in the time span between the April,1973, RESAR preliminary design j l and the submittal of the respective FSAR amendments, Westinghouse i issued WCAP 8185 in September 1973 and this Amendment 1 to the Reference 17 x 17 Core Report. This amendment provides an l updated Section 1.5 on R & D, Testing and Verification; an updated Chapter 4 which incorporates text changes responding to information
-I-
j requested by the AEC during their review of RESAR Revision 3 and ' the original September topical report; a transient analysis Chapter 15 and technical specifications Chapter 16 for four loop plants, three loop plants and for the ice condenser plants with an upper head injection (UHI) system. Chapter 4 presented herein, is to be considered as a generic description of a four-loop plant (3411 MW core power) and a three-t loop plant (2652 MW corepower). Chapter 4 information for UHI t system plants is contained in the FSAR's for those plants. Where appropriate to illustrate the effects of the fuel design change, these parameters are comphred with those of the respective four-and three-loop 15 x 15 designs presented in the Trojan FSAR (Docket ! No. 50-344) and the Beaver Valley Unit 1 FSAR (Docket No. 50-334). I In general, the core parameters described herein are applicable on a reference basis to any Westinghouse PWR with the same core power rating. Differences in the reactor internals design or thermal hydraulic or other characteristics of a specific plant will, of course, be addressed in the specific SAR's. Similarly, any special requirements for power shaping or control (such as required for ECCS performance in plants with Ice Condenser Containment and 15 15 x 15 fuel assemblies) which are not accommodated by the margins inherent in the 17 x 17 core design are addressed in this Amendment and subsequently in the specific SAR amendments. i l The material presented in this report complies with the " Standard l Format and Content of Safety Analysis Reports for Nuclear Power l Plants, (Revision 1)", issued October,1972 including section numbering. Where references are given to material other than those included herein, it is to be understood that this, in general, refers to RESAR 3. However, if this Topical Report is incorporated by reference in a SAR for a specific plant, then references in this report to other Chapters are understood to be those of the SAR of that plant.
-II-l l
TABLE OF CONTENTS Section Title Page FOREWORD' 1 1.5 REQUIREMENTS FOR FURTHER TECHNICAL INFORMATION 1.5-1 1.5.1 Verification Test (17 x 17) 1.5-2 1.5.1.1 Rod Cluster Control Spider Test 1.5-3 1.5.1.2 Grid Test 1.5-4 1.5.1.3 Fuel Assen61y Structural Tests 1.5-5 ; 1.5.1.4 Guide Tube Test 1.5-7 1.5.1.5 Prototype Assembly Test 1.5-10 1.5.1.6 Departure from Nucleate Boiling (DNB) 1.5-12 1.5.1.7 Incore Flow Mixing 1.5-14 1.5.2 Inpile Fuel Densification 1.5-15 1.5.3 LOCA Heat Transfer Test (17 x 17) 1.5-16 1.5.3.1 Verification Facility Testing 1.5-16 2 1.5.3.2 D NB Test 1.5-19 1.5.3.3 Single Rod Burst Test (SRBT) 1.5-22 1.5.3.4 Power-Flow Mismatch 1.5-24 1.5.4 Westinghouse Test Engineering Laboratcry Facility 1.5-25 1.5.4.1 Introduction 1.5-25 1.5.4.2 Test Loops and Equipment 1.5-27 1.5.5 References 1.5-36 i LIST OF FIGURES Figure Title Page
- 1. 5-1 Schematic of ECCS Verification Test Facility 1.5-37 ,
1.5-2 DNB Test Facility Schematic 1.5-38 l 1-1
.. -. - . - - . . _ - . . - - - _ ~ _ = - - .- - . .- - . -
l 1.5 REQUIREMENTS FOR FURTHER TECHNICAL INFORMATION Reference [1] presents descriptions of the safety related Research and Development l , Programs which are being carried out for, or by, or in conjunction with, Westinghouse l Nuclear Energy Systems, and which are applicable to Westinghouse pressurized l water reactors. ! For each program still in progress the safety related program is first introduced, followed, where appropriate, by background information. There is, then, a description of the program which relates the program objectives to the problem and presents pertinent recent results. Finally, a back up position may be given for programs --generally experimental rather than analytical -- which have not yet reached a stage where it is reasonably certain that the results confirm the expectation. The back up position is one that might be used if the results are unfavorable; it is not necessarily the only course that might be taken. The term "research and development", as used in this report, is the same as that used by the Commission in S,ection 50.2 of its regulations, that is: l
~
(n) 'Research and development' means (1) theoretical analysis, exploration or experimentation; or (2) the extension of investigative findings and theories of a scientific nature into pmotical application for experimental and demonstmtion purposes including the experimental production and testing of modete, devices, equipment, materiate, and processes. The technical information generated by these research and development programs j !. will be used either to demonstrate the safety of the design and more sharply define margins of conservatism, or will lead to design improvements. ) The most up-to-date expected completion dates for the programs relevant l to each plant will be incorporated in the Applicant's Safety Analysis l 1.5-1
l Report. The schedules for developing this technical information are compatible with plant schedules such that definitive results will be available before each plant design is complete and in time to consider alternatives in the program or changes in design or in plant operating conditions should the program results not corroborate their objectives. Progress in these development programs will be reported semiannually. New safety related research and development programs, which include existing programs which become safety related, will also be described. Included in the overall Research and Development effort are the programs below which are applicable to the new 17 x 17 fuel assembly. The description of these programs, as presented in RESAR-3 Amendment 3, have been updated to indicate the status of each program as of October 31, 1973. 1.5.1 VERIFICATION TESTS (17 x 17) Design of the reactor described in this amendment uses a 17 x 17 square array of fuel rods and thimbles in a fuel assembly and is conceptually similar to but geometrically different from the 15 x 15 array used in previous designs. The 17 x 17 design is considered to be a relatively small extrapolation of the 15 x 15 design. Comprehensive testing has been planned, however, to verify that the extrapolation is sufficiently conservative. Preliminary evaluation of the data obtained to date has not revealed any anomalies. Design changes, if necessary, will be made to the reference 17 x 17 hardware in the unlikely event that any of the experimental results fall outside the conservative design values used in analysis. Westinghouse maintains that no plant need be designated a prototype and instrumented to verify the 17 x 17 fuel design. The change in flow induced vibration response of the internals from a 15 x 15 to a 17 x 17 fuel design will be minimal for the following reasons: O 1.5-2 l
1
- 1. The only structural changes in the internals resulting from the design change from the 15 x 15 to the 17 x 17 fuel assembly (gD are the guide tube and control rod drive line.
- 2. The guide tube is rigidly attached at the upper core support plate only. The upper core plate serves only to align the guide tubes. Because of this type of support arrangement the guide tube has a minimal contribution to the vibrational response of the core barrel and other internals components.
- 3. The effective flow area of the 17 x 17 guide tube is essentially the same as that of the 15 x 15 and therefore there are no significant differences in the flow distribution in the upper plenum.
- 4. The differences in mass and spring rate between the 15 x 15 and 17 x 17 fuel assemblies are very small (approximately 3%).
This insures that the effects of the fuel on the vibrational response of the reactor internals will remain essentially un-( ) changed. The preoperational hot functional flow testing pre-
- sented in Chapter 14 is considered the most conservative test condition since higher flow rates exist.
[ l S. More adequate and meaningful tests to verify the change from 15 x 15 to 17 x 17 would be to test the new guide tube and i fuel assembly designs individually in a special test facility such as the loop test facilities at the Westinghouse Forest , Hills site. This type of program is in fact being conducted and being reported in subsequent sections of this document. 1.5.1.1 Rod Cluster Control Spider Tests t Test Purpose and Parameters The 17 x 17 RCC spider (Section 4.2.3.2) is conceptually similar to, but geometrically different from the 15 x 15 spider. The 1.5-3
i 17 x 17 spider. supports 24 rodlets (the 15 x 15 design supports 20) , with no vane supporting more than two rodlets (same as 15 x 15 design). l( The purpose of the RCC spider tests is to verify the structural adequacy of the design. The spider vane to hub joint will be tested for structural adequacy by (a) vertical static load test to failure and (b) vertical fatigue test to approximately three million steps. The spider vane to hub joint tests and the spring tests are similar to tests performed on the 15 x 15 spider. The spring pack, within the spider hub will be tested to determine the spring load-deflection characteristic as a function of the loading cycles seen by the spring. It is planned to terminate the test after one thousand cycles. . Facility The 17 x 17 spider tests, as were previous spider tests, are being
- /7 performed at the Westinghouse Engineering Mechanics 1.aboratory.
(SeeSection 1.5.4.2.15). Status Spider tests are underway. A vertical static load test approximately seven times the design load did not result in spider vane to hub joint failure. A spider has been tested to 1.8 x 106steps without failure. l l 1.5.1.2 Grid Tests Test Purpose and Parameters I The 17 x 17 grid (Section 4.2.1.2) is conceptually similar but geometrically different from the 15 x 15 "R" grid. The purpose d
- O 1.5-4
l of the grid tests is to v:rify the structural ad:quacy of the L t grid design. The grid tests are similar to tests performed q on the various 15 x 15 grid designs. . .U l Load-deflection tests have been made on the grid spring and dimple. This information will verify that the fuel rod clad wear evaluation has been based on conservative values of these parameters. The grid buckling strength has been determined. The tests had short sections of fuel tubing inserted in the grid in place of fuel rods. Static and dynamic buckling were investigated along with impact damping. These tests are used to verify that grid buckling during a postulated seismic occurrence will not interfere with control rod insertion. Facility The grid tests were conducted in the Westinghouse Forest Hills Engineering Mechanics Laboratory (See Section 1.5.4.2.15). Status The grid tests have been completed. Test results are in agreement with pretest design values. The test results, along with fuel assembly structural test results, are being factored into the seismic analysis. 1.5.1.3 Fuel Assembly Structural Tests Test Purpose and Parameters The 17 x 17 fuel assembly (Section 4.2.1.2) was tested for mechanical strength including internal damping, lateral and axial stiffness, and lateral and axial impact responses for both static and dynamic loading. The lateral impact response will be used for verification .J l 1.5-5 1
l 2 of seismic analysis. The axial impact response is for verification of blowdown force analysis. The remaining tests are primarily I h,m to verify the assembly has sufficient mechanical strength to prevent damage during shipment, normal handling, and normal operation. These tests are similar to tests previously perfoyed on the 15 x 15 fuel assembly. Lateral tests were accomplished with both nozzles fixed in place and forces applied to various grids. The lateral stiffness is found by incrementally increasing and decreasing the static load. l The fuel assembly (lateral) natural frequencies and modes shapes j were obtained with the nozzles supported by core pins. An electro-
]
dynamic shaker was used to provide excitation. Fuel assembly l internal damping values were also obtained from these tests. l Lateral impact tests were perfonned by displacing the center of the assembly with the nozzles fixed in place. The assembly was i fm released and allowed to impact on lateral restraints at each of the five center grid locations. The axial stiffness was found by incrementally increasing the static load (compressive) and then incrementally decreasing the j static load. The axial impact response and damping were found by dropping the fuel assembly from various heights. The axial impact test was performed with the fuel assembly in the upright position. Facility These tests are being conducted at the Westinghouse Engineering Mechanics Laboratory (See Section 1.5.4.2.15). i (v j l.5-6 l l l
Status d(~% The fuel assembly structural tests have been completed. Preliminary j evaluation indicates lateral stiffness, frequency, and mode shape were conservative relative to predictions. All other results are in agreement with predictions. The fuel assembly structural test results , are being factored into the seismic and blowdown analyses. I l
- 1. 5.1.4 - Guide Tube Tests Test Purpose and Parameters A new rod cluster control guide tube is being designed to accommodate the 24 rodlet pattern adopted for 17 x 17 cores and which is sufficiently strong to assure that the support column function is not impaired and to provide increased margins of safety over present guide tubes.
- A high degree of interchangeability of parts has been designed into three new guide tube designs (96 and 150 inch for 4-loop cores; 125 inch for 3-loop cores). The main features of the new designs are full length enclosures and. cylindrical upper guide tubes. The 17 x 17 rodlet pattern reduced the central area available for drive line passage significantly, thus necessitating a generally tighter design of the rod guidance elements.
I To verify one structural adequacy of the new guide tubes, an extensive I series M tests will be conducted to determine guide tube deflection with simulated blowdown forces comparable to those expected during l ! a Loss-of-Coolant Accident and to determine the maximum acceptable deflection which assures insertion of a control rod by free fall. l Additional tests will be conducted to determine fatigue strength, displacement as a function of strain and the natural frequencies l of. the guide tubes for use in dynamic analyses. This test series is described below.
/"N-1.5-7 9
)
Upper Internals Scale Model Flow Test i q 1-Q In this series of tests, a 1/7 scale model of the upper internals will be employed to determine the radial and lateral induced vibration and flow forces on the guide tubes and support columns. This test will be conducted in the Westinghouse Test Engineering Laboratory H-Loop. (SeeSection1.5.4.2.5). l 17 x 17 Guide Tube Dynamic Characteristics This test will determine the strain versus displacement characteristic and the natural frequency in both air and water. This test will l be conducted in the Westinghouse Test Engineering Laboratory Autoclave Pit. (See Section 1.5.4.2.11). 17 x 17 Guide Tube Fatigue Tests The fatigue strength will be determined for each of the three new guide tube designs (i.e. three lengths for various 3 loop and 4 loop applications). Tests will be conducted in the Westinghouse Test Engineering Laboratory Autoclave Pit. Full Scale Flow Distribution Test l Using one full scale 17 x 17 guide tube (96 inch style) and support column, tests will be conducted to detennine the flow distribution and pressure balance between the guide tube and support column during normal operation and during blowdown. The normal operation tests will be conducted in the H-Loop and the blowdown tests in the E-Loop, respectively, of the Westinghouse Test Engineering Laboratory Facility. (See Section 1.5.4.2.3). I l Guide Tube Drop and Deflection Test In a simulated blowdown condition, tests will be conducted to determine r~ the guide tube deflection and control rod drop time. In addition, LU l.5-8 l
( the guide tube deflection which prevents control rod drop will l be detennined. These-tests are to be conducted at the Westinghouse Test Engineering Laboratory Facility. Facility
'Section 1.5.4 covers the Westinghouse Test Engineering Laboratory Facility description and capability.
Status Upper Internals 1/7 Scale Model-Flow Test Support column and guide tube calibration tests have been completed, the model assembled, and data col,lection has begun. Approximately 50% of the data required to complete the initial phase of testing. has been acquired. Preliminary data ~ analysis shows tilat none of the measured forces exceed the maximum calculated forces. Dynamic Tests Dynamic test to determine mode shapes and natural frequencies have been completed. The. first beam mode natural frequency is in good agreement with the calculated value. Fatigue Test A fatigue test is in progress. The guide tube has been vibrated for 107 cycles at the initial fatigue test vibratory amplitude and for 106cycles at twice the initial test amplitude with no damage detected. . Testing will be continued'at higher amplitudes until damage is detected. O 1.5-9 ! 1 I
, _. - , . _ b
1 Flow Distribution Tests j i I A, The flow distribution tests are in progress and are to be completed by the end of October. Preliminary analysis shows that the present hardware provides satisifactory flow distribution to the top of the core.
)
Static Deflection Tests (96 Inch Guide Tube) l Static load tests to measure strains and deflections for several support conditions were completed. 1.5.1.5 Prototype Assembly Tests The purpose of these tests is to demonstrate that the 17 x 17 fuel assembly and control rod hardware designs will perfonn as predicted. Two prototype assemblies will be sequentially tested in order to obtain the required experimental data. A single set of control rod hardware, including driveline, will be used in the tests. The fuel assemblies will be subjected to flow and system conditions covering l those most likely to occur in a plant during normal operation as v well as during a pump overspeed transient. These tests will be used to verify the integrated fuel assembly and RCC performance in several areas. Data to be obtained includes pressures and pressure drops throughout the system, hydraulic loadings on the fuel assembly and drive line, control rod drop time and stall velocity, fuel rod vibration and control rod, drive-line, guide tube, and guide l thimble wear during a lifetime of operation. Specifically, two full-size 17 x 17 fuel assemblies, one control rod, drive shaft and control rod drive mechanism will be installed and tested in the 24-in. I.D. x 40 foot high D-Loop at the Westinghouse l Test Engineering Laboratory Facility. J 1.5-10
l Fue'; Assembly Life Test (Phase I) p The first fuel assembly will be subjected to the maximum expected
-V control rod travel during fuel irradiation. The nominal test condi-tions will be a flow velocity of 119% of the design flow rate, tempera-ture of 585 F and pressure of 2000 psig. These conditions represent an extreme set of conditions.
Using a fully instrumented 17 x 17 prototype fuel assembly, guide tube and RCC drive assembly, tests will be conducted in the D-Loop to obtain information on the following:
- 1. Evaluate mechanical integrity and performance
- 2. Determine drop time
- 3. Fuel rod vibration 4 Control rod velocity
- 5. Hydraulic lift force
- 6. Guide thimble dashpot pressure Following this, the prototype' fuel assembly will undergo a complete V post test evaluation and the guide tubes and drive line will be inspected for any abnormal wear conditions.
Guide Tube and RCC Life Test (Phase II) The second fuel assembly will then be installed to continue the test at the same flow and temperature until 3,000,000 total steps of the drive mechanism are accumulated. These tests will be directed toward:
- 1. Life wear evaluation
- 2. Drop time 3 Stepping forces in drive line
- 4. Rod stall
- 5. Guide tube strains O
O 1.5-11 l l
Wh:n ab:ut one-third of the life test is completed, the test assembly will be inspected to determine guide tube and drive line wear charac-O gJ teristics. This inspection will be repeated at the end of the test.
\
Facility Section 1.5.4 covers the Westinghouse Test Engineering Laboratory Facility description and capability. Status Phase 1 of the D-loop testing is in progress. Preliminary data j indicates that fuel rod vibration frequencies are higher and vibration l amplitudes are lower than predicted. This apparent performance is I conservative relative to predictions. 1.5.1.6 Departure from Nucleate Boiling (DNB) Test Purpose and Parameters
~
The effect of the 17 x 17 fuel assembly geometry on the DNB heat flux will be determined experimentally and will be incorporated in a modified spacer factor for use with the W-3 correlation. The effect of cold-wall thimble cells in the 17 x 17 geometry will also be quantified. A similar program was conducted to quantify the DNB performance of.the R-type mixing vane grid as developed for the 15 x 15 fuel assembly design [2,3 . The results of that program were used to develop a modified spacer factor which quantifies the power capability associated with the use of the R mixing vane grid as well as the change in power capability due to the axial spacing of the grids. The modified spacer factor, along with the W-3 correlation with the cold-wall factor, was shown to be applicable to cold-wall thimble ! cells in the 15 x 15 geometry [3] , 1.5-12 I L
;Y
l The experimental program will consist of three test series employing rod bundles which are representative of the 17 x 17 fuel assembly geometry. Two of the tests will employ all heated rods; one test (V'] section being eight feet long and the other being fourteen feet long. The third test will have one simulated cold-wall thimble tube. All three tests will employ a uniform axial heat flux. The applicability of DNB data obtained using a uniform heat flux to a non-uniform heat flux has been well established by use of an axial flux shape factor. Tong [5]firstdevelopedtheformofthefactor. This same form l with some minor change in the empirical constants has been confirmed byWilson.bb3 This method of analysis has roven correct for non-uniform rod bundle data as shown by Rosal b , Motley [2] and Wilson [6] , Facility These tests will continue to be conducted in the high temperature and high pressure loop that was constructed by Westinghouse at the Columbia University Heat Transfer Laboratories. The loop charac-teristics of this facility are listed below: l /~T b Flow Rate 400 gpm maximum 40 gpm minimum Working Pressure 3500 psia maximum Test Section Inlet Temperature: 650*F maximum Test Section Outlet Temperature: 700"F maximum Test Section Heated Length: 16 feet maximum i l l Power Input to Test Section: 7.5 MWe maximum I v 1.5-13 l
l l l The 17 x 17 DNB tests will be performed parametrically for various combinations of inlet temperature and flow rate by increasing the bundle power incrementally until DNB occurs. Status ! i Design and fabrication of the test section hardware is complete. ! The 17 x 17 DNB testing began in mid-October. The test programs are , scheduled for completion during 1974 in order to s@ port the initial l loading of the first 17 x 17 fuel scheduled for early 1975. ' 1.5.1.7 Incore Flow Mixing Test Purpose and Parameters In the thermal-hydraulic design 'of a reactor core, the effect ) of mixing or turbulent energy transfer within the hot assembly is evaluated using the THINC code. The rate of turbulent energy g transfer is formulated in the THINC analysis in tems of a themal diffusion coefficient (TDC). A program to determine the proper value of TDC for the R grid vane, as used in the 15 x 15 fuel assenbly design, has been completed and showed that a design value of 0.038 (for 26 inch spacing) can be used for TDC. These results also showed that TDC was independent of Reynold's number, mass velocity, pressure, and quality over the ranges tested. A new TDC experimental program will employ a geometry typical of the 17 x 17 fuel assembly to determine the effects of the geometry on mixing and to detemine an appropriate value for TDC. A uniform axial heat flux will be used. There is no analytical reason to expect that the mixing coefficient would be affected by a non-uniform axial heat flux. The THINC computer code considers the mixing in each increment along the heated length and within that increment v 1.5-14
l l l the heat flux is considered uniform. The tests reported by Cadek[8] 1 s indicate that there was no difference, within experimental accuracy, j
.g)
( between a test section with a uniform flux (Pitt) and one half of i a cosine flux (Columbia). The heat flux will vary between the simulated fuel rods in the test section to create a thermal gradient in the l radial direction. Using different flow rates and inlet temperatures, the TDC for the 17 x 17 geometry will be determined. l Facility These tests will be conducted at the Columbia University Heat Transfer Laboratories in the equipment described in 1.5.1.6. Status The TDC tests are scheduled to follow the DNB tests and have not been initiated.
, 1.5.2 INPILE FUEL DENSIFICATION
]
Recent operating experience with uranium dioxide fuel has indicated that the fuel may densify under irradiation, to a greater density than that to which it was manufactured. This densification can lead to shorter active fuel length stacks, increased initial rod-i to-clad radial gaps, and pellet-to-pellet axial gaps. The shorter fuel stack length gives rise to a small increase in overall, average linear power density (kW/ft). Increased radial gap dimensions result in reduced gap conductance and lead to higher pellet temperatures. Axial gaps give rise to local power peaking due to decreased neutron absorption. ! Westinghouse fuel densification research is directed toward producing i fuel with a structure which minimizes inpile densification (hereafter
' called stable fuel). The objective of the program is to define material characteristics and manufacturing processes which lead to stable fuel. Stable fuel is defined as fuel whose densification is small.
(4), Any residual effects of densification will be evaluated on an appropriate 1.5-15
l l model davaloped in this program. A more detailed description of l the program and results will be presented in " Safety-Related Research and Development for Westinghouse Pressurized Water Reactors."U } 1.5.3 LOCAHEATTRANSFERTESTS(17x17) Extensive experimental programs have been completed or are in progress to detemine the thermal hydraulic characteristics of 15 x 15 fuel assemblies, and to obtain experimental reflooding heat transfer data under simulated loss-of-coolant accident conditions. Complementary experimental programs will be performed with a simulated 17 x 17 assembly to determine its behavior under similar LOCA conditions. The 17 x 17 test will be conducted in a new test loop which is presently being activated. Initial tests in this loop will be 'perfomed on 15 x 15 fuel assemblies to provide a basis for normalization to previous
)
results of the "FLECHT" series. ! Results from the 17 x 17 programs will be compared with data from ) the 15 x 15 assembly test programs and will be used to confim predictions made by correlations and codes based on the 15 x 15 test results. 1.5.3.1 Verification Facility Testing { Test Purpose and Parameters a The Verification Facility will provide experimental measurements of clad surface temperature and fluid flow conditions as a function of time during the period 2 to 8 and 8 to 20 seconds of blowdown, refill and reflood. It is believed that this transient blowdown L heat transfer information can be the subject of revised Westinghouse correlation. The Verification Facility will also provide experimental information on the refill and reflood phases of the accident. The Verification I O 1.5-16
- . _ - - - - . _ - - - - . . - . - . - - - - ~ _ .
Facility has the correct volume configuration to handle these phases. The major limitation of the verification facility during these Q phases is that system representation is limited in this facility to an orifice plate and downstream control valve, controlling back pressure. The large rod bundles of the Verification Facility should provide an accurate represer'ation of these important phases of the accident with experimental data providing the basis for new heat ; transfer correlations. The steady state runs have provided experimental data applicable to the small break behavior. This data will be used to check the current foaming or bubble rise model. This data will be useful in modifications of the current models. Experiments will be performed
.to determine the minimum water height required to cool a rod bundle as a function of bundle power.
Facility Description The Verification Facility has the capability of simulating blowdown in either of two core assemblies, the first contains an array of 529 rods of the type in the 15 x 15 assembly - 480 of which are heated rods and 99 of which are instrument rods; the second contains j an array of fuel rods of the type used in the 17 x 17 fuel assembly. i The facility can simulate unidirectional flow during the blowdown portion beginning as early as 2 seconds after depressurization if radial power zoning is utilized. The facility can simulate the ECCS ; behavior during blowdown as well as during refill and reflood. This simulation is limited to orifices in the upstream and downstream portions of the loop. The facility can simulate small break behaviors since its pressure capability is 2000 psi. Figure 1.5-1 shows a schematic of the ECCS Verification Test Facility. The Verification Test Facility was designed in accordance with the following considerations: O 1.5-17
- 1. The fluid volume below the active heater rods is 14.8 ft3. It [
l was obtained by scaling the volume below the actual fuel rods in a 4-loop PWR which is 1215 ft3 , l{
- 2. The vertical distance from the bottom of the active heater rods to the lowest elevation of the connecting pipes is 49 inches.
This is typical of the vertical distance in a PWR from the bottom of the active fuel rods to the lower surface of the core support > plate.
- 3. The flow resistance from the bottom of the active heater rods l to the side tank is t.,e same as the resistance in a PWR from ,
the bottom cf the active fuel rods to the lower plenum. l
- 4. The downcomer height from the lowest elevation of the connecting ,
pipes to the horizontal outlet pipe is 242 inches. This is the same as the vertical distance in a PWR from the lower surface of the core support plate to the lower surface of the outlet ! nozzles. 5 The downcomer cross sectional flow area is scaled from the combined
, cross sectional area of the PWR downcomer and barrel-baffle region.
Status The test facility began operation in December,1972. Six months of testing and debugging was devoted to small break testing where relationships were devel~oped for core void fraction as a function of pressure and downcomer head, and decay heat flux. In June, 1973 large break testing was initiated. At this time, UHI blowdown tests have been successfully completed to evaluate the effect of variations in UHI injection temperature, flow rate, flow maldistribution, core flow rate and inlet quality, power level, pressure decay rate,.and l initial clad temperature. Improvements to the facility are being considered. The future plans are to continue systematic investigation l 1.5-18
l of test conditiens and extend the testing to the phases of the LOCA following blowdown. A
'd 1.5.3.2 2
D NB Test Test Purpose and Parameters 2 The objective of the D NB (Delayed Departure from Nucleate Boiling) Test is to determine the time that DNB occurs under LOCA conditions. - 1 Currently, the AEC Interim Policy Statement on " Criteria for Emergency Core Cooling Systems for Light-Water Power Reactors" defines the basis, evaluation models, and conservative assumptions to be used in the evaluation of the perfomance of Emergency Core Cooling Systems (ECCS). Westinghouse believes that some of the conservatism of. the criteria is associated with.how transient DNB phenomena are treated in the evaluation models. Transient critical heat flux data presented at the 1972 specialists meeting of the Comittee on Reactor Safety Technology (CREST) indicated that the time to DNB can be delayed by several seconds. To demonstrate the conservatism l of the Interim Acceptance Criteria (IAC), Westinghouse has initiated ) an accelerated program to experimentally simulate the blowdown l phaseofaloss-of-coolar,iaccident(LOCA). 2 The D NB tests, which are part of this LOCA program, will be used to confim the predictions made with the new W transient DNB correlation. The motivation for conducting the D NB test is independent of the change over to the 17 x 17 fuel and are being conducted with a 15 x 15 geometry. However, the results of the steady state DNB program, described in Section 1.5.1.6, will be used to assure that the minimal geometric difference between the 17 x 17 and 15 x 15 arrays can i be correctly treated in transient correlations. The program is divided into two phases. Phase I will provide data directly applicable to the PWK to permit definition of the time delay associated with onset of DNB. Tests in this phase will cover l O
- 1.5-19 1
I
a range of cold and hot leg breaks, with particular emphasis on the large double-ended guillotine cold leg break. All tests in Phase I will be started upon establishment of typical steady state operating conditions. The fluid transient would then be initiated, and the rod power decayed in such a manner as to simulate the actual heat input of fuel rods. The transient will be required to follow a predetermined behavior as predicted by Westinghouse computer codes and the as-designed system hydraulics. The test would be terminated when the heater rod temperatures reach a predetermined limit (dependent on power level). The parameters to be studied under Phase I testing are shown below. Parameters Range INITIAL STEADY STATE CONDITIONS Pressure 2200 to 2250 psia 6 Test section mass velocity 2.0 to 2.6 x 10 lb/hr-ft 2 Inlet coolant temperature 540 to 560 F TRANSIENT CONDITIONS Simulated breaks Various break sizes will be ! simulated to cover range of typical large and small breaks ; Phase II will provide separate effects data to permit heat transfer correlation development.
. The Phase II tests will also start from steady state conditions, with sufficient power to maintain nucleate boiling throughout the bundle.
Controlled ramps of decreasing test section pressure or O 1.5-20 e
flow will initiate DNB. By applying a series of controlled conditions, investigation of the DNB will be studied over a range of qualities and flows, and at pressures relevant to a PWR blowdown. To obtain
, qualities higher than in the Phase I tests, a steam generator will be added to the unbroken loop prior to the Phase II tests.
The parameters to be studied in Phase II testing are shown below. Parameter Range INITIAL STEADY STATE CONDITIONS Pressure 200 to 2250 psia 6 Test section mass velocity 0.1 to 2.5 x 10 lb/hr-ft 2 Test section quality 0 to superheat (depending on system pressure) TRANSIENT RAMP CONDITIONS Pressure decrease 100 to 50,000 psi /sec (subcooleddepressurization) Flow decrease . 10 to 1500 %/sec Flow reversal +100% to -50% in 0.05 sec Quality increase 10 to 50 %/sec The experiments in the Delayed DNB Facility will result in cladding temperature and fluid properties measured as a function of time through-out the blowdown range from 0 to 20 seconds. Facility The experimental program will be conducted in the J-Loop at the Westing-house Test Engineering Laboratory Facility with a full length 5 x 5 rod bundle simulating a section of a 15 x 15 assembly to determine DNB occurrence under loss of coolant accident conditions. 2 The schematic for the D NB Facility is shown in Figure 1.5-2. (See also Section 1.5.4.2.6). O 1.5-21 r
The heater rod bundles to be used in this program will be assembled (
) using internally-heated rods. The proprietary heater rods are designed
!. j. for high reliability, long life, and high power density. The maximum j power is 18.8 kw/ft, and the total power is 136 kw for extended periods I over the 12-foot heated length of the rod. Heat is generated internally by means of a varying cross-section, rugged, tubular resistor which 2 approximates a U cos U power distribution, skewed to the bottom. Each rod is adequately instrumented with a total of 20 thermocouples , (8 inside resistor,12 sheath thermocouples). Status 1 The construction of the D2 NB facility is complete. Shakedown testing ! has been completed and all instruments have been calibrated. Phase I testing is presently starting and will continue to the end of the year. 1.5.3.3 Single Rod Burst Test (SRBT) I i ig Test Purpose and Parameters The single rod burst test results are used to quantify the maximum assembly flow blockage which is to be assumed in LOCA analyses. Previously, single rod and multi-rod burst test (MRBT) have been completed on 15 x 15 fuel assembly rods under conditions which exist during the loss-of-coolant accident. The conclusion of these tests were that fuel rods burst in a staggered manner so that maximum average assembly wise flow area blockage is 55 percent during blowdown and 65 percent during reflood based on the characteristics of the pressurized PWR fuel rod and the conservative peak clad temperature predicted during the LOCA transient.
,a
() 1.5-22
The single rod burst test program for the 17 x 17 fuel assembly rods consisted of testing specimens, at the two internal pressures and the three heating rates listed below in a steam atmosphere. Heating Rate Internal Pressure (725'F to 1940 F) psi 5 F/sec 1200, 1800 25*F/sec 1200, 1800 100*F/sec 1200, 1800 All specimens were then heated 5'F/sec from 1940*F to about 2300*F held for a short time and then cooled 5 F/see to 1200 F. Meta 11ography was done on specimens to determine the degree of wall thinning and the extent of oxygen embrittlement. In addition, tests were run on 15 x 15 fuel assembly rods to insure reproducibility of the 1972 single rod burst test results. O Facility The SRB tests were conducted in the Westinghouse Engineering Mechanics Laboratory in an electrically heated furnace. Status The single rod burst tests have been completed. These tests showed that the LOCA behavior of 17 x 17 clad in comparison to that of 15 x 15 clad exhibited no'significant differences in failure ductility. Because of the result and the geometric sealing, the flow blockage (90) as dttermined by 15 x 15 MRBT simulation can be used for 17 x'17 fuel geometry. 1.5-23
. _ . _ __. .. ~ . _. _ _ _m ._ __ _ _ _ . _ _ _ . . _ . . _ _ _ _ _ . _ . . _ _
l 1.5.3.4 Pownr-Flow Mismatch i i Design emphasis has been placed on reliable and effective control ) and protection systems for the reactor core and engineered safety features to ensure adequate margins within which incidents can be terminated before the onset of fuel failure. For this reason, investigations into the mechanisms of fuel failure and its propa-gation, and phenomena of molten fuel-coolant interaction have been limited. I Obtaining complete answers to the question related to fuel rod failure will require extensive testing because of the multiplicity of para-meters involved. Such an extensive inpile test program has been proposed by the Aerojet Nuclear Corporation under AEC sponsorship for the Power Burst Facility (PBF). l The proposed PBF program is expected to simulate conditions appro-priate to major accidents pos'tulated for reactor systems, i.e., loss-of-flow', loss-of-coolant, and reactivity-initiated accidents. Phase I of the proposed PBF program is expected to be completed approxi-mately two years after facility checkout or by the end of 1975. Phase I is expected to cancentrate upon establishing the thresholds for and the consequences of fuel failure and utilizes fuel rod clusters of various sizes to determine the extent of failure propagation. This program would broaden the experimental basis for evaluating reactor safety, but should not be considered essential for the design and safe operation. Westinghouse will closely follow any such experimental program or analytical studies that may become available. Until such inform-ation is available, clearly demonstrating tnat local fuel melting is an acceptable condition, the emphasis in design will continue to be placed on providing adequate margins to minimize the prob-ability of fuel melting.. The margins incorporated in the design within which incidents can be terminated before the onset of fuel failure, provide a sound basis for safe operation. L 1.5-24
. _ - . ..-. . __- - _ . . _ - =-. --- _ .- 1.5.4 WESTINGHOUSE TEST ENGINEERING LABORATORY FACILITY 1.5.4.1 Introductdon l The Test Engineering Laboratory at Forest Hills, Pennsylvania, has l 4 long been the major Westinghouse center for nuclear research and development. The Test Engineering Laboratory is totally involved with the design and implementation of facilities and programs to prove the reliability of Westinghouse PWR concepts and components. The Test Engineering Laboratory has full in-house capabilities to design , and construct pressurized water reactor loops for both hydraulic and heat transfer testing programs. The most vital current project is Emergency Core Cooling Systems (ECCS), which involves scale-model tests, run on three separate facilities. The "G" Loop consists of a test vessel, which presently contains a 480 heater rod bundle, the largest such test facility in the world. It also has a steam supply to provide the proper environment during system j blowdown, and the capability to test high-pressure and low-pressure ECCS. "G" Loop operates at pressures up to 2000 psi and temperatures up to 650*F. It is designed to start operation at 8 seconds after a LOCA (Loss-of-Coolant Accident), and is presently capable of investigating the current ECCS, Upper-Head Injection and other spray systems.
"J" Loop consists of a test vessel, which contains a 25 heater rod l array, a broken loop simulation and an unbroken loop simulation. The loop is designed to operate at 2500 psi and 650 F, and is capable of simulating the first 20 seconds of a LOCA with primary emphasis on Delayed Departure from Nucleate Boiling (DDNB).
I FLECHT-SET consists of a test vessel, which contains a 100 heater rod and thimble array, and is used to investigate the reflood phase of the current ECCS with plant system effects measured with scaled piping anJ two scale-model steam generators. The facility is designed to operate at up to 100 psia. 1.5-25 l
l Fiva g:n=ral purpose hydraulic loops are also involved in the devel- 1 opment of improved water reactor components, as well as the reliability - ! testing of current and prototype PWR components. Details of these l
, loops, as well as the heat transfer loops mentioned above, are included in the enclosed folders.
Historically the Test Engineering Laboratory has been in a state of transition, depending upon the current need for its services. Today's great need is for ECCS data and the verification of many new PWR system ; l components. Past needs and accomplishments have included the development of supercritical heat transfer once through loops; rod cluster control L dirve mechanisms; fuel assemblies; underwater handling tools; and fuel assembly grid design, among many other earlier projects. Testing has included air filter tests; water chemistry tests; in-pile testing I for fuel rods; single fuel rod burst tests; hydraulic studies on fuel assemblies; and corrosion testing of zircaloy and other PWR components and materials, with and without heat transfer. l The Test Engineering Laboratory is a very flexible installation, one I which will continue to expand and develop as future needs for its services arise. Its staff, too, varies according to requirements. There are currently more than 100 persons involved in Laboratory projects, including 12 electrical and mechanical engineers, more than 75 highly skilled technicians, and some 30 specialists from other divisions of Westinghouse. The Test Engineering Laboratory has the option of obtaining personnel from the entire Cormration, depending upon the need for specific skills, knowledge and experience. Ongoing research performed at the Test Engineering Laboratory continues to demonstrate the reliability of Westinghouse PWR plant components and greatly facilitates the development of improved reactor system components. As the test center for Westinghouse Nuclear Energy Systems, the Test Engineering Laboratory is totally comitted to the advancement i of the nuclear energy industry. r !O
- 1.5-26 i
i
l.5.4.2 Test Loops and Equipment l This section contains a brief description of the major test loops [ and test equipment at the Westinghouse Test Engineering Laboratory Facility. 1.5.4.2.1 "A" & "B" Loops. Low-Flow /High-Pressure Hydraulic Facilities These loops are small, high pressure, stainless steel facilities, used for testing small components and individual parts of larger components under normal working conditions. A canned motor pump circulates water in both "A" Loop and "B" Loop at 150 gpm. Operating temperatures are obtained from the conversion of the pumping power into heat, as ! ! well as from external heaters. Typical tests run in these loops are: ' a) full-scale gate and check valves; b) material corrosion-erosion, with variable water chemistry; and, c) corrosion product release ^and transport properties of crud. ! Characteristics of "A" & "B" Loops l Maximum Flow Rate 150 gpm at 300 ft. Maximum Pump Head' 335 ft. at 60 gpm Maximum Allowable Temperature 650 F Normal Working Pressure 2000 psi Normal Working Temperature 600 F
-1.5.4.2.2 "D" Loop. Medium-Flow /High-Pressure Hydraulic Facility
- The "D" Loop is a flexible test facility used for demonstrating the interplay of reactor subsystems and evaluating component design concepts. It contains a canned motor pump, which produces a l
l I 1.5-27
- O l
t g - - , . , c - - ,
l 290 ft. head at 3000 gpm. All piping (10-inch Schedule 160) in contact with the primary water is stainless steel. Loop pressure is established and maintained by an air driven charging pump operating in conjunction with a gas loaded back pressure valve. Most of the power required to establish and maintain loop temperature is derived from the circulating ! pump operation, and 75 Kw of heat is available from electric strip heaters. l The "D" Loop services a 24" ID x 40' long test vessel, which acconno-dates full-scale models of large PWR core components for operational i studies. studies. l l Characteristics of "D" Loop Maximum Flow Rate 4400 gpm Maximum Allowable Pressure 2400 psi' Maximum Allowable Temperature 650*F Normal Working Pressure 2000 psi Normal Working Temperature 600'F O) U Pump Head at 3000 gpm 290 ft. Maximum Pump Head 340 ft. (at 1500 gpm) Main Loop Flow Measurement 10" Venturi Auxiliary Flow Measurement 6" Venturis (2" Branch Lines) 1.5.4.2.3 "E" Loop. Low-Flow / Low-Pressure Hydraulic Facility The "E" Loop is a low-pressure, six-inch, stainless steel loop, with two circulating pumps. These pumps may be connected in l parallel, giving 2000 gpm at 130 ft, head, or in series, giving j 1000 gpm at 260 ft. head. Flow and vibration studies are conducted l with this loop, and, because of its low pressure, plastic models for visual observation or photography may be used. In addition, a 4-inch Rockwell water meter in a branch line permits the calibration of flow meters up to 800 gpm. 1.5-28
1 Characteristics of "E" Loop l , liaximum Flow Rate 2000 gpm at 130 f t. (] l 1000 gpm at 360 ft. I Maximum Working Pressure Pump Head 1.5.4.2.4 "G" Loop. Emergency Core Cooling System Facility The "G" Loop is a nigh-pressure, emergency core cooling (ECCS) test facility designed and fabricated to AS!4E Section 1 for 2000 psi and 650 F. It consists of a main test section and vessel, exhaust system, piping, separators and muffler, flash chamber steam supply system, and high-pressure / low-pressure cooling systems. This loop is basically designed to obtain test data for analysis of LOCA, for breaks up to and including double-ended pipe breaks for Pressurized Water Reactors. Tests are initiated at simulated (
) conditions existing 8 seconds after the start of a LOCA (Loss-of-Coolant Accident). A typical run consists of constant power and pressure, followed by pressure blowdown, power decay and ECCS.
"G" Loop is capable of performing the following methods of ECCS:
Current, Upper Head Injection (UHI), UHI w/ Current and other core spray systems. It may also be used for constant temperature / pressure small leg break tests (core uncovering tests). These consist of boiling off water at a constant bundle power input until the rods can no longer be cooled. The "G" Loop test bundle consists of 480 electrically heated rods,16 grid support thimbles, and 33 spray thimbles bounded by an octagonal stainless steel baffle and arranged as per a 4-loop 15 x 15 rod bundle configuration. The loop is controlled (fully automated during transients) through a PDP-II-DEC-16K 13 V 1.5-29
computer with a 600 point Computer Products A-D Converter opera-ting at a sweep rate of 40,000 pts /second for data acquisition.
"G" Loop System Components & Characteristics Rated Typical Operating Press. Temp. Press. Temp.
Component Material (psi) ( F) (psi) ( F) Test Vessel Carbon Steel 2000 650 1000 545 Downcomer Side Tank Carbon Steel 2000 650 1000 54 5 In-Line Mixer Carbon Steel 2000 650 1000 545 Mixer Accumulator Stainless Steel 2500 650 1800 100 Flash Chamber Carbon Steel 3000 700 2800 660 Separators Nos.1 & 2 Carbon Steel 2000 650 1000 545 Spray Accumulators kos. 1 & 2 Carbon Steel 2000 650 1800 150 Spray Accumulator No. 3 Stainless Steel 2500 650 1800 150 Reflood Tank Stainless Steel Atmos. 212 Atmos. 150 Primary Piping Carbon Steel 2000 650 1000 545 C 1.5.4.2.5 "H" Loop. High-Flow Hydraulic Facility The "H" Loop constitutes a versatile hydraulic facility, capablo of supplying 14,000 gpm of water at a developed head of 600 ft. and at temperatures as high as 200 F. This 4-loop system can simultaneously ! handle either full-scale prototype test assemblies, or one large-scale reactor model. The major purpose of "H" Loop is to permit ! the use of 1/7 scale reactor models and full-scale fuel assemblies for conducting mixing studies, flow distribution studies, and similar low-temperature / low-pressure hydraulic tests. I O v 1.5-30
L I Characteristics of "H" Loop i l l o 1 > Lh L Maximum Flow Rate Pressure Drop Across Vessel Model 14,000 gpu 120 psi Minimum Vessel Outlet Pressure 10 psig- l Flow Accuracy 1/2 % ! 1 Water Temperature Range 70-200 F
]
Maximum Loop-to-Loop Temp. Variation 2*F i Max. Loop-to-Loop Flow Rate Variation ' 3% i L 1.5.4.2.6 '"J" Loop. Delayed Departure from Nucleate Boiling L , Heat Transfer Facility ! l The "J" Loop is a completely instrumented pressurized water. test facility for verifying DDdB phenomena during a LOCA (Loss-of-Coolant Accident), and for conducting steady state heat transfer studies. This test loop is a full-size, single-loop simulation of a typical 4-loop reactor system; it will accept a full-length 5 x 5 bundle of internally heated " fuel rods." '"J" Loop is designed to operate at 2500 psia at 650"F, and at variable flow rates of up to 450 gpm. 1 During LOCA tests, fluid input to the " reactor vessel" is closely controlled by two servo-controlled mixers, which inject a two-phase water / steam mixture into the test vessel, to simulate flow from the unbroken loops. ! Characteristics of "J" Loop Test Fluid Demineralized Water Design Pressure 2500 psia Design Temperature 650 F Maximum Flow Rate (hot) 450 gpm Power Input to Test Vessel 3,500,000 watts (max.) Primary Test Heat Exchanger Rating 11,400,000 BTV/HR O 1.5-31
l 1.5.4.2.7 "K" Loop. Boron Thennal Regeneration Test "X" Loop, the Boron Thermal Regeneration System (BTRS) test facility, is used to study the performance and to verify the component sizing of both the currently available THERM I and the improved THERM II BTR systems. The function of this system is to process baron-containing l effluents from the Reactor Coolant System (RCS) to yield a high- ! boron concentration fraction, which can be sued to borate the RCS. f A relatively boron-free fraction is also processed, which can be l used to dilute the RCS, such as that required in load-follow operations. Characteristics of "K" Loop Total Tank Capacity 30,000 gal. Chiller Capacity 48 ice-tons 3 Max. Ion Exchange Resin Test Volume 75 ft Max. Test Process Rate Capability 10 gpm/ft2 bed area Max. Flow Test Capability 200 gpm
- Min. Boron Storage Mo/e Fluid Temp. 50*F Max. Boron Release Mode Fluid Temp.
160 F 1.5.4.2.8 FLECHT-SET, Emergency Core Cooling System Facility The FLECHT-SET is a low-pressure facility, designed to provide experimental data on the influence of system effects on ECCS during the reflood phase of a Loss-of-Coolant Accident (LOCA). The facility consists of a once-through system, including an elec-trically heated test section (" fuel rods" and housing), accumulator, steam generator simulators, pressurizer, catch vessels, instrumenta-tion, and the necessary piping to simulate the reactor primary coolant loop. Data acquisition is accomplished through a PDP-II-DEC-16K ! Computer with a 256 point Computer Products A-D Converter, operating at a sweep rate of 1200 pts /second. I 10 e 1.5-32 w -
_ _ __, _ _ _._-._ m . _ _ m. _ _
. .m.__.__.__.__..____ _ _ . _ _ _ _ _ _ _. -
Characteristics of FLECHT-SET I 100 Rod Bundle Maximum Power 1000 Kw Maximum Bundle Flooding Rate 86 gpm {' Water Temperature Range 100-200 F System Pressure 0-60 psia l 1.5.4.2.9 Single-Rod Loop. Heater Rod Development Facility l The single-rod test loop is used to evaluate prototype heater rods-and for in-depth study of existing rods in pressurized water systems. l The test section of the loop is easily replaced to facilitate the ) !' installation of various length and diameter heater rods. The l-Single-Rod Loop is electrically controlled and operated by one person. Steady. state and blowdown at various conditions can be
- simualted in the loop. The main test section can be replaced with a quartz tube, and DWB phenomenon can be observed on a single rod with a remotely operated camera.
Characteristics of Single-Rod Test Loop Maximum Operating Pressure 2250 psia Maximum Operating Temperature 650 F Maximum Flow Rate 10 gpm System capacity 5 gal. Maximum Power' Available 200 Kw Piping Size 1" and 3" 1.5.4.2.10 Hydraulic Model Testing Miscellaneous hydraulic tests on mock-ups of reactor system parts and components are routinely performed at the Test Engineering Labora tory. Typical of this type of testing are the two discussed below, which were recently completed: lO L 1.5-33 \
, ~ . _ _ , ._ , _ . . , , . . . , _ _ , . - , _ _ .
Emergency Core Cooling Flow Distribution: As shown above, a 10 x 10 rod bundle was installed in a plastic housing with a water supply at the top. A grid-collection unit at the bottom of the bundle collected the water as it flowed through the model and diverted it to the measur-i ing tubes at the base. Knowledge of the flow distribution in the bundle was obtained in this manner. Sample System Mixing Test: This test used one thermocouple to measure the temperature of water from four locations in a reactor. The purpose of the procedure was to determine whether the indication from the single thermocouple was representative of the average temperature of the four water supplies. A mock-up of the mixing chamber was constructed I l so that hot or cold water - at closely controlled pressure - could be supplied to any of the four inlets. By running combinations of hot and cold inlets and making simultaneous recordings of the various tem-peratures, highly useful information was obtained. l l I 1.5.4.2.11 Autoclave Testing The Test Engineering Laboratory is equipped with autoclaves ranging in size from 1/2 gallon to 100 gallons. These devices are in constant use to determine the effect of various water chemistries on core I components, as well as to perform corrosion tests. The units have ! also been used as boilers to provide steam for miscellaneous develop-ment tests, including acoustic leak detection. I 1.5.4.2.12 Mechanical Component and Vibration Tests Full-scale mechanical and vibration tests are performed at the Test l Engineering Laboratory on plant and reactor components to prove the reliability of equipment design. Vibration testing of reactor components is also performed in this Laboratory, using electronically excited shaker heads. Three sizes are available (2 lbs., 50 lbs., and 150 lbs.) for regular scale model testing for frequencies from 5 Hz to 50 Hz.
'v l.5-34
{ . _
1 l 1.5.4.2.13 Electronic Component Assembly ( Highly skilled technicians are available at the Test Engineering Laboratory for constructing complex control and instrumentation systems. Work is initiated with engineering ideas and sketches,
, and includes mounting of process controllers, recorders, meters, relay logic, protection circuits, switches and indicators.
l l Point-to-point wiring or PCB's are used, as required. Final I
"as-built" drawings are prepared, inspection and thorough elec-trical checkout is performed before installation in a facility.
1.5.4.2.14 Surveillance System Development Surveillance Systems provide on-line monitoring of pressure vessels for flaws. Electronic components are being developed at the Test Engineering Laboratory for an acoustic emission monitoring system for in-service inspection of operating plant vessels and piping. This system is designed to detect and locate initiation and propagation of cracks at various locations, such as welds and stress risers. Vessel flaw growth and rupture data have been obtained through joint programs at the National Reactor Testing Station in Idaho, and at the Oak Ridge National Laboratories. Pipe rupture data have been obtained from AEC sponsored tests, and hydrostatic test data, operational noise and attenuation characteristics have been measured at various Westinghouse operating plants. 1.5.4.2.15 Engineering Mechanics Laboratory Bench tests are performed in fixtures designed for the particular test using standard test equipment and techniques. 1.5-35
1.
5.5 REFERENCES
- 1. " Safety Related Research and Development for Westinghouse Pressuri-zed Water Reactors, Program Summaries, Spring-Fall 1973,"
WCAP-8204.
- 2. F. E. Motley and F. F. Cadek, "DNB Results for New Mixing Vane Grid (R)": WCAP-7695-L, July 1972, (Westinghouse NES Proprietary);
and WCAP-7958, October 1972.
- 3. F. E. Motley and F. F. Cadek, "DNB Test Results for R Grids with Thimble Cold Wall Cells" : WCAP-7695-L, Addendum 1, October 1972, (Westinghouse NES Proprietary); and WCAP-7958, Addendum 1, October 1972.
I l 4 F. F. Cadek, and F. E. Motley and D. P. Dominicis, "Effect ' of Axial Spacing on Interchannel Thermal Mixing with R Mixing Vane Grid," WCAP-7941-L, June 1972, (Westinghouse NES Proprietary); and WCAP-7959, October 1972. O 5. L. S. Tong, " Prediction of Departure From Nucleate Boiling for j an Axially Non-Unifonn Heat Flux Distribution", J. Nucl. Energy,21,pp.241-248(1967). i
- 6. R. H. Wilson, L. J. Stanek, J. S. Gellerstedt, R. A. Lee, I
" Critical Heat Flux in a Non-uniformly Heated Rod Bundle,"
in Two-Phase Flow and Heat Transfer in Rod Bundles, pp. 56-62, ASME New York, November 1969.
- 7. E. R. Rosal, et. al., " Rod Bundle Axial Non-Uniform Heat Flux Tests and Data," WCAP-7411, December 1969 (Westinghouse NES Propr_ietary), and WCAP-7813, December 1971
- 8. F. F. Cadek, "Interchannel Thermal Mixing with Mixing Vane ;
Grids," WCAP-7667-L, May 1971 (Westinghouse NES Proprietary), and WCAP-7755, September 1971. 1.5-36
l ll ' CsI O R E HB SM AA i LH M FC r f -
}
[N - i v ja i R E x o i M [ r- M E (N g 6 N R O s M v I L T r N y RA L r, . I i t E l U i c 1x M WU a F C C c t s A t F e D T L O n F
- i o
IN t A a M L i c E f i TS r SS e
, EE 'v TV
, " \ S C
O
" C m E f
o
,, I.s n i
c
' c t a
_" s u
. o m S' /.
,I e
?
l c
! M n w *2 h c
F
- o o
R O S
- F i1 1
5
/
T A R A 1 K P 5 9! A E 1
- ) S f l i >
F'l F 4 A Y} e r 7 _ u F. . g NM ~ l V .. g F i l
,, 3
=
F- -
&D " )
O - LO LLK I FFN EEA 6 S R 2f. Ml F
' ,l Ti p
~RRT E
- F t. r F
L M O .
- h F F T A
% U M : R:
A* P
~ E
- S O
".y tl
. [
t c ! I . ; ' ! i. . ' O ~[ ~:4 {' . }
/
I l N k e e m o n.% _ ) s. . s .t e n i
. a0
. 3
_ . 6 --
- e.
- o. .
s .
. e _.
,n u ,
+
o W. g . t.
- n. .
p+ ' 1 c f e De 5Ts 3 _ f a e t r e e e .. at r te . a 6C i C eA 2 t s . i c
' m t u
c c a a t m E s t S S e
" n t v h F .i 1 c 0
- 1 1
S 1 t P y t N) i
~
~
)
s f li c
'( t O
s a st F es e t f s rfe 5 e 9t b r s e Pt e .t a
. T t g . t t e i B
- a. s ee
& _ s N
.p g ,
t e t e as I e D m 1 -
- e.
,a
* .. . 2 e -
P -
" 5 m
F 4, _- ,y 1 v e e e r f r u t, s
,- d t.. s g e , ,, a l e i l
l ' . S F ei * . t s
- e 5 b 9
* . . fp t
- c e o .
8 0
- et 4
0 t Y f . l ri = _ nu .e P o . 4 +- -- . e *g e-t t a t se u( n
,i ,',l/ . c' o
e
,.h .
. s;[l_
#m
.r
. m,wm
are equipped with guide thimbles joined to the grids and the t:p and bottom nozzles. Depending upon the position of the assembly in the core, the guide thimbles are used as core locations for rod cluster control assemblies, neutron source assemblies, and burnable poison rods. Otherwise, the guide thimbles are fitted with plugging devices to limit bypass flow. The bottom nozzle is a box-like structure which serves as a bottom structural element of the fuel assembly and directs the coolant flow distribution to the assembly. The top nozzle assembly functions as the upper ' +.ructural element of the fuel assembly in addition to providing a partial protective housing for the rod cluster control assembly or other components. The rod cluster control assemblies each consist of a group of individual absorber rods fastened at the top end to a comon hub or spider assembly. These assemblies are of two types, those with rods containing full length absorber material to control the reactivity of the core under operating conditions, and those with rods containing a 36 inch part length absorber section to control axial power distribu-tion. The control rod drive mechanisms for the full length rod cluster control assemblies are of the magnetic latch type. The latches are controlled by three magnetic coils. They are so designed that upon a loss of power to the coils, the rod cluster control assembly is released and falls by gravity to shutdown the reactor. The control rod drive mechanisms for the part length control rods are of a roller nut type mechanism which move at slow speed and stop motion on complete loss of power. Loss of power to these drive mechanisms will not result in any reactivity change due to rod motion. O 4.1-2
l l The components of the reactor internals are divided into three parts consisting of the lower core support structure (including j the entire core barrel and neutron shield pad assembly), the upper l core support structure and the in-core instrumentation support l structure. The reactor internals support the core, maintain fuel alignment, limit fuel assembly movement, maintain alignment between l t
)
fuel assemblies and control rod drive mechanisms, direct coolant j l j flow past the fuel elements and to the pressure vessel head, provide l gamma and neutron shielding, and provide guides for the in-core instrumentation. l l The nuclear design analyses and evaluation establish physical locations l for control rods and burnable poisons and physical parameters such as fuel enrichments and boron concentration in the coolant such that the reactor core has inherent characteristics which together with corrective actions of the reactor control, protective and i emergency cooling systems provide adequate reactivity control even if the highest reactivity worth rod cluster control assembly is stuck in the fully withdrawn position. O ( l The design also provides for inherent stability against diametral l and azimuthal power oscillations and for control of induced axial l power oscillations through the use of the part length control rods. l The themal-hydraulic design analyses and evaluation establish coolant flow parameters which assure that adequate heat transfer is assured between the fuel cladding and the reactor coolant. The themal design takes into account local variations in dimensions, I power generation, flow distribution and mixing. The mixing vanes incorporated in the fuel assembly spacer grid design induces additional flow mixing between the various flow channels within a fuel assembly as well as between adjacent assemblies. Instrumentation is provided in and out of the core to monitor the nuclear, thermal-hydraulic, and mechanical performance of the l reactor and to provide inputs to automatic control functions. l f V i 4.1-3 l-l
The reactor core design together with corrective actions of the reactor control, protection and emergency cooling systems can ( meet the reactor performance and safety criteria specified in Section 4.2. This Topical Report constitutes a Final Design Report for the # Westinghouse 17 x 17 core design and will be followed by amendments r4 to individual Final Safety Analysis Reports of those plants which will employ the 17il7 fuel. T6 illustrate the effects of the change in fuel design on the four loop plant, Table 4.1-1A presents a comparison of the principal nuclear, thermal-hydraulic and mechanical design parameters between the four loop plant with jd x 17 fuel assemblies ipcluding the effects of fuel densification, gtherTrojan~ Nuclear Plant (Docket 50-344) with 15 x 15 fuel assemblies (without fuel densification). A more general comparison 1 showing the difference between plants with and without loop stop valves without the effects of fuel densification, fs presented in Table 4.1-1 of RESAR-3. 1 Similarly, Table 4.1-1B presents a comparison of the three-loop l (] D plant with 17 x 17 fuel assemblies including the effects of fuel densification and the Beaver Valley Unit 1 (Docket No 50-334) with 15 x 15 fuel assemblies (without fuel densification). The effects of fuel densification for both the four loop and three 1 loop plants were evaluated with the methods described in Reference 1. The contents of Chapter 4, for the greater part, are applicable to both the four-loop and three-loop plants. Where required, specific differences are noted in the text. In Sections 4.1 and 4.2 suffixes A and B are employed in the table and figure numbers in a few cases where the four-loop and three-loop data are different. In Sections 4.3 and 4.4, however, the primary group of tables and figures include the four-loop data and that which is comon to three and four loops. The three-loop tables and figures which show different data are appended to the back of the respective Sections 4.3 and 4.4. These are not differentiated p by suffixes, but are identified as three loop parameters in the U title. DECEMBER, 1973 4.1-4
The analysis techniques employed in the core design are tabulated in Table 4.1-2. The loading conditions considered in general for the core internals and components are tabulated in Table 4.1-
. 3. Specific or limiting loads considered for design purposes of the various components are listed as follows: fuel assemblies in Section 4.2.1.1.2; reactor internals in Section 4.2.2.3 and Table 5.2-2; neutron absorber rods, burnable poison rods, neutron source rods and thimble plug assemblies in Section 4.2.3.1.3; full-and part-length control rod drive mechanisms in Section 4.2.3.1.4.
The dynamic analyses, input forcing functions, and response loadings are presented in Section 3.9. 4.
1.1 REFERENCES
- 1. J. M. Hellman (Ed.), " Fuel Densification Experimental Results and Model for Reactor Application," WCAP-8218, October 1973, I (Westinghouse NES Proprietary ); and WCAP-8219, October 1973.
O O 4.1-5 DECEMBER,1973
l TABLE 4.1-1A REACTOR DESIGN COMPARISON TABLE (FOURLOOPPLANT) l l REFERENCE PLANT TROJAN 17x17 FUEL ASSEMBLY 15x15 FUEL ASSEMBLY I WITH DENSIFICATION WITHOUT THERMAL AND HYDRAULIC DESIGN PARAMETERS EFFECTS DENSIFICATION EFFECTS
- 1. Reactor Core Heat Output, W t 3411 3411 6
- 2. Reactor Core Heat Output, Btu /hr 11,641.7 x 10 11,641.7 x 10 6
- 3. Heat Generated in Fuel, % 97.4 97.4
- 4. System Pressure, Nominal, psia 2250 2250 l
- 5. System Pressure, Min. Steady State, psia 2220 2220
- 6. Minimum DNBR for Design Transients >l.30 >l.30 Coolant Flow
- 7. Total Thermal Flow Rate,1b/hr 132.7 x 10 6 132.7 x 10 6
8 Effective Flow Rate for Heat Transfer,1b/hr 0 6 126.7 x 10 126.7 x 10
- 9. Effective Flow Area for Heat 2
Transfer, ft 51 .1 51 .2
- 10. Average Velocity Along Fuel Rods, ft/sec 15.7 15.5 !
2 0 6
- 11. Average Mass Velocity,1b/hr-ft 2.48 x 10 2.47 x 10 l Coolant Temperature, F 12.. Nominal Inlet 552.5 552.5
- 13. Average Rise in Vessel 64.2 64.2
- 14. Average Rise in Core 66.9 66.9
- 15. Average in Core 585.9 585.9
- 16. Average.in Vessel 584.7 584.7 Heat Transfer
- 17. Active Heat Transfer, Surface 2
Area, ft 59,700 52,200 4.1-6 DECEMBER, 1973
1 TABLE 4.1-1A (Continued) REACTOR DESIGN COMPARIS0N TABLE (FOUR LOOP PLANT)
' REFERENCE PLANT- TROJAN 17x17 FUEL ASSEMBLY 15x15 FUEL ASSEMBLY WITH DENSIFICATION WITHOUT I
- THERMAL AND HYDRAULIC DESIGN PARAMETERS EFFECTS DENSIFICATION EFFECTS
- 18. Average Heat Flux, Btu /hr-ft 2 189,800 217,200
- 19. Maximum Heat Flux for Normal Operation, Btu /hr-f t 2 474,500[a] 521,300
- 20. Average Thermal Output, kw/ft 5.45 7.03
-21. Maximum Thermal Output for Normal Operation, kw/ft 13.6[a] 16.9 l 22. Maximum Thermal Output at Maximum Overpower Trip Point (118%
power),kw/ft 15.2[c] I 21.1
- 23. Heat Flux Hot Channel Factor, F g 2.50[b] 2.40 Fuel Central Temperature,- F
- 24. Peak at 100% Power 3500 3950
- 25. Peak at Maximum Thennal Output for Maximum Overpower Trip Point 3800 4500 l
[a] This limit.is associated with the value of Fq = 2.50 ! [b] Includes the effect of fuel densification l [c] See Section 4.3.2.2.6 1 l i 4.l-7 DECEMBER, 1973 i w -- , . . , -, - - , , -~ - - ,
- . ... _. - . .___ .~ ._. . - .
l l TABLE 4.1-1A (Continued)
/'N U PEACTOR DESIGN COMPARIS0N TABLE l
(FourLoopPlant) 1 I REFERENCE PLANT TROJAN l7x17FUEk. ASSEMBLY 15x15 FUEL ASSEMBLY WITHOUT I WITH DENSIFICATION CORE MECHANICAL DESIGN PARAMETERS EFFECTS DENSIFICATION EFFECTS i Fuel Assemblies RCC Canless RCC Canless
- 26. Design 193 193
- 27. Number of Fuel Assemblies 264 204
- 28. UO Rods per Assembly 2
Rod Pitch, in. 0.496 ' O.563
- 29. '
y
- 30. Overall Dimensions, in. 8.426 x 8.426 yd' ,8.426 x 8.426
- 31. Fuel Weight (as U02 ), p unds 222,739 JM 218,367 50,913 4 4,424
- 32. Zircalloy Weight, lbs.
Number of Grids per Assembly 7-Type R 7 - Type L 1 e- 33. Loading Technique 3 region non-uniform 3 region non-uniform 34. Fuel Rods 50,952 39,372
- 35. Number _
Outside Diameter, in. 0.374% 0.422 36. Diametral Gap, in. , Regions 1,2, and 3 0.0065% 0.0075 37. Clad Thickness, in. 0.0225f4 0.024
- 38. ~
Zircaloy-4% Zircaloy-4
- 39. Clad Material Fuel Pellets Material 002 Sintered UO Sintered
- 40. 2 95g 94
- 41. Density (% of Theoretical)
Diameter, in. , Regions 1, 2, and 3 0.3229 0.3659 42. 0.530 / 0.600
- 43. Length, in.
0 4.1-8 DENR,1973
TABLE 4.1-1 A (Continued) (v/ REACTOR DESIGN COMPARISON TABLE (FourLoopPlant) REFERENCE PLANT TROJAN 17x17 FUEL ASSEMBLY 15x15 FUEL ASSEMBLY WITH DENSIFICATION - WITHOUT l CORE MECHANICAL DESIGN PARAMETERS EFFECTS DENSIFICATION EFFECTS i Rod Cluster Control Assemblies
- 44. Neutron Absorber, Full and Part Length Ag-In-Cd Ag-In-Cd
- 45. Cladding Material Type 304 Type 304 SS-Cold SS-Cold Worked Worked
- 46. Clad Thickness, in. 0.0185 .019
- 47. Number of Clusters, Full /Part Length 53/8 53/8 l
- 48. Number of Absorber Rods per
,q Cluster 24 20
(/ i Core Structure
- 49. Core Barrel I .D./0.D. , in. 148.0/152.5 148.0/152.5
- 50. Thermal Shield Neutron Pad Neutron Pad Design Design f l
l i l , /5 i DECEMBER, 1973 4.1-9
.. .. -- -... .. . . . _ - - ~ . . . . _
f TABLE 4.1-1A(Continued) REACTOR DESIGN COMPARIS0N TABLE
, (FourLoopPlant) ;
REFERENCE PLANT TROJAN j 17x17 FUEL ASSEMBLY 15x15 FUEL ASSEMBLY WITH DENSIFICATION - WITHOUT I ' NUCLEAR DESIGN PARAMETERS EFFECTS DENSIFICATION EFFECTS-Structure Characteristics
- 51. Core Diameter, in. (Equivalent) 132.7 132.7 1 l
- 52. Core Average Active Fuel Height, in. 143.7 144 l t
t Reflector:Thickneds and Composition
- 53. Top - Water plus Steel, in, s10 ) s10
- 54. Bottom - Water plus Steel, in. s10 s10
- 55. Side - Water plus Steel, in. $15 ) sl5
- 56. H 0/U, Cold Molecular Ratio 2
Lattice . 3.43) 3.55 Feed Enrichment, w/o :
- 57. Region 1; Z?ic g 2.25
- 58. Region 2i 2;60 % 2.80
- 59. Region 4 3.10) 3.30 i
i-l \ c 4.1-10 DECEMBER,1973
[ TABLE 4.1-1B REACTOR DESIGN COMPARISON TABLE (ThreeLoopPlant) 17x17 Reference BVPS Unit 1 l Plant With 15x15 Without Densification Densification Thermal and Hydraulic Design Parameters Effects Effects l l 1. Reactor Core Heat Output, MWt 26E2 2652 6 6
- 2. Reactor Core Heat Output, Btu /hr 9051x10 9051x10
- 3. Heat Generated in Fuel, % 97.4 97.4 l 4 System Pressure, Nominal, psia 2250 2250
- 5. System Pressure, Min. Steady l State, psia 2220 2220 6 Minimum DNBR for Design Transients 1.30 1.30 Coolant Flow 6 6
- 7. Total Thermal Flow Rate, lb/hr 100.9x10 100.9x10
- 8. Effective Flow Rate for Heat 6 6 Transfer, lb/hr 96.3x10 96.3x10
- 9. Effective Figw Area for Heat Transfer, ft' 41.5 41.8
- 10. Average Velocity Along Fuel Rods, ft/sec 14.4 14.4 6 6
- 11. Average Mass Velocity, lb/hr-ft 2 2.32x10 2.31x10 Coolant Temperature, F
- 12. Nominal Inlet 542.5 542.5
- 13. Average Rise in Vessel 67.4 67.4
- 14. Average Rise in Core 70.3 70.3
- 15. Average in Core 579.3 579.3
- 16. Average in Vessel 576.2 576.2 Heat Transfer
- 17. Ac e H at Transfer, Surfac 48,600 42,460 2
- 18. Average Heat Flux, Btu /hr-ft 181,400 207,600 4.1-11 DECEMBER,1973 l
TABLE 4.1-1B (Continued) REACTOR DESIGN COMPARISON TABLE
\
(Three Loop Plant) 17x17 Reference BVPS Unit 1 Plant With 15x15 Without Densification Densi fication I Themal and Hydraulic Design Parameters Effects Effects
- 19. Maximum Heat Flux for Normal Operation, Btu /hr-ft2 453,500[a] 580,000
- 20. Average Thermal Output, kw/ft 5.2 6.7 ,
- 21. Maximum Thermal Output for Normal I Operation, kw/ft 13.0{3) 18.8
- 22. Maximum Themal Output at Maximum Overpower Trip Point (118% power), kw/ftl4.6[c] 21.1
- 23. Heat Flux Hot Channel Factor, F 1 Fuel Central Temperature, F Q 2.50[b] 2.72(1)
- 24. Peak at 100% Power 3400 4250
- 25. Peak at Maximum Thermal Output for Maximum Overpower Trip Point, "F 3700 4500 Core Mechanical Design Parameters !
Fuel Assemblies
- 26. Design RCC Canless RCC Canless
- 27. Number of Fuel Assemblies 157 157
- 28. U0 2Rods per Assembly 264 204
- 29. Rod Pitch, in. 0.496 0.563 )
- 30. Overall Dimensions, in. 8.426x8.426 8.426x8.426 i
- 31. Fuel Weight (as U02), Pounds 181,205 176,200
- 32. Zircalloy Weight, lbs. 38,230 36,300
- 33. Number of Grids per Assembly 7 - Type R 7 1
- 34. Loading Technique 3 Region 3 Region Non-uniform Non-uniform Fuel Rods
- 35. Number 41,448 32,028
- 36. Outside Diameter, in. 0.374 0.422 .
[a] This limit is associabd with the value of F =2.50 q [b] Includes the effect of fuel densification. [c]Seesection4.3.2.2.6 1 4.1-12 DECEMBER, 1973
TABLE 4.1-1B (Continued) REACTOR DESIGN COMPARISON TABLE I. (Three Loop Plant) _ 17x17 Reference BVPS Unit 1 Plant With 15x15 Without Densification Densification I l Core Mechanical Design Parameters Effects Effects
- 37. Diametral Gap, in. Regions 1,2 0.0065 0.0075 (Region 3) (0.0065) (0.0085)
- 38. Clad Thickness, in. 0.0225 0.0243
- 39. Clad Material Zircaloy-4 Zircaloy-4 Fuel Pellets ,
- 40. Material 00 U0 2.! 2 !
Sinterk Sintered '
- 41. Density (% of Theoretical) 95 94
- ' 42. Diameter, in. , Regions 1,2 0.324 0.3659 y (Region 3) (0.325) (0.3649)
- 43. Length, in. - 0.!iB0 0.600 Rod Cluster Control Assemblies
- 44. Neutron Absorber Ag-In-Cd Ag-In-Cd
- 45. Cladding Material Type 304SS Type 304SS Cold Worked Cold Worked
- 46. Clad Thickness, in. 0.0185 0.019
- 47. Number of Clusters, Full /Part Length 48/5 48/5
- 48. Nunber of Absorber Rods per Cluster 24 20 Core Structure
- 49. Core Barrel , I .D./0.D. , in. 133.85/137.875 133.85/137.875
- 50. Thermal Shield, I.D./0.D., in. Neutron Pads Neutron Pads O 4.1-13 DECEMBER, 1973
I l l l TABLE 4.1-1B (Continued) REACTOR DESIGN COMPARISON TABLE (Three Loop Plant) 17x17 Reference BVPS Unit 1 Plant With 15x15 Without Densification Densification I I Nuclear Design Parameters Effects Effects Structure Characteristics j 51. Core Diameter, in. (Equivalent) 119.7 119.7
- 52. Core Average Active Fuel Height, in. 143.7 144 l1 Reflector Thickness and Composition l l 53. Top - Water plus Steel, in, s10 s10 i
- 54. Bottom - Water plus Steel, in, s10 s10
- 55. Side - Water plus Steel, in. sl5 s15 g 56. H 20/U, Cold Molecular Ratio, (Lattice) 3.43 3.55 V
l Feed Enrichment, w/o
- 57. Region 1 2.10 2.00
- 58. Region 2 2.60 2.70
- 59. Region 3 3.10 3.35 (I) The design value of F = 2.72 was employed for nuclear and thermal-q hydraulic design. However, to meet the Interim Acceptance Criteria for emergency core cooling, the allowed gF is 2.32.
O 4.1-14 DECEMBER, 1973
* "4 #%. 'i.
'\ ,i &! N,)
TABLE 4.1-2 ANALYTIC TECHNIQUES IN CORE DESIGN Analysis Technique Computer Code Section Referenced Mechanical Design of Core Internals l1 Loads, Deflections, and Static and Dynamic Blowdown code, FORCE, 3.7.2.1 Stress Analysis Modeling Finite element structural 3.9.1 analysis code, and others 3.9.3 Fuel Rod Design Fuel Performance Characteristics Semi-empirical thermal Westinghouse fuel rod 4.2.1.3.1 (temperature, internal pressure, model of fuel rod with design model 4.3.3.1 clad stress, etc.) consideration of fuel 4.4.2.2 density changes, heat 4.4.3.4.2 transfer, fission gas f* release, etc. .'m l Nuclear Design
- 1) Cross Sections and Group Microscopic data Modified ENDF/B library 4.3.3.2 1 Macroscopic constants LEOPARD / CINDER type 4.3.3.2 Constants for homogenized core regions Group constants for control HAMMER-AIM 4.3.3.2 rods with self-shielding c,
R 2-D, 2-Group Diffusion TURTLE 4.3.3.3
- 2) X-Y Power Distributions, Fuel Depletion, Critical Theory
"!! Boron Concentrations, x-y ? Xenon Disbributions, 53 Reactivity Coefficients to
'~5 A r"~N
[d U U TABLE 4.1-2 (Continued) ANALYTIC TECHNIQUES IN CORE DESIGN Analysis Technique Computer Code Section Referenced Nuclear Design (Continued)
- 3) Axial Power Distributions, 1-D, 2-Group Diffusion PANDA 4.3.3.3 Control Rod Worths, and Theory Axial Xenon Distribution
- 4) Fuel Rod Power Integral' Transport Theory LASER 4.3.3.1 Effective Resonance Monte Carlo Weighting REPAD Temperature Function Thennal-Hydraulic Design
?
7 1) Steady-state Subchannel analysis of THINC-IV 4.4.3.4.1 g local fluid conditions in rod bundles, including inertial and crossflow resistance tems, solution progresses from' core-wide to hot assembly to hot channel
- 2) Transient DNB Analysis Subchannel analysis of local THINC-I (THINC-III) 4.4.3.4.1 fluid conditions in rod
- bundles during transients by including accumulation terms in conservation equations; solution progresses from core-wide to hat assembly to hot channel
- 3) ' Hydrodynamic Stability Prediction of flow instability HYDNA 4.4.3.5 in parallel closed channels
L TABLE 4.1-3 DESIGN LOADING CONDITIONS FOR REACTOR CORE COMP 0NENTS _ 1. Fuel Assembly Weight
- 2. Fuel Assembly Spring Forces ,
- 3. Internals Weight
- 4. Control Rod Scram (equivalent static load)
- 5. Differential Pressure
- 6. Spring Preloads
- 7. Coolant Flow Forces (static)
- 8. Temperature Gradients
- 9. Differences in thennal expansion
- a. Due to temperature differences
- b. Due to expansion of different materials
- 10. Interference between components
- 11. Vibration (mechanically or hydraulically induced) '
- 12. One or more loops out of service V 13. All operational transients listed in Table 5.2-2
- 14. Pump overspeed
- 15. Seismic loads (operation basis earthquake and design basis earthquake)
- 16. Blowdown forces (due to cold and hot leg break) 1 l
I l 1 i i l O 4.1-17 l l i
i I ! l l 4.2 MECHANICAL DESIGN l The plant conditions for design are divided into four categories L in accordance with their anticipated frequency of occurrence and risk to the public: Condition I - Normal Operation; Condition II - ! Incidents of Moderate Frequency; Condition III - Infrequent Incidents; 1 Condition IV - Limiting Faults. The reactor is designed so that its components meet the following performance and safety criteria:
- 1. The mechanical design of the reactor core components and their j physical arrangement, together with corrective actions of the reactor control, protection, and emergency cooling systems (when applicable) assure that: -
- a. Fuel damage
- is not expected during Condition I and Condition j II events. It is not possible, however, to preclude a very small number of rod failures. These are within the O capability of the plant cleanup system and are consistent with the plant design bases,
- b. The reactor can be brought to a safe state following a
- Condition III event with only a small fraction of fuel rods damaged
- although sufficient fuel damage might occur to preclude resumption of operation without considerable outage time.
l
- c. The reactor can be brought to a safe state and the core can be kept subcritical with acceptable heat transfer geometry following transients arising from Condition IV events.
l [
- Fuel damage as used here is defined as penetration of the fission product barrier (i.e.' the fuel rod clad).
4.2-1
2 The fuel assemblies are designed to accommodate expected con- i ditions for design for handling during assembly inspection and refueling operations and shipping loads.
- 3. The fuel assemblies are designed to accept control rod insertions in order to provide the required reactivity control for power operations and reactivity shutdown conditions.
- 4. All fuel assemblies have provisions for the insertion of in-core instrumentation ne.essary for plant operation.
l l 5. The reactor internals in conjunction with the fuel assemblies direct reactor coolant through the core to achieve acceptable flow distribution and to restrict bypass flow so that the heat transfer performance requirements can be met for all modes of operation. In addition, the internals provide core support and distribute coolant flow to the pressure vessel head so that the temperature differences between the vessel flange and head do not result in leakage from toe flange during the Condition I and Il modes of operation. Required in-service i inspection can be carried out as the internals are removable ; and provide access to the inside of the pressure vessel. 4.2.1 FUEL 1 4.2.1.1 Design Bases l l l The fuel rod and fuel assembly design bases are established to satisfy the general performance and safety criteria presented in Section 4.2 and specific criteria noted below. 4.2.1.1.1 Fuel Rods 1 ! The integrity of the fuel rods is ensured by designing to prevent excessive fuel temperatures, excessive internal rod gas pressures due to fiss'on gas releases, and excessive cladding stresses and b v l 4.2-2
i strains. This is achieved by designing the fuel rods so that the , following conservative design bases are satisfied during Condition ! I and Condition II events over the fuel lifetime: , 9
- 1. Fuel Pellet Temperatures - The center temperature of the hottest pellet is to be below the melting temperature of the U02 (melting ;
point of 5080 F b3 unirradiated and reducing by 58 F per 10,000 141D/MTU) . While a limited amount of center melting can be I tolerated, the design conservatively precludes center meltina. ) A calculated centerline fuel temperature of 4700 F has been l selected as an overpower limit to assure no fuel melting. This provides sufficient margin for uncertainties, as described I in Section 4.4.1.2 and 4.4.2.10.1.
- 2. Internal Gas Pressure - The internal gas pressure is less than the nominal coolant design pressure. This conservative limit precludes primary tensile stresses in the clad.
- 3. Clad Stress - The clad stresses are less than the Zircaloy yield stress, with due consideration of temperature and irradiation effects. While the clad has some capability for accommodating ]
plastic strain, the yield stress has been accepted as a conser-vative design basis. , 1
- 4. Clad Tensile Strain - The clad tensile strain is less than one percent. This limit is consistent with proven practice.
I S. Strain Fatigue - The cumulative strain fatigue cycles are j less than the design strain fatigue life. This basis is consistent 1 l with proven practice. L The fuel rods are designed for a peak pellet burnup of approximately 50,000 Megawatt Days per Metric Ton of Uranium (MWD /MTU) in the fuel cycle equilibrium condition, while the fuel is designed for
- a region average discharge burnup of 33,000 MWD /t4TU (Nuclear Design i
!O l 4.2-3 DECEMBER,1973
l Basis, Section 4.3.1) for the equilibrium cycle. The detailed fuel rod design establishes such parameters as pellet size and density, O clad-pellet diametral gap, gas plenum size, and helium pre-pressure. ' The design also considers effects such as fuel density changes, l fission gas release, clad creep, and other physical properties which vary with burnup. An extensive irradiation testing and fuel surveillance operational
~
experience program is being conducted to verify the adequacy of E the fuel performance and design bases. This program is discussed ' in Sections 4.2.1.3.3, 4.2.1.3.4 and Sections 8 and 17 of Reference [6]. Fuel surveillance and testing results, as they becone available, are used to improve fuel rod design and manufacturino processes
- and assure that the design bases and safety criteria are satisfied. I i 4.2.1.1.2 Fuel Assembly Structure Structural integrity of the fuel assemblies is assured by setting
! limits on stresses and deformations due to various loads and by determining that the assemblies do not interfere with the functioning of other components. Three types of loads are considered.
- 1. Non-operational loads such as those due to shipping and handling, l
- 2. Normal and abnormal loads which are defined for Conditions '
I and II,
- 3. Abnormal loads which are defined for Conditions III and IV.
These criteria are applied to the design and evaluation of the top and bottom nozzles, the guide thimbles, the grids and the thimble l joints. l l -The design bases for evaluating the structural integrity of the fuel assemblies are: t LO L 4.2-4
- - . - ~ . _ - - . - - - - - . - - - . - - - - . _ - - . . - _ . . _ .
I
- 1. Non-Operational - 69 loading with dimensional stability,
- 2. Normal Operation (Condition I) and Incidents of Moderate Frequency I (Condition II),
For normal operating conditions, Section III of the ASME Boiler and Pressure Vessel Code has been used as a basis for evaluating acceptability of calculated stresses. Both static and alternating stress intensities are considered. For materials not covered by the code, allowable stresses are established in the same manner as used in the code for ma'terials of similar properties.
- 3. Abnormal loads during Conditions III or IV - worst cases repre-sented by combined seismic and blowdown loads,
- a. Deflections of components cannot interfere with the reactor shutdown or emergency cooling of the fuel rods,
- b. The fuel assembly components are designed in accordance O with ASME Section III, Appendix F, " Rules for Evaluation 1 of Faulted Conditions." The criteria for the grid component are based on experimental tests used in conjunction with the limits established in Table F-1322.2-1 of Appendix F.
4.2.1.2 Design Description The standard fuel assembly and fuel rod design data are given in Table 4.3-1. Two hundred and sixty-four fuel rods, twenty-four guide thimble tubes and one instrumentation thimble tube are arranged within a supporting structure to form a fuel assembly. The instrumentation thimble is located in the center position and provides a channel for insertion of an incore neutron detector, if the fuel assembly is located in an instrt.mented core position. The guide thimbles provide channels for insertion of either a rod cluster control assembly, DECEMBER, 1973 4.2-5
I a neutron source assembly, a burnable poison assembly or a plugging i device, depending on the position of the particular fuel assembly () in the core. Figure 4.2-1 shows a cross-section of the fuel assenbly array, and Figure 4.2-2 shows a fuel assembly full length view. The fuel rods are loaded into the fuel assembly structure so that there is clearance between the fuel rod ends and the top and bottom , nozzles. All standard fuel assemblies in the core are identical l l in mechanical construction. ! l Each fuel assembly is installed vertically in the reactor vessel and stands upright on the lower core plate, which is fitted with I: alignment pins to locate and orient the assembly. After all fuel assemblies are set in place, the upper support structure is installed. Alignment pins, built into the upper core plate, engage and locate the upper ends of the fuel assemblies. The upper core plate then I bears downward against the fuel assembly top nozzle via the holddown springs to hold the fuel assemblies in place. 4.2.1.2.1 Fuel Rods O The fuel rods consist of uranium dioxide ceramic pellets contained in slightly cold worked Zircaloy-4 tubing which is plugged and seal l welded at the ends to encapsulate the fuel. A schematic of the fuel rod is shown in iigure 4.2-3. The fuel pellets are richt circular cylinders consisting of slightly enriched uranium-dioxide powder which has been compacted by cold pressing and then sintered to the required density. The ends of each pellet are dished slightly to allow greater axial exp6nsion at the center of the pellets. ; To avoid overstressing of the cladding or seal welds, void volume and clearances are provided within the rods to accommodate fission gases released from the fuel, differential thermal expansion between the cladding and the fuel, and fuel density changes during burnup. Shifting of the fuel within the cladding during handling or shipping prior to core loading is prevented by a stainless steel helical spring which bears on top of the fuel. At assembly the pellets Q/ ' 4.2-6 l
. .- _ . _ . _ _ . _- - - - _ = . - . - _ - . . . . _- .
are stacked in the cladding to the required fuel height, the spring is then inserted into the top end of the fuel tube and i! the end plugs pressed into the ends of the tube and welded. All fuel rods are internally pressurized with helium during the welding process in order to minimize compressive clad stresses and creep due to coolant operating pressures. The helium cas ore-pressurization is not defined until the final j fuel rod design, which is completed after a FSAR submittal. The helium pre,-pressurization may be different for each fuel region and for various plants, since a fuel rod pressurization is dependent on the planned fuel burnup as well as other fuel design parameters and fuel characteristics (particularly densification potential). 1 l The fuel rods are presently being pre-pressurized and designed ; so that (1) the-internekget pressure wf1Tnot" exceed"the nominal systmo coolant' pressure evemduring-anticipated transientse(Condition II), and-(2) clad flattening,wilL notroccur duriner t W fueb core li fe.- l ( 4.2.1.2.2 Fuel Assembly Structure The fuel assembly structure consists of a bottom nozzle, top nozzle, guide thimbles and grids, as shown in Figure 4.2-2. Bottom Nozzle The bottom nozzle is a box-like structure which serves as a bottom structural element of the fuel assembly and directs the coolant flow distribution to the assembly. The square nozzle is fabricated from type 304 stainless steel and consists of a perforated plate and four angle legs with bearing plates as shown in Figure 4.2-2. The legs form a plenum for the inlet coolant flow to the fuel assembly. The plate itself acts to prevent a downward ejection of the fuel rods from their fuel assembly. The bottom nozzle is fastened to the fuel assembly guide tubes by weld-locked screws which penetrate through the nozzle and mate with an inside fitting in each guide tube. . DECE!SER, 1973 4.2-7
l Coolant flow through the fuel assembly is directed from the plenum in the bottom nozzle upward through the penetrations in the plate to the channels between the fuel rods. The penetrations in the Y plate are positioned between the rows of the fuel rods. I Axial loads (holddown) imposed on the fuel assembly and the weight of the fuel assembly are transmitted through the bottom nozzle to the lower core plate. Indexing and positioning of the fuel assembly is controlled by alignment holes in two diagonally ooposite bearing ! plates which mate with locating pins in the lower core plate. Any lateral loads on the fuel assembly are transmitted to the lower core plate through the locating oins. 1 Top Nozzle ' The top nozzle assembly functions as the upper structural element of the fuel assembly in addition to providing a nartial protective housing for the rod cluster control assembly or other comoonents. It consists of an adapter pla.te, enclosure, top plate, hold down ' /~'T springs, clamps, and pads as shown in Figure 4.2-2. The sprinos and bolts are made of Inconel 718 and Inconel 600 respectively, whereas other components are made of type 304 stainless steel. i The square adapter plate is provided with round and obround penetra-tions to permit the flow of coolant upward through the top nozzle. Other round holes are provided to accept the guide thimbles which are then mechanically attached to the adapter plate. The ligaments in the plate cover the tops of the fuel rods and prevent their upward ejection from the fuel assembly. The enclosure is a sheet metal shroud which sets the distance between the adapter plate and the top plate. The top plate has a large square hole in the center to permit access for the control rods and the control rod spiders. l Holddown springs are mounted on the top plate and are fastened in l place by bolts and clamps located at two diagonally opposite corners. On the other two corners integral pads are positioned which contain alignment holes for locating the upper end of the fuel assembly. O V i 4.2-8 i l l
, -. - - . ~ . - - - . - _ - . - _ . . - . . . . . - _ . . . . . - , . . ~ . . 1 l I Guide and Instrument Thimbles The guide thimbles are structural members which also provide channels for the neutron abscrber rods, burnable poison rods or neutron source I assemblies. Each one is fabricated from Zircaloy-4 tubing having two different diameters. The larger diameter at the top provides a relatively large annular area to permit rapid insertion of the control rods during a reactor trip as well as to accommodate the flow of coolant during normal operation. Four holes are provided l on the thimble tube above the dashpot to reduce the rod drop time. The lower portion of the guide thimbles has a reduced diameter to produce a dashpot action near the end of the control rod travel l l during normal operation and to accommodate the outflow of water I from the dashpot during a reactor trip. The dashpot is closed at the bottom by means of an end plug which is provided with a small flow port to avoid fluid stagnation in the dashpot volume durina normal operation. The top end of the guide thimble is fastene'd into the top nozzle adapter plate. The lower end of the guide thimble is fitted with an end plug which is then fastened into the bottom - I nozzle by a weld locked screw. The central instrumentation thimble of each fuel assembly is not attached to either the top or bottom nozzles, but the thimble is constrained by its seating in counterbores of each nozzle. The thimbles internal diameter does not vary, and in-core neutron detectors pass through the bottom nozzle's large counterbore into the center I thimbl e. Grid Assemblies i I The fuel rods, as shown in Figure 4.2-2, are supported laterally i at intervals along their length by grid assenblies which maintain the lateral spacing between the rods throughout the design life of the assembly. Each fuel rod is afforded lateral support at six contact points within each grid by the combination of support dimples and springs. The grid assembly consists of individual slotted straps l 1 O 4.2-9
. ~ _ . _ _ _ _ _ . - .. _ =.. _ . - . _ . _ . . _ . _ _ . _ _ .._ . _ _
interlocked and brazed in an " egg-crate" arrangement to join the straps pemanently at their points of intersection. The straps entain spring fingers, support dimples and mixing vanes. The grid material is Inconel 718, chosen because of its corrosion resistance and high strength properties. The magnitude of the crid restraining force on the fuel rod is set high enouoh to minimize
- possible fretting, without overstressing the cladding at the points l of contact between the grids and fuel rods. The grid _ asse
- tblies
- also allow axial themal expansion of the fuel rods without imposing restraint sufficient to develop buckling or distortion of the fuel rods.
Two types of grid assemblies are used in each fuel assemSly. One type, with mixing vanes projecting from the edges of the straps I~ into the coolant stream, is used in the high heat flux reoion of l the fuel assemblies to promote mixing of the coolant. The other type, located at the ends of the assembly, does not contain mixing i vanes on the internal straps. The outside straps on all grids contain mixing vanes which, in addition to their mixing function, aid in guiding the grids and fuel assemblies past projecting surfaces during handling or during loading and unloading of the core. i 4.2.1.3 Design Evaluation 4.2.1.3.1 Fuel Rods The fuel rods are designed to assure the design bases are satisfied for Condition I and II events. This assures that the fuel performance ) and safety criteria (Section 4.2) are satisfied. Materials - Fuel Cladding The desired fuel rod cladding is a material which has a superior
- combination of neutron economy (low absorption cross section), high I
l 4.2-10 . l l rw- - - ~ ,
- . . - - . _ - - . .~ -- - - - -. - . . . .-
strength (to resist deformation due to differential pressures and I mechanical interaction between fuel and clad), high corrosion resistance (to coolant, fuel and fission products), and high reliability. Zircaloy-4 has this desired combination of cladding properties. As shown in Reference 9, there is considerable PWR operating experience on the capability of Zircaloy as a cladding material. Clad hydriding has not been a significant cause of clad perforation since current controls on fuel contained moisture levels were instituted.E93 l Materials - Fuel Pellets 1 Sintered, high density uranium dioxide fuel reacts only slightly with the cladding, at core operating temperatures and pressures . In the event of cladding defects, the high resistan,ce of uranium dioxide to attack by water protects against fuel deter-ioration although limited fuel erosion can occur. As has been shown by operating experience and extensive experimental work, the themal design parameters conservatively account for changes in the thermal performance of the fuel elements due to pellet fracture which may occur during power operation. The consequences of defects in the cladding are greatly reduced by the ability of uranium dioxide to retain fission products including those which are gaseous or highly volatile.
- Observations from several operating Westinghouse PWR's (References 6,9) has shown that fuel pellets can densify under irradiation to a density ,
higher than the manufactured values. Fuel densification and subsequent i incomplete settling of the fuel pellets results in local and distributed gaps in the fuel rods. An extensive analytical and experimental effort is underway in Westinghouse to characterize the fuel densification phenomenon and identify improvements in pellet manufacturing to eliminate or minimize this anomaly (see ; l Section 1.5). ' l The effects c'/ fuel densification have been taken into account in the nuclear and thermal-hydraulic design of the reactor described herein in Sections 4.3 and 4.4, respectively. 4.2-11 l l 1 _ _ _ . ._ ,
l l Materials - Strength Considerations l
\
O b One of the most important limiting factors in fuel element duty is the mechanical interaction of fuel and cladding. This fuel-l cladding interaction produces cyclic stresses and strains in the { cladding, and these in turn consume cladding fatigue life. The j reduction of fuel-cladding interaction is therefore a principal ' goal of design. In order to achieve this goal and to enhance the cyclic operational capability of the fuel rod, the technology for using pre-pressurized fuel rods in Westinghouse PWR's has been developed. Initially the gap between the fuel and cladding is sufficient to prevent hard contact between the two. However, during power operation a gradual compressive creep of the cladding onto the fuel pellet occurs due to the external pressure exerted on the rod by the coolant. Cladding compressive creep eventually results in hard fuel-cladding contact. During this period of fuel-cladding contact, changes in power level could result in significant changes [ in cladding stresses and strains. By using pre-pressurized fuel rods to partially offset the effect of the coolant external pressure, the rate of cladding-creep toward the surface of the fuel is reduced.
~
Fuel rod pre-pressurization delays the time at which substantial fuel-cladding interaction and hard contact occur and hence significantly reduces the number and extent of cyclic stresses and strains experienced by the cladding both before and after fuel-cladding contact. These factors result in an increase in the fatigue life margin of the cladding and lead to greater cladding reliability. If gaps should form in the fuel stacks, clad flattening will be prevented by the rod pre-pressurization so that the flattening time will be greater than the fuel core life. Steady State performance Evaluation In the calculation of the steadp state performance of a nuclear fuel rod, the following interacting factors must be considered: O v - 4.2-12
- 1. Clad creep and elastic deflection,
- 2. Pellet density changes, thermal expansion, gas release, and l thermal properties as a function of temperature and fuel burnup, ,
- 3. Internal pressure as a function of fission gas release, rod l t
geometry, and temperature distribution. ; These effects are evaluated using an overall fuel rod design model[3] , The model modifications for time dependent fuel densification are given l 1 in Reference 10. With these interacting factors considered, the 3
)
r.odel determines the fuel rod performance characteristics for a ' given rod geometry, power history, and axial power shape. In particular, internal gas pressure, fuel and cladding temperatures, and cladding deflections are calculated. The fuel rod is divided lengthwise into several sections and radially into a number of annular zones. Fuel density changes, cladding stresses, strains and deformations, and fission gas-releases are calculated separately for each segment. The effects are integrated to o,btain the internal rod pressure. The gap conductance between the pellet surface cnd the claddina inner diameter is calculated as a function of the composition, temperature, and pressure of the gas mixture, and the gap size or contact pressure between clad and pellet. After computing the fuel temperature for each pellet annular zone, the fractional fission gas release is assessed from the diffusion-trapping model described by lleisman, et al.E43 The total amount of gas released is based on the average fractional release within each axial and radial zone and the gas generation rate which in turn is a function of burnup. Finally, the gas released is summed over all zones and the pressure is calculated. The code shows good agreement in fit for a variety of published and proprietary data on fission gas release, fuel temperatures and I clad deflection.[3] Included in this spectrum are variations in power, time, fuel density, and geometry. The in-pile fuel temperature measurement comparisons used are referenced in Section 4.4.2.2. , O DECEMBER,1973 4.2-13
Typical fuel clad inner diameter and the fuel pellet outer diameter I- as a function of exposure are presented in Figure 4.2-4. The cycle to cycle changes in the pellet outer diameter represent the effects of power changes as the fuel is noved into different positions as a result of refueling. The gap size at any time is merely the difference between clad inner diameter and pellet outer diameter. Total clad-pellet surface contact occurs near the end of Cycle 2. The figure represents hot fuel dimensions for a fuel rod operating at the oower level shown in Figure 4.2-5. Figure 4.2-5 illustrates representative fuel rod internal gas pressure and linear power for the lead burnuo l- rod vs. irradiation time. In addition, it outlines the typical j operating range of internal gas pressures which is applicable to the total fuel rod population within a region. The "best estimate" fission gas relaase model was used in determining the internal gas pressures as a function of irradiation time. The plenum height of the fuel rod has been designed to ensure that the maximum internal pressure of the fuel rod will not exceed the design pressure of the reactor cqolant. The clad stresses during steady state operation are low. Compressive stresses are created by the pressure differential between the coolant pressure and the rod internal gas pressure. Because of the pre-l. pressurization, the compressive stresses are always less than410,000
- psi at the pre-pressurization level used in this fuel- rod desion.
l Tensile stresses could be created once the clad has come in contact with the pellet. These stresses would be induced by the fuel pellet swelling during irradiation. As shown in Figure 4.2-4, there is very limited clad pushout after pellet-clad contact. The extended time period over which clad pushout occurs allows for stress relief by clad creep and, as a result, the clad tensile stress is very low in the steady state. j' Tensile stresses could also be created if the internal gas pressure were higher than system pressure. The internal gas pressure criterion O 4.2-14
1 1 prevents the internal gas pressure from exceeding system pressure during steady state operation and, therefore, no tensile stress f 79 is created by high internal pressure. () ! Transient Evaluation Method A " modified CYGR0", which retains the basic design approach of l the referenced CYGR0[5] , is used as the Westinghouse design code to investigate the mechanical integrity of fuel rods during power transients. l l Modifications to the referenced CYGR0 were made in the following l areas: )
- 1. Consideration of an axial power shape, i
- 2. Calculation of temperatures in order to agree with Westinghouse j design codes and experimental data, !
fm 3. Calculation of fission gas pressure based upon the power time history,
- 4. Use of cladding creep and pellet behavior models consistent with Westinghouse data.
The " modified CYGR0" Westinghouse design code has been compared l to the fuel rod design model discussed in the Steady State Performance I Evaluation Section. In all cases, the conformity of the codes is satisfactory. Referring to Figure 4.2-4, as a result of power changes the fuel pellet outer diameter changes approximately 0.1 percent per kW/ft due to themal expansion. Power escalations or spikes early in life do not lead to hard clad-pellet interaction since the pellet merely expands into the gap. Power increases which occur after considerable gap closure will result in hard l clad-pellet interaction. The extent of the interaction detemines the clad stress level. Detailed clad stress analyses are strongly
,m affected by the control rod management which is selected during the final nuclear design and the specific fuel rod design. This information is not available at the time of the initial FSAR submittal.
l 4.2-15
- . . . - - . - . - - - . - - . ~ . ~ . - _ . - _ - - . ..- -
t The potential effects'of operation with waterlogged fuel are disdussed in Section 4.4.3.6. Waterlogging is not considered to be a concern during operational transients. 4.2.1.3.2 Fuel Assembly Structure Stresses and Deflections l The potential sources of high stresses in the assembly are avoided by the design. For_ example, stresses in the fuel rod due to tFermal l expansion and Zircaloy irradiation growth are limited by the relative motion of the rod as it slips over the grid spring and dimple surfaces. Clearances between the fuel rod ends and nozzles are provided so that Zircaloy irradiation growth will not result in end interferences. As another example, stresses due to holddown springs in cpposit; ion to the hydraulic lift force ar'e limited by the deflection character-istic of the springs. Stresses in the fuel assembly caused by tripping of the rod cluster control assembly have little influence on fatigue because of the small number of events during the life of an assembly. Assecbly components and prototype fuel assemblies made from production parts will be subjected to structural tests to verify that the design j bases requirements are met (see Section 1.5). j 1 l Dimensional Stability 1 l A prototype fuel assembly will be subjected to column loads in excess l of those expected in normal service and faulted conditions (see j Section 1.5). l l The coolant flow channels are established and maintained by the I structure composed of grids and guide thimbles. The lateral spacing between fuel rods is provided and controlled by the support dimoles of adjacent grid cells. Contact of the fuel rods on the dimples . is assured by the clamping force provided by the grid springs. Lateral motion of the fuel rods is opposed by the spring force and the internal moments generated between the spring and the support dimples. Grid i testing is discussed in Section 1.5. 4.2-16
l 7 Vibration and Wear l l The effect of a flow induced vibration on the fuel assembly and
)
individual fuel rods is minimal. The cyclic stress range associated l with deflections of such small magnitude is insignificant and has no effect on the structural integrity of the fuel rod.
]
The reaction on the grid support due to vibration motions is also correspondingly small and definitely much less than the spring l preload. Firin contact is therefore maintained. No significant wear of the cladding or grid supports is expected during the life of the fuel assembly, as described in the following section. Clad fretting and fuel rod vibration will be experimentally investi-gated (see Section 1.5).
- 4. 2.1. 3. 3 Operational Experience l
The fuel rods to be used in this reactor are very similar to the l fuel rods manufactured by Westinghouse in the past. The materials, manufacturing techniques, and quality control used are identical j to that used in manufacture of the larger diameter fuel rods. The ! current fuel rod clad thickness to diameter ratio is slightly increased. Application of the reactor operating experience of the larger diameter l Westinghouse manufactured fuel rods to the fuel in this reactor is appropriate because of the great similarity of the fuel rod l design and manufacture. A discussion of fuel operating experience ' 1 is given in Reference 9. 4.2.1.3.4 High Power Fuel Rod Development A high power fuel test program in progress has the object of defining failure limits for the combined effects of linear heat generation rate and burnup, providing increased assurance that current plants have an adequate performancy and design margin to falure threshold, and verifying the adequacy of design methods and computer codes. l The program is updated approximately every 6 months. This program is described in Section 8 of Reference 6. l 4.2-17 DECEPEER, 1973 h I
i i 4.2.1.4 Tests and Inspections ; 4.2.1.4.1 Quality Assurance Program The Quality Assurance Program Plan of the. Westinghouse Nuclear Fuel Division, as summarized in Reference 7, has been developed to serve the division in planning and monitoring its activities for the design and manufacture of nuclear fuel assemblies and associated components. l The program provides for control over all activities affecting product i quality, coninencing with design and development and continuing through ! procurement, materials handling, fabrication, testing and inspection, l storage, and transportation. The program also provides for the j indoctrination and training of personnel and for the auditing of ! i activities affecting product quality through a formal auditing program. 4.2.1.4.2 Manufacturing Quality Control philosophy is generally based on the following inspec-l tions being performed to a 95% confidence that at least 95% of the ; product meets specification, unless otherwise noted, using either l a hypergeometric function with zero defectives for small lots or the latest revision of Mil-105D for large lots. This confidence . level has been based on past experience-gained during the manufacturing of uranium cores. The following inspections are included. i 1. Component Parts l All parts received are inspected to a 95 x 95 confidence level. The characteristics inspected depends upon the component parts and includes dimensional, visual, check audits of test reports, material certification and non-destructive testing such as X-ray and ultrasonic. Westinghouse materials process and components specifications specify in detail the inspection to be perforned. i All material used in the manufacture of this core is accepted i and released by Quality Control. 4.z-18
i 2 Pellets Inspection is performed to a 95 x 95 confidence level for the dimensional characteristics such as diameter, density, length and squareness of ends. Additional visual inspections are performed for cracks, chips and porosity according to standards l established at the beginning of production. These standards are based upon standards used in previous cores which have in turn served as standards for over 125 million pellets manufactured and used in operating cores. Density is determined in terms I of weight per unit length and is plotted on zone charts used l l in controlling the process. Chemical analyses are taken on a daily sample basis throughout pellet production.
- 3. Rod Inspection Rod inspection consists of the following 100% non-destructive inspection and is based on the experience, specifications, procedures and standards established on previously manufactured and operated cores.
- a. Leak Testing Each rod is tested to a known leak using mass spectrometry with helium being the detectable gas. This is the system used previously on the leak test of over 500,000 rods.
- b. X-ray All fuel rod weld enclosures are X-rayed using weld
, correction forms. X-rays are taken in accord with ASTM I l 1 E-142-68, using 2-2T as the basis of acceptance.
- c. Dimensional j l l All rods are dimensionally inspected prior to final release and upgrading. The requirements include such items as ;
. length, camber, and visual inspection. l j 4.2-19
- d. Fluo roscope
(~N -100% of the fuel rods are irspected to insure V) proper plenum dimensions and that no significant gaps exist between pellets. , I
- e. Gamma Scanning 100% of the fuel rods are actively gamma scanned to verify the enrichment control.
- f. Traceability Full traceability of fuel rods and fuel rod components j is also established by Quality Control.
- 4. Assembly Inspection consists of 100 percent inspection for drawing l /N requi rements .
N)
- 5. Other Inspection i The following inspections are perfomed as part of the routine inspection operation:
- a. Measurements other than those specified above which are critical to thermal and hydraulic analyses are obtained to enable evaluation of manufacturing variations to a 95% confidence level,
- b. Tool and gage inspection and control including standard-ization to primary and secondary working standards.
i Tool inspection is performed at prescribed intervals on all serialized tools. Complete records are kept of calibration and condition of tools. rs i N l' 4.2-20 DECEMBER, 1973 l 4
1 I
- c. Check audit inspection of all inspection activities and '
records to assure that prescribed methods are followed l and that all records are correct and properly maintained. l
- d. Surveillance of outside contractors, including approval l
of standards and methods are performed where necessary. However, all final acceptance is based upon inspection performed by Westinghouse personnel. l l l 6. Process Control ; l i l To prevent the possibility of mixing enrichments during J l' fuel manufacture and assembly, strict enrichment segregation and meticulous process control are exercised, j l The U0 p wder is kept in sealed containers by blend. The I 2 contents are fully identified both by descriptive tagging and preselected color coding. A Westinghouse identification tag completely describing the contents is affixed to the containers before transfer to powder storage. Isotopic content is confirmed by sample isotopic analysis or 100% gamma scanning of the powder containers. i Powder withdrawal from storage can be made by only one authorized group, which directs the powder to the correct pellet production l line. All pellet production lines are physically separated 1 i from each other and pellets of only a single enrichment and l density-are produced in a given production line. Finished pellets are placed on trays having the same color code as the powder containers and transferred to segragated storage racks within the confiaes of the pelleting area. Samples l from each pellet lot are tested for isotopic content and ( impurity prior to acceptance by Quality Control. Physical barriers prevent mixing of pellets of different densities { and enrichments in this storage area. Unused powder and sub-O 4.2-21 l
standard pellets to be analyzed and reprocessed are returned to storage in the original color coded containers. Loading of pellets into the cladding is performed in isolated production lines and again only one density and enrichment is loaded on a line at a time. A serialized traceability sticker is placed on each fuel tube i which identifies the contract and enrichment. The sticker i is color coded to the original pellet tray code for visual identification. The end plugs are inserted; the bottom end plug is permanently identified to the contract and enrichment; and welded to seal the tube. The fuel tube remains color coded and traceability identified until just prior to installation in the fuel assembly. The color coding and end plug identification character and traceability stickers provide a cioss reference of the fuel contained in tne fuel rods. All fuel rods are gamma scanned over the full length for isotopic content prior to acceptance for assembly loading. l At the time of installation into an assembly, the color coding ; and traceability stickers are removed and a matrix is generated to identify each rod in its position within a given assembly. An inspector verifies that all fuel rods in an assembly have i the same end plug identification, and that the top nozzle to be used on the assembly carries the correct identification character describing the fuel enrichment and density for the core region being fabricated. The top nozzle identification ' then becomes the permanent description of the fuel contained in the assenbly. 4,,2.1.4.3 On-Site Inspection On-site inspection of fuel assenblies, control rods, and reactor internals will be provided in the Applicant's Safety Analysis Reports. O - 4.2-22
.. . . - - - . . . - . . - - - -_.- _ _ ~ - _ . - - . - . -
l 4.2.2 REACTOR VESSEL INTERNALS 4.2.2.1 Design Bases The design bases for the mechanical design of the reactor vessel internals components are as follows: l 1. The reactor intemals in conjunction with the fuel assemblies shall direct reactor coolant through the core to achieve acceptable , I flow distribution and to restrict bypass flow so that the heat ) transfer perfonnance requirements are met for all modes of operation. In addition, required cooling for the pressure vessel head shall be provided so that the temperature differences between tne vessel flange and head do not result in leakage from the flange during reactor operation.
- 2. In addition to neutron shielding provided by the reactor coolant, a separate neutron pad asserably is provided to limit the exposure of the pressure vessel in order to maintain the required ductility of the material for all modes of operation.
- 3. Provisions shall be made for installing in-core instrumentation useful for the plant operation and vessel material test specimens required for a pressure vessel irradiation surveillance program.
- 4. The core internals are designed to withstand mechanical loads ;
arising from operating basis earthquake, design basis earthquake and pipe ruptures and meet the requirement of Item 5 below.
- 5. The reactor shall have mechanical provisions which are sufficient to adequately support the core and internals and to assure that the core is intact'with acceptable heat transfer geometry following transients arising from abnormal operating conditions. ,
O 1 4.2-23 1
_ .. _ _. . . . ._. . __ -~ _ -- - - . . - - - l
- 6. Following the design basis accident, the plant shall be capable of being shutdown and cooled in an orderly fashion so that (n)
- v. fuel cladding temperature is kept within specified limits.
This inplies that the deformation of certain critical reactor internals must be kept sufficiently saall to allow core cooling. l The functional limitations for the core structures during the design basis accident are shown in Taule 4.2-1. To insure no column loading l of rod cluster control guide tubes, the upper core plate deflection i is limited to not exceed the value shown in Taole 4.2-1. 1 Details of the dynauic analyses, input forcing functions, and response i loadings are presented in Section 3.9. l 4.2.2.2 Description and Drawings The reactor vessel internals are described as follows: The components of the reactor internals are divided into three parts q consisting of the lower core support structure (including the entire core barrel and neutron shield pad assembly), the upper core support i structure and the in-core instrumentation support structure. The reactor internals support the core, maintain fuel alignment, limit I fuel assembly movement, maintain alignment between fuel assemblies and control rod drive mechanisms, direct coolent flow past the fuel elements, direct coolant flow to the pressure vessel head, provide gamma and neutron shielding, and guides for the in-core instrumentation. The coolant flows from the vessel inlet nozzles down the annulus between the core barrel and the vessel wall and then into a plenum at the bottom of the vessel. It then reverses and flows up through the core support and through the lower core plate. The lower core plate is sized to provide the desired inlet flow distribution to the core. After passing through the core, the coolant enters the region of the upper support structure and then flows radially to the core barrel outlet nozzles and directly through the vessel outlet nozzles. A small portion of the coolant flows between the baffle plates and the core barrel to provide additional cooling of the 4.2-24
g . . _ _ _ . . _ ._. . . _ _ - . _._._ _ -___ - ... _ _ _ . _ . _ _ . _ _ _ - _ barrel. Similarly, a small amount of the entering flow is directed l into the vessel head plenum and exits through the vessel outlet ! l nozzles, i i All the major material for the reactor internals is Type 304 stain-less steel. Parts not fabricated from Type 304 stainless steel ;
. include bolts and dowel pins which are fabricated from Type 316 i l stainless steel and radial support key bolts which are fabricated f of inconel. ;
All reactor internals are removable from the vessel for the purpose of their inspection as well as the inspection of the vessel internal surface. Lower Core Support Structure The major containment and support member of the reactor internals is the lower core support structure, shown in Figure 4.2-6A (three loop) and Figure 4.2-6B (four loop). This support structure assembly consists of the core barrel, the core baffle, and the lower core plate and support columns, the neutron shield pads, and tne core j support which is welded to the core barrel. ' All the major material ar this structure is Type 304 stainless steel. The lower core , i support structure is supported at its upper flange from a ledge in l the reactor vessel head flange and its lower end is restrained in ] its transverse movement by a radial support system attached to the i vessel wall. Within the core barrel are an txial baffle and a lower core plate, both of which are attached to the core barrel wall and I form the enclosure periphery of the assembled core. The lower core support structure and principally the core barrel serve to provide passageways and control . for the coolant flow. The lower core plate is positioned at the bottom level of the core below the baffle plates and provides support and orientation for the fuel assemblies. The lower core plate is a member through which the necessary flow distribution holes for each fuel assembly are machined. Fuel assembly 4.2-25 l
locating pins (two for each assembly) are also inserted into this plate. Columns are placed between this plate and the core support of the core barrel in order to provide stiffness and to transmit i the core load to the core support. Adequate coolant distribution l is obtained through the use of the lower core plate and core support. ( l l The neutron shield pad assembly consists of four pads that are bolted l and pinned to the outside of the core barrel. These pads are constructed of type 304 stainless steel and are approximately 36 to 48 inches ! wide by 148 inches long by 2.7 to 2.8 inches thick. The pads are located azimuthally to provide the required degree of vessel protection. Rectangular tubing in which material surveillance samples can be ! inserted and irradiated during reactor operation are a" ached to I the pads. The samples are held in the rectangular tubing by a preloaded spring device at the top and bottom to prevent sample movement.
- Additional details of the neutron shielding pads and irradiation l specinen holders is given in Reference [8]. I i
Vertically downward loads from weight, fuel assembly preload, control rod dynamic loading, hydraulic loads and earthquake acceleration
\- are carried by the lower core plate partially into the lower core l plate support flange on the core barrel shell and partially through ,
the lower support columns to the core support and thence through the core barrel shell to the core barrel flange supported by the vessel head flange. Transverse loads from earthquake acceleration, coolant cross flow, and vibration are carried by the core barrel shell to be distributed by the lower radial support to the vessel wall, and to the core barrel flange. Transverse acceleration of the fuel assemblies is transmitted to the core barrel shell by direct connection of the lower core plate to the barrel wall and by a radial support type connection of the upper core plate to slab sided pins pressed into the core barrel. l The main radial support system of the core barrel is accomplished ( by " key" and " keyway" joints to the reactor vessel wall. At equally spaced points around the circucaference, an Inconel block is welded O l 4.2-26 i
to the vessel inner dianeter. Another Inconel block is bolted to each of these blocks, and has a " keyway" geometry. Opposite each of these is a " key" which is attached to the internals. At assembly, as the internals are lowered into the vessel, the keys engage the keyways in 'the axial direction. With this design, the internals are provided witn a support at the furthest extremity, and may be viewed as a beam fixed at the top and simply supported at the bottom. Radial and axial expansions of the core barrel are acconinodated but transverse movement of the core barrel is restricted by tnis design. With this system, cyclic stresses in the internal structures are within the ASME Section III limits. In the event of an abnormal downward vertical displacement of the internals following a hypothetical failure, energy absorbing devices limit the displacement after contacting the vessel bottom head. The load is then transferred through the energy absorbing devices of the internals to the vessel. The energy absorbers, cylindrical in shape, are contoured on their bottom surface to the reactor vessel bottom head geometry. Their number and design are determined so as to limit the stresses imposed on all components except the energy absorber to less than yield (ASME Code Section III valves). Assuming a downward vertical displacement, the potential energy of the system is absorbed mostly by the strain energy of the energy absorbing devices. Upper Core Support Assembly The upper core support assembly, shown in Figures 4.2-7A and 8A (three loop) and 4.2-7B and 8B (four loop) consists of the top support plate, assembly, and the upper core plate between which are contained support columns and guide tube assemblies. The support columns establish the spacing between the top support plate assembly and the upper core plate and are fastened at top and bottom to these plates. The support columns transmit the mechanical loadings between the two plates and serve the supplementary function of supporting 4.2-27
thermocouple guides. The guide tube assemblies, sheath and guide g the control rod drive shafts and control rods. They are fastened () to the top support plate and are guided by pins in the upper core plate for proper orientation and support. Additional guidance for the control rod drive shaf ts is provided by the upper guide tube which is attached to the upper support plate and guide tube. The upper core support assembly, which is removed as a unit during refueling operation, is positioned in its proper orientation with respect to the lower support structure by slots in the upper core plant which engage flat-sided pins pressed into the core barrel. At an elevation in the core barrel where the upper core plate is positioned, the flat-sided pins are located at angular positions of 90 from each other. Four slots are milled into the core plate at the same positions. As the upper support structure is lowered into the lower internals, the slots in the plate engage the flat-sided pins in the axial direction. Lateral displacement of the plate and of the upper support assembly is restricted by this design. Fuel assembly locating pins protrude from the bottom () of the upper core plate and engage the fuel assemblies as the upper assembly is lowered into place. Proper alignment of the lower core support structure, the upper core support assembly, the fuel assemblies and control rods are thereby assured by this system of locating pins and guidance arrangement. The upper core support assembly ' is restrained from any axial movements by c large circumferential spring which rests between the upper barrel flange and the upper core support assembly and is compressed by the reactor vessel head flange. Vertical loads from weight, arthquake acceleration, hydraulic loads and fuel assembly preload are transmitted through the upper core plate via the support columns to the top support plate assembly and then the reactor vessel head. Transverse loads from coolant cross flow, earthquake acceleration, and possible vibrations are l distributed by the support columns to the top support plate and lg upper core plate. The top support plate is particularly stiff to () minimize deflection. 4.2-28 I
__ _ m _. _. . . . _ _ . _ _ _ - - . _ _ _ - _ . _ . . _ _ _ _ . _. . _ . .- . _ . . In-Core Instrumentation Support Structures
,m
( ) The in-core instrumentation support structures consist of an upper U. system to convey and supporc thermocouples penetrating the vessel through the head and a lower system to convey and support flux thimbles penetrating the vessel through the bottom (Figure 7.7-9 shows the Basic Flux-Mapping System). i i The upper system utilizes the reactor vessel head penetrations. ' Instrumentation port columns are slip-connected to in-line columns that are in turn fastened to the upper support plate. These port ; i columns protrude through the head penetrations. The thermocouples- l are carried through these' port columns and the upper support plate at positions above their readout locations. The thermocouple conduits are supported from the columns of the upper core support system. The thermocouple conduits are stainless steel tubes. l In addition to the upper in-core instrumentation, there are reactor vessel bottom port columns (See Figure 4.2-6) which carry the retractable, i n cold worked stainless steel flux thimbles that are pushed upward into the reactor core. Conduits extend from the bottom of the reactor vessel down through the concrete shield area and up to a thimble seal line. The minimum bend radii are about 144 inches and the trail-ing ends of the thimbles sat the seal line) are extracted approximately 15 feet during refueling of the reactor in order to avoid interference
*:. thin the core. -The thimbles are closed at the leading ends and serve as the pressure' barrier between the reactor pressurized water and the containment atmosphere.
Mechanical seals between the retractable thimbles and conduits are , provided at the seal line. During normal operation, the retractable thimbles are stationary and move only during refueling or for maintenance, at which time a space of approximately 15 feet above the seal line is cleared for the retraction operation. The in-core instrumentation support structure is designed for adequate ( support of instrumentation during reactor operation and is rugged 4.2-29
~ v 'a m-
enough to resist damage or distortion under the conditions imposed oy handling during the refueling sequence. These are tne only a l i conditions which affect the in-core instrumentation support structure. Reactor vessel surveillance specimen capsules are covered in Section l 5.4.3.6. That section is expanded in detail for a particular plant, as a part of the FSAR, and all the necessary details with regard to irradiation surveillance are added including a cross-section of the i reactor showing the capsule identity and location. 4.2.2.3 Design Loading Conditions The design loading conditions that provide the basis for the design of the reactor internals are:
- 1. Fuel Assembly Weight
- 2. Fuel Assembly Spring Forces
- 3. Internals Weight
- 4. Control Rod Scram (equivalent static load)
- 5. Differential Pressure
) 6. Spring Preloads
- 7. Coolant Flow Forces (static)
- 8. Temperature Gradients it 9. Differences in Thermal Expansion
- a. Due to temperature differences
- b. Due to expansion of different materials
- 10. Interference between components
- 11. Vibration (mechanically or hydraulically induced)
- 12. One or more loops out of service
- 13. All operational transients listed in Table 5.2-2
- 14. Pump overspeed
- 15. Seismic loads (operation basis earthquake and design basis earthquake)
- 16. Blowdown forces (due to cold and hot leg break)
The main objectives of the design analysis are to satisfy allowable stress limits, to assure an adequate design margin, and to establish O i t G 4.2-30
i ! I deformation limits which are concerned primarily with the functioning of the components. The stress limits are established not only [') to assure that peak stresses will not reach unacceptable values, but aise limit the amplitude of the oscillatory stress component j in consideration of fatigue characteristics of the materials. Both low and high cycle fatigue stresses are considered when the allowable i amplitude of oscillation is established. Dynamic analysis on the l reactor internals are provided in Section 3.9. As part of the evaluation of design loading conditions, extensive l testing and inspections are performed from the initial selection of raw materials up to and including component installation and ! plant operation. Among these tests and inspections are those performed during component fabrication, plant construction, startup and checkout, and during plant operation. 4.2.2.4 Design Loading Categories The combination of design loadings fit into either the normal, upset or faulted conditions as defined in the ASME Section III (] Code. Loads and deflections imposed on components due to shock and vibration are determined analytically and experimentally in both scaled models and operating reactors. The cyclic stresses due to these dynamic loads and deflections are combined with the stresses imposed by loads from component weights,, hydraulic forces and thermal gradients for the determination of the total stresses of the internals. The reactor internals are designed to withstand stresses originating from various operating conditions as summarized in Table 5.2-2. The scope of the stress analysis problem is very large requiring many differeri techniques and methods, both static and dynamic. The analysis performed depends on the mode of operation under consideration.
,/~\
4.2-31
Allowable Deflections ( For normal operating conditions, downward vertical deflection of ! the lower core support plate is negligible. i l For the loss of coolant accident plus the design basis earthquake condition, the deflection criteria of critical internal structures are the limiting values given in Table 4.2-1. The corresponding no loss of function limits are included in Table 4.2-1 for comparison purposes with the allowed criteria. l The criteria for the core drop accident are based upon analyses which have been performed to detennine the total downward displacement of the internal structures following a hypothesized core drop resulting from itss of the normal core barrel supports. The initial clearance between the secondary core support structures and the reactor vessel lower head in the hot condition is approximately one half inch. l An additional displacement of approximately 3/4 inch would occur due to strain of the energy absorbing devices of the secondary
. core support; thus the total drop distance is about 1-1/4 inches which is insufficient to permit the grips of the rod cluster control assembly to come out of the guide thimble in the fuel assemblies.
l l Specifically, the secondary core support is a device which will never be used, except during a hypothetical accident of the core i support (core barrel, barrel flange, etc.). There are 4 supports ) in each reactor. This device limits the fall of the core and absorbs the energy of the fall which otherwise would be imparted to the vessel. The energy of the fall is calculated assuming a complete and instantaneous failure of the primary core support and l is absorbed during the plastic deformation of the contrulled volume of stainless steel, loaded in tension. The maximum deformation of this austenitic stainless piece is limited to approximately 15 percent, after which a positive step is provided to insure support. I .
~
4.2-32 l l e
Fcr additicnal infcrmation on design loading categories, see Section 3.9. q 4.2.2.5 Design Criteria Basis l 1 The basis' for the design stress and deflection criteria is identified ! below: ) l I Allowable Stress .
. i l
For normal operating conditions, Section III of the ASME Nuclear { l Power Plant Components Code is used as a basis for evaluating acceptability ! of calculated stresses. Both static and alternating stress intensities are considered. For materials not covered by the code, c.g. cold worked type 316 stainless steel used as a bolting material, allowable stresses are established in the same manner as used in the code for materials of similar properties. It should be noted that the allowable stresses in Section III of the ASME code are based on unirradiated material properties. In view of the fact that irradiation l j increases the strength of the 304 stainless steel used for the ! j internals, although decreasing its elongation, it is considered
- l that use of the allowable stresses in Section III is appropriate and conservative for irradiated internal structures.
l l The allowable stress limits during the design basis accident used j for the core support structures are based on the January 1971 draft of the ASME Code for Core Support Structures, Subsection NG, and the Criteria for Faulted Conditions. 1 4.2.3 REACTIVITY CONTRCL SYSTEM 1 4.2.3.1 Design Bases Bases for temperature, stress on structural members, and material compatibility are imposed on the design of the reactivity control components. 4.2-33 { 1 ( -
4.2.3.1.1 Design Stresses The reactivity control system is designed to withstand stresses originating from various operating conditions as summarized in Tabl e 5.2-2. Allowable Stresses: For normal operating conditions Section III ' of the ASME Boiler and Pressure Code is used. All components are l analyzed as Class I components under Article 14B-3000. Dynamic Analysis: The cyclic stresses due to dynamic loads and deflections are combined with the ' stresses imposed by loads from component weights, hydraulic forces and thermal gradients for the determination of the total stresses of the reactivity control system. 4.2.3.1.2 Material Compatibility Materials ~ are selected for compatibility in a Pressurized Water Reactor environment, for adequate mechanical properties at room and operating temperature, for resistance to adverse property changes in a radioactive environment, and for compatibility with interfacing components. 4.2.3.1.3 Reactivity Control Components The reactivity control components are subdivided into two categories:
- 1. Permanent devices used to control or monitor the core and, i
- 2. Temporary devices used to control or monitor the core, i The permanent type components are the full and partial length rod i l cluster control assemblias, control rod drive assemblies, neutron I source assemblies, and thimble plug assemblies. Although the thimble plug assembly does not directly contribute to the reactivity control i O
4.2-34 l
,- . ,,m. - -- - . -
of the reactor, it is presented as a reactivity control system ! component in this document because it is needed to restrict bypass j f) flow through those thimbles not occupied by absorber, source or 1 V burnable poison rods. I i l The temporary component is the burnable poison assembly which is ' normally used only in the initial core. The design bases for each of the mentioned components are in the following paragraphs. Absorber Rods The following are considered design conditions under Article NB-3000 of the ASME Boiler and Pressure Vessel Code Section III.
- 1. The external pressure equal to the Reactor Coolant System operating pressure. l
- 2. The wear allowance equivalent to 1,000 reactor trips. l
/ 3. Bending of the rod due to a misalignment in the guide tube.
- 4. Forces imposed on the rods during rod drop, 1
- 5. Loads caused by accelerations imposed by the control rod drive mechanism.
- 6. Radiation exposure for maximum core life.
The absorber material temperature shall not exceed its melting temperature (1470*F for Ag-In-Cd absorber material)b23 Burnable Poison Rods The burnable poison rod clad is designed as a class I component under Article NB-3000 of the ASME Boiler and Pressure Vessel code, Section III,1973 for Conditions I and II. For abnormal loads (
%)
DECEMBER, 1973 4.2-35
during Conditions III and IV code stresses are not considered limiting. Failures of the burnable poison rods during these conditions must l C not interfere with reactor shutdown or emergency cooling of the
\
fuel rods. The burnable poison absorber material is non-structural. The structural l elements of the burnable poison rod are designed to maintain the absorber geometry even if the absorber material is fractured. The rods are designeo so that the absorber material is below its softening temperature.
-m_.
(1492 F* for reference 12.5 w/o boron rods). In addition. tne structural elements are designed to prevent excessive slumping. Neutron Source Rods l The neutron source rods are designed to withstand tne following: 1
- 1. The external pressure equal to the Reactor Coolant System operating pressure and An internal pressure equal to the pressure generated by released
]
2. gases over the source rod life. Thirable Plug Assembly l The thimble plug assemblies satisfy the following:
- 1. Acconnodate the differential thennal expansion between the l fuel assembly and the core internals, l
- 2. Maintain positive contact with the fuel assembly and the core internals.
l
- Borosilicate glass is accepted for use in burnable poison rods if the softening temperature is 1510 1,18 F. The softening temperature is defined in ASTM C 338.
r3 4.2-36 l
I i
- 3. Be inserted into or withdrawn from the fuel assembly by a
,q force not exceeding 25 pounds.
b 4.2.3.1,4 Control Rod Drive Mechanisms The mechanisms are Class I components designed to meet the stress requirements for normal operating conditions of Section III of the ASME Boiler and Pressure Vessel Code. Both static and alternating stress intensities are considered. The stresses originating from the required design transients are included in the analysis. A dynamic seismic analysis is required on the full length control rod drive mechanism when a seismic disturbance has been postulated to confirm the ability of the mechanism to meet ASME Code, Section III allowable stresses and to confirm its ability to trip when subjected to the seismic disturbance. The part length control rod drive mechanism meets only the stress limits defined in the ASME Code, Section III, in order to maintain L structural integrity when subjected to seismic loads since it is , a non-tripping mechanism. Full Length Control Rod Drive Mechanism Operational Requirements The basic operational requirements for the full length Control Rod Drive Mechanisms are:
- 1. 5/8-inch step,
- 2. 150-inch travel,
- 3. ' 360-pound maximum load, l 4. Step in or out at 45 inches / min (72 steps / min),
p 5. Power interruption shall initiate release of drive rod assembly, L 4.2-37 I
- 6. Trip delay of less than 150 ms - Free fall of drive rod assembly O shall begin less than 150 ms after. power interruption no matter L
what holding or stepping action is being executed with any )
-load and coolant temperatures of 100*F to 550"F.
'7 ' 40-year design life with normal refurbishment,
- 8. .28,000 complete travel excursions which is 13 x 100 steps with normal refurbishment. I Part Length Control Rod Drive Mechanism Operational Requirements The basic operational requirements for the part length Control' Rod Drive Mechanisms'are:
- 1. 3/8 inch / revolution leadscrew translation,
- 2. 150 inch travel, .
ID D ~'
- 3. - Lifting Capacity, (a). ' Design Life Load' - 150 pounds. at rated speed I
.(b) Performance Total Load - 700 pounds at rated speed, (c)'. Maximum Lif1 Force - 1400 pounds at stall- .
l
- 4. Minimum incremental movement of 0.050 inches.
- 5. Shall operate in or out at 15 inches / min. )
i
~
6.. Power interruption shall cause the leadscrew to remain in l position no matter if holding or moving at any load at coolant temperatures of 100'F to 550 F.
- 7. ' 40-year design life'with normal refurbishment.
;d LU 4.2-38 l
l l-l 8. 28,000 complete in-out travel excursions with normal refurbish- ; ment.
'v _ i
- 9. 4000 couplete in-out trave 1 excursions without refurbishment. !
l 4.2.3.2 Design Description Reactivity control is provided by neutron absorbing rods and a soluble chemical' neutron absorber (boric acid). The boric acid concentration is varied to control long-tenn reactivity changes j such as: 1
\
- 1. Fuel ' depletion and fission product buildup.
- 2. Cold to hot, zero power reactivity change.
- 3. Reactivity change produced by intermediate-term fission products such as xenon and samarium.
rn d 4. Burnable poison depletion Chemical and Voleme Control is covered in Section 9.3.4. The rod cluster control assemblies provide reactivity control for:
- 1. Shutdown,
- 2. Reactivity changes due to coolant temperature changes in the power range.
- 3. Reactivity changes associated with the power coefficient of reactivity.
- 4. Reactivity changes due to void fonnation.
4.2-39
The first fuel cycle contains more excess reactivity than subsequent : cycles due to the loading of all fresh (unburned) fuel. If soluble i (V] poron were the sole means of control, the moderator temperature : coefficient would be positive. It is desirable to have a negative , moderator temperature coefficient throughout the entire cycle in - order to re' duce possible deleterious effects caused by a positive coefficient during loss of coolant or loss of flow accidents. This is accomplished by installation of burnable poison assemblies. Tbe neutron source assemblies provide a means of monitoring the core during periods of low neutron activity. J The most effective reactivity control components are the full and partial length rod cluster control assemblies and their corresponding drive rod assemblies which are the only kinetic parts in the reactor. Figure 4.2-9 identifies the full length rod cluster control and drive rod assembly, in addition to the arrangement of these components in the reactor relative to the interfacing fuel assembly, guide tubes, and control rod drive mechanism. In the foilowing paragraphs,
' each reactivity control component is described in detail.
4.2.3.2.1 Reactivity Control Components Full Length Rod Cluster Control Assembly [ The full length. rod cluster control assemblies are divided into two categories: control and shutdown. The control groups compensate
.for reactivity changes due to variations in operating conditions of the reactor, i.e., power and temperature variations. Two criteria have been employed for selection of the control groups. First the total reactivity worth must be adequate to meet the nuclear requirements of the reactor. Second, in view of the fact that some of these. rods may be partially inserted at power operation, the total power _ peaking factor should be low enough to ensure that the power t
!A
- O-
, 4.2-40 l l 9 q - -
capability is met. The control and shutdown aroups provide adequate shutdown margin which is defined as the amount by which the core would be subcritical at hot shutdown if all rod cluster control (V] assemblies are tripped assuming that the hiahest worth assembly remains fully withdrawn and assumina no changes in xenon or boron concentration or part length rod cluster control position. A rod cluster control assembly comprises a aroup of individual neutron absorber rods fastened at the top end to a common spider assembly, as illustrated in Figure 4.2-10. The absorber material used in the control rods is silver-indium-cadmium alloy which is essentially " black" to thermal neutrons and has sufficient additional resonance absorption to significantly increase its worth. The alloy is in the form of extruded rods. which are sealed in stainless steel tubes to prevent the rods from coming in direct contact with the coolant. In construction, the 4 silver-indium-cadmium rods are inserted into cold-worked stainless steel tubing which is then sealed at the bottom and the top by welded end plugs as shown in Figure 4.2-11. Sufficient diametral (~} C and end clearance is provided to acconnodate relative thermal expansions. The bottom plugs are made bullet-nosed to reduce the hydraulic drag during reactor trip and to guide smoothly into the dashpot section of the fuel assembly guide thimbles. The upper plug is
- threaded for assembly to the spider and has a reduced end section to make the joint more flexible.
The spider assembly ~is in the form of a central hub with radial vanes containing cylindrical fingers from which the absorber rods are sus-pended. Handling detents and detents for connection to the drive rod assembly are machined into the upper end of the hub. A coil spring inside the spider body absorbs the impact energy at the end of a trip insertion. The radial vanes are joined to the hub by tack welds and brazing and the fingers are joined to the vanes by furnace O 4.2-41
brazing. A centerpost which holds the spring and its retainer
'is threaded into the hub within the skirt and welded to prevent O loosening in service. All components of the spider assembly are made from type 304 and 308 stainless steel except for the retainer which is of 17-4 PH material and the springs which are Inconel 718 alloy or oil tempered carbon steel where the springs do not contact the coolant.
The absorber rods are fastened securely to the spider to assure trouble free service. The rods are first threaded into the spider fingers and then pinned to maintain joint tightness, after which the pins are welded in place. The end plug below the pin position is designed with a reduced section to pennit flexing of the rods to correct for small operating or assembly misalignments. The overall-length is such that when the assembly is withdrawn through its full travel the tips of the absorber rods remain engaged in the guide thimbles so that alignment between rods and thimbles is always maintained. Since the rods are long and slender, they C) the guide thimble. Part length Rod Cluster Control Assembly The function of the part length rods is to control axial neutron flux shape and axial xenon oscillations should they occur. The part length rods are on manual control. The axial position of the part length rods depends on the full length rod insertion, power level, xenon distribution, etc. Since generally a considerable portion of the neutron power generation takes place in the region
~
containing the part length rods, their effect on possible reactivity insertion rates and power shapes is significant. The part length rod cluster control assemblies as shown in Figure 4.2-12 are identical to the full length rod cluster control assemblies 4.2-42 L I
--i.-...-..i.>w..r -i ' '
! l except for two physical differences. The most obvious difference is the attachment configuration for the drive rod assembly. The forward end of the spider hub is slightly longer than the control rod spider to accommodate a slot which matches a key in the drive rod assembly. This key / slot arrangement prevents the drive rod assembly from rotating. The movement of the roller nut arrangement in the part length mechanism results in only translation of the l part length drive rod assembly. The other difference is that only l the bottom 36 inches of each rod contains the silver-indium-cadmium ) absorber material. The balance of the void length is partially filled with aluminum oxide to compensate for the. void left by the reduced amount of absorption material. The same diametral clearances are provided as in the full length i absorber rods thus giving somewhat similar temperatures in the absorber material. Also the same end plugs are used giving the same mechanical attachment of the rod to the spider fingers as in the full length absorber rods.
.] The materials are identical to those of the full length rod cluster control assemblies except for the aluminum oxide in the absorber rod. The aluminum oxide is intended to act only as a filler material and is not intended to absorb neutrons.
? j ; -
. . . ' ._3.t-c .. . ..
Each burnable poison assembly consists of burnable poison rods attached to a hold down assembly. Conceptual burnable poison assemblies are shown in Figure 4.2-13. The poison rods consist of borosilicate glass tubes contained within Type 304 stainless steel tubular cladding which is plugged and seal welded at the ends to encapsulate the glass. The glass is i ("\ 4.2-43
l t , 1 also supported along the . length of its inside diameter by a thin l7 wall tubular inner liner of Type 304 stainless steel. The top (f end of the liner is open to permit the diffused helium to pass ' into the void. volume $d the liner overhangs the glass. The liner has an outward flange at the bottom end to maintain the position of the liner with the glass. A typical burnable poison rod is shown in longitudinal and transverse cross-sections in Figure 4.2-14. The rods are statically suspended and positioned in selected guide thimbles within specified fuel assemblies. The poison rods in each fuel assembly are grouped and attached together at the I ! top end of the rods to a hold down assembly by a flat, perforated retaining plate which fits within the fuel assembly top nozzle and rests on the adaptor plate. The retaining plate (and the poison I t rods) is held down and restrained against vertical motion through ; a spring pack which is attached to the plate and is compressed by the j
- upper core plate when the reactor upper internals assembly is lowered
- into the . reactor. This arrangement assures that the poison rods f~ cannot be ejected from the core by flow forces. Each rod is permanently
,I attached to the base plate by a nut which is lock welded into place. ] The clad in the rod assemblies is 10 percent cold worked Type 304 stainless steel. All other structural materials are 304 or 308 stainless steel except for the springs which are Inconel 718 1 The borosilicate glass tube provides sufficient boron content to l meet the criteria discussed in Section 4.3.1. Neutron Source Assembly The purpose of the neutron source assembly is to provide a base neutron level to insure that the detectors are operational and responding to core multiplication neutrons. Since there is very little neutron activity during loading, refueling, shutdown, and approach to criticality, a neutron source is placed in the reactor to provide a positive neutron count of at least 2 counts per second on the source range
.f . . '~
4.2-44 DECEMBER, 1973 O
detectors attributable to core neutrons. The detectors, called g source range detectors, are used primarily when the core is subcritical V and during special suberitical modes of operations. The source assembly also permits detection of changes in the core multiplication factor during core loading refueling, and approach to criticality. This can be done since the multiplication factor is related to an inverse function of the detector count rate. Therefore a change in the multiplication factor can be detected during addition of fuel assenblies while loading the core, a change in control rod positions, and changes in baron concentration. Both primary and secondary neutron source rods are used. The primary source rod, containing a radioactive material, spontaneously emits neutrons during initial core loading and reactor startup. After the primary source rod decays beyond the desired neutron flux level, neutrons are then supplied by the secondary source rod. The secondary source rod, contains a stable material, which must be activated by neutron bombardment during reactor operation. ( ,/ The activation results in the subsequent release of neutrons. This 1 becomes a source of neutrons during periods of low neutron flux, such as during refueling and the subsequent startups. The three loop reactor ct re employs two source assemblies. Each source assembly contains one primary source rod and three secondary source rods. A conceptual source assembly is shown in Figure 4.2-15A. The four loop reactor core employs four source assemblies; two primary source assemblies and two secondary source assemblies. Each primary source assembly contains one primary source rod and between zero and twenty-three burnable poison rods. Each secondary source assembly contains a symmetrical grouping of four secondary source rods and between zero and twenty burnable poison rods. Locations not filled with a source or burnable poison rod contain a thimble
, plug. Conceptual source assemblies are shown in Figures 4.2-15B
() and 4.2-15C. DECEMBER, 1973 4.7 45
Neutron source assemblies are employed at diametrically opposite sides of the core. The assemblies are inserted into the rod cluster control guide thimbles in fuel assemblies at selected unrodded locations. ! The other source assemblies contain a holddown assembly identical to that of the burnable poison assembly. [ The primary and secondary source rods bcth utilize the same cladding material as the absorber rods. The secondary source rods contain Sb-Be pellets stacked to a height of approximately 88 inches. The primary source rods contain capsules of Californium (Pu-Be is j ! a possible alternate) source material and alumina spacer rods ) to position the source material within the cladding. The rods in each assembly are permanently fastened at the top end to a l
' holddown assembly, which is identical to that of the burnable poison assemblies.
The other structural members are constructed of type 304 stainless steel except for the springs. The springs exposed to the reactor coolant are wound from an age hardened nickel base alloy for corrosion resistance and high strength. The springs, when contained within the rods where corrosion resistance is not necessary, are oil ! tempered carbon steel. i Thimble Plug Assembly In order to limit bypass flow through the rod cluster control guide thimbles.in fuel assemblies which do not contain either control rods, source rods, or burnable poison rods, the fuel assemblies at those locations are fitted with thimble plug assemblies. The thimble plug assemblies as shown in Figure 4.2-16 consist of a flat base plate with short rods suspended from the bottom surface i 4.2-46 l 4
and a spring pack assesly. The twenty-four short rods, called l ,, thimble plugs, project into the upper ends of the guide thisles ( to reduce the bypass flow area. Similar short rods are also used on the source assemblies and burnable poison assemblies to plug the ends of all vacant fuel assembly guide thimbles. At installation in the core, the thimble plug assemblies interface with both the upper core plate and with the fuel assembly top nozzles by resting on the adaptor plate. The spring pack is compressed by the upper core plate when the upper internals assembly is lowered into place. Each thinble plug is permanently attached to the base plate by a nut which is locked to the threaded end of the plug by a small lock-bar welded to the nut. All components in the thimble plug assembly, except for the springs, are constructed from type 304 stainless steel. The springs are wound from an age hardened nickel base alloy for corrosion resistance and high strength. _ 4.2.3.2.2 Control Rod Drive Mechanism
\
(G All parts exposed to reactor coolant are made of metals which resist the corrosive action of the water. Three types of metals are used exclusively: stainless steels, Inconel-X and cobalt based alloys. Wherever magnetic flux is carried by parts exposed to the main coolant, 400 series stainless steel is used. Cobalt based alloys are used for the pins and latch tips. Inconel-X is used for the springs of both latch assemblies and 304 stainless steel is used for all pressure containing parts. Hard chrome plating provides wear surfaces on the sliding parts and prevents galling between mating parts. A position indicator assembly slides over the full length control rod drive mechanism rod travel housing. It detects the drive l rod assembly position by means of 42 discrete coils that magnetically sense the entry and presence of the rod drive line through its i center line over the normal length of the drive rod travel. s'~' J DECEMBER,1973 4.2-47
l Similarly, a position indicator asscely slides over the motor l l tube of the part length control rod drive mechanism and provides
- j a means of detecting the leadscrew position. '
C Full Length Control Rod Drive Mechanism Control rod drive mechanisms are located on the dome of the reactor ; I vessel. They are coupled to rod contml clusters which have absorber material over the entire length of the control rods and derive l their Hme from this feature. The full length control rod drive mechanism is shown in Figure 4.2-17 and schematically in Figure j l 4.2-18. The primary function of the full length control rod drive mechanism is to insert or withdraw rod control clusters within the core to contml average core temperature and to shut down 1.he reactor. The full length control rod drive mechanism is a magnetically operated jack. A negnetic jack is an arrangenent of three electro-magnets which are energized in a controlled sequence by a power cycler to insert or withdraw rod control clusters in the reactor core in discrete steps. The control rod drive mechanism consists of four separate subassedlies. They are the pressure vessel, coil stack assedly, the latch assedly, and the drive rod assedly.
- 1. The pressure vessel includes a latch housing and a rod travel housing which are connected by a threaded, seal welded, maintenance joint which facilitates replacement of the latch assedly. The !
closure at the top of the rod travel housing is a threaded plug j with a canopy seal weld for pressure integrity. The late.h housing is the lower portion of the vessel and contains the latch assedly. The rod travel housing is the upper portion ! of the vessel and provides space for the drive rod during its upwaro movement as the contml mds are withdrawn fron: C the core. DECEMPER, 1973 4.2-48 l
- 2. The coil stack assembly includes the coil housings, an electrical conduit and connector, and three operating coils; 1)the r3 i
'Q stationary gripper coil, 2) the moveable gripper coil, and
- 3) the lift coil.
The coil stack assembly is a separate unit which is installed on the drive mechanism by sliding i+ over the outside of the latch housing. It rests on the base of the latch housing without mechanical attachment. Energizing of the operation coils causes movement of tne pole pieces and latches in the latch assembly. 1
- 3. The latch assembly includes the guide tube, stationary pole I pieces, moveable pole pieces, and two sets of latches; 1) l the moveable gripper latch, and 2) the stationary gripper latch.
The latches engage grooves in the drive rod assembly. The l(q g moveable gripper latches are moved up or down in 5/8 inch steps by the lift pole to raise or lower the drive rod. The stationary gripper latches hold the drive rod assembly while the moveable gripper latches are repositioned for the next l 5/8 inch step. 4 The drive rod assembly includes a flexible coupling, a drive ] rod, a disconnect button, a disconnect rod, and a locking l- button. l l l The drive rod has 5/8 inch grooves which receive the latches during holding or moving of the drive rod. The flexible coupling is attached to the drive rod and produces the means for coupling to the rod control cluster assembly. 4 l dp. 4.2-49 1 a
~ . - . . . - -.
The disconnect button, disconnect rod, and locking button ' provide positive locking of the coupling to the rod control [ cluster assembly and permits remote disconnection of the drive rod. The control rod drive mechanism is a trip design. Tripping can j occur during any part of the power cycler sequencing if power to the coils is interrupted.- The control rod drive mechanism is threaded and seal welded on an adaptor on top of the reactor vessel and is coupled to the rod control cluster assembly directly below. The mechanism is capalle of handling a 360 pound load, including , the drive rod weight, n - of 45 inches / minute. Withdrawal of the rod control cluster is accomplished by magnetic forces while insertion is by gravity. l
- The mechanism internals are designed to operate in 650"F reactor
-coolant. The pressure vessel is designed to contain reactor coolant
, at 650 F and 2500 psia. The three operating coils are designed to operate at 392*F with forced air cooling required to maintain that temperature. The full length control rod drive mechanism shown schematically in Figure 4.2-18 withdraws and inserts its control rod as electrical pulses are received by the operator coils. An ON or 0FF sequence, I- repeated oy silicon controlled rectifiers in the power prograniner, j causes either withdrawal or insertion of the control rod. Position ! of the control rod is measured by 42 discrete coils mounted on J the position indicator assembly surrounding the rod travel housing. ! Each coil magnetically senses the entry and presence of the top
, of the ferro-magnetic drive rod assembly as it moves through the coil center line.
4 4.2-50
. - . ..- . . .- . - . - - . ~ - . . _ . . - . . --
During plant operation the stationary gripper coil of the drive mechanism holds the control rod withdrawn from the core in a static I Q position until the movable gripper coil is energized. Rod Cluster Control Assembly Withdrawal The control rod is withdrawn by repetition of the following sequence of events: l
- 1. Movable Gripper Coil (B) - ON ;
1 !
- The latch locking plunger raises and swings the movable gripper latches into the drive rod assembly groove. A 1/16 inch axial clearance exists between the latch teeth and the drive rod.
2 Stationary Gripper Coil (A) - 0FF The force of gravity, acting upon the drive rod assembly and attached control rod, causes the stationary gripper latches and plunger to move downward 1/16 inch until the load of the drive rod assembly and attached control rod is transferred to the movable gripper latches. The plunger continues to move downward and swings the stationary gripper latches out of the drive rod assembly groove. 3 Lift Coil (C) - ON The 5/8 inch gap benveen the movable gripper pole and the lift pole closes and the drive rod assembly raises one step length (5/8 irch). 4 Stationary Gripper Coil (A) - ON l The plunger raises and closes the gap below the stationary gripper pole. The three links, pinned to the plunger, swing i A. l-(f 4.2-51 I
l i and the stationary gripper latches into a drive rod assembly i l
,._ groove. The latches contact the drive rod assembly and lif t j
(/ it (and the attached control rod) 1/16 inch. The 1/16 inch vertical drive rod assembly movement transfers the drive rod assembly load from the movable gripper latches to the stationary gripper latches. ) l
- 5. Movable Gripper Coil (B) - 0FF i l
The latch locking plunger separates from the movable gripper pole under the force of a spring and gravity. Three links, pinned to the plunger, swing the three aiovable gripper latches I out of the drive rod assembly groove. I
- 6. Lift Coil (C) - 0FF The gap between the movable gripper pole and lift pole opens. l The movable gripper latches drop 5/8 inch to a position adjacent l to a drive rod assembly groove.
.-(3 C/
'7. Repeat Step 1 The sequence described above (1 thru 6) is termed as one step or one cycle. The control rod moves 5/8 inch for each step or cycle. The sequence is repeated at a rate of up to 72 steps per minute and the drive rod assembly (which has a 5/8 j inch groove pitch) is raised 72 grooves per minute. The control rod is thus withdrawn at a rate up to 45 inches per minute.
Rod Cluster Control Aswably Insertion i Tne sequence for control rod insertion is similar to that for control rod withdrawal, except the timing of lift coil (C) ON and 0FF is changed to permit lowering the control rod. uO 4.2-52
)
- 1. Lift Coil (C) - ON O
(f . The 5/8 inch gap between tb, movable gripper and lift pole closes. The movable gripper latches are raised to a position adjacent to a drive rod assembly groove.
- 2. Movable Gripper Coil- (B) - ON The latch locking plunger raises and swings the movable gripper latches into a drive rod assembly groove. A 1/16 inch axial clearance exists between the latch teeth and the drive rod assembly.
l
- 3. Stationary Gripper Coil (A) - 0FF The force of gravity, acting upon the drive rod assembly and attached control rod, causes the stationary gripper latches and plunger to' move downward 1/16 inch until the load of the drive rod assembly and attached control rod is transferred j to the movable gripper latches. The plunger continues to l ()
move downward and swings the stationary gripper latches out l of the drive rod assembly groove.
- 4. Lif t Coil (C) - 0FF l
The force of gravity separates the movable gripper pole from the lift pole and the drive rod assembly and attached control j rod drop down 5/8 inch.
- l. 5. Stationary Gripper (A) - ON l ;
i The plunger raises and closes the gap below the stationary gripper pole. The three links, pinned to the plunger, swing the three stationary gripper latches into a drive rod assembly
\
v 4.2-53
groove. The latches contact the drive rod assembly and lif t /~~N it (and the attached control rod) 1/16 inch. The 1/16 inch U vertical drive rod assembly movement transfers the drive rod ) assembly load from the movable gripper latches to the stationary gripper latches.
- 6. Movable Gripper Coil (B) - 0FF The latch locking plunger separates from the movable gripper pole under the force of a spring and gravity. Three links, !
pinned to the plunger, swing the three movable gripper latches out of the drive rod assembly groove.
- 7. Repeat Step 1 The sequence is repeated, as for control rod withdrawal, up to 72 times per minute which gives a control rod insertion rate of 45 inches per minute. !
CT ; C Holding and Tripping of the Control Rods I During most of the plant operating time, the control rod drive mechan-isms hold the control rods withdrawn from tha ccre in a static position. In the holcing mode, only one coil, the stationary gripper coil (A), is energized on each mechanism. The drive rod assembly and attached control rod hang suspended from the three latches. If power to the stationary gripper coil is cut off, the combined weight of the drive rod assembly and the rod cluster control assembly is sufficient to move latches out of the drive rod assembly groove. The control rod falls by gravity into the core. The trip occurs as the magnetic field, holding the stationary gripper plunger half against the stationary gripper pole, collapses and the stationary gripper plunger half is forced down by the weight acting upon the latches. After the drive rod assembly is released by the mechanism, (] it falls freely until the control rods enter the buffer section kJ of their thimble tubes. 4.2-54
Part Length Control Rod Drive Mechanism
,a
( ) Control rod drive mechanisms are located on the dome of the reactor a vessel head. Part length control rod drive mechanism's are coupled to rod control clusters which have absorber material only on the lower 25% of the length of the control rods and derive the name part length from this feature. The part length control rod drive I mechanism is shown in Figure 4.2-19. The primary function of the part length control rod drive mechanism is to position rod control cluster tips within the core to control the axial power distribution during xenon transients and to limit 1 axial xenon oscillations should they occur. The part length control rod drive mechanism is an electro-mechanical roller-nut mechanism. The control rod drive mechanism consists of fuor separate subassemblies.
,, They are the pressure vessel, the stator, the rotor, and the translating
) leadscrew.
- 1. The pressure vessel includes an adapter and a motor tube which are connected by a threaded, seal welded maintenance joint which facilitates replacement of the control rod drive mechanism internals.
The adapter is the lower portion of the vessel and contains the lower rotor assembly. The motor tube is the upper portion of the vessel. It contains the upper rotor and provides space for the leadscrew during its upward movement as the control rods are withdrawn from the core.
- 2. The stator assembly includes a cooling shroud, electrical conduit and connector, and the basic stator jacket, stack, and windings.
L) DECEMBER, 1973 4.2-55
l l Stator energization causes the brake arms on the rotor to swing outward and release the mechanical brake. When the stator windings are energized in a controlled sequence by
- - a power cycler, a rotating magnetic field is produced within the rotor causing it to rotate.
- 3. The rotor assembly includes a thru< t bearing, roller assembly, lower rotor, upper rotor, brake arms, and a radial bearing.
The rotary motion of the. rotor which is induced by the stator causes linear travel of the leadscrew. 4 The leadscrew assembly includes a flexible coupling, a leadscrew, a disconnect button, a disconnect rod, and a locking button. Tne leadscrew has threads which are engaged with the roller assembly and cause linear travel of the leadscrew when the rotor rotates. The leadscrew travel is 3/8 inch /rev. The flexible coupling is attached to the leadscrew and provides j the means for coupling to the rod control cluster assembly. The disconnect button, disconnect end, and locking button provide positive locking of the coupling to the rod control cluster assembly and permits remote disconnection of the leadscrew. l The control rod drive mechanism is a non-trip design. When power to the stator is interrupted, the mechanical brake arms swing in and engage the mechanical brake which stops the leadscrew and the rod control cluster assembly motion. ! The control rod drive mechanism is threaded and seal welded to an adaptor on the top of the reactor vessel and is coupled to the od control cluster assembly directly below. I + i i I O l 4.2-56
- =
F )
- i i
! l
The mechanism is capable of developing a 700 pound lifting force at a rate of 15 inches / minute. The insertion force is limited by design to the weight of the leadscrew and tne rod control cluster assembly. The motor stalling torque is 12 foot-pounds. The mechanism internals are designed to operate in 650 F reactor cool ant. The pressure vessel is designed to contain reactor coolant at 650 F and 2500 psia. The stator windings are designed to operate at 300 F with forced air cooling required to maintain that temperature. The internal rotor assembly is the operating center of the mechanism. The lead screw assembly is permanently attached to the rotor assembly and is disconnected with special tools only in the case of exceptional maintenance. During refueling, the lead screw is disconnected from the control rod and driven into the travel housing for storage. The rotor assembly is free to rotate in and held in place within the pressure vessel by ball bearing assemblies. Five free rotating
" rollers" are held captive in the lower cylindrical portion of the rotor and are canted to match the lead angle of the drive rod threads. As the internal rotor assembly rotates, the roller-nut turns within the threads of the drive rod, translating vertical motion to it, much as a turning nut would cause a bolt to rise or fall in a slot which prevents the bolt from rotating. In the case of the drive mechanism, the rotational torque is taken by the rod cluster control assembly in the core.
e In operation, at least two of the six windings are energized by the sequencing to obtain the required motion. In the holding mode, two or three of the windings are energized thereby holding the final position of the armature and preventing rod motion even though the brakes are disengaged. Loss of power to the mechanism will engage the brakes and prevent rod motion. 4.2-57
However, the holding force created by a single winding is sufficient to overcome the rundown torque produced by the mechanism load.
.( Therefore, the rod cannot move except under the control of the l
, power supply.
The rotational energy is supplied in sequential pulses to the arnature which rotates directionally 15 degrees per pulse as controlled by the power supply. 4.2.3.3 Design Evaluation 4.2.3.3.1 Reactivity Control Components The components are analyzed for loads corresponding to normul, upset, emergency and faulted conditions. The analysis performed depenas on tne mode of operation under consideration. The scope of the analysis requires many different techniques and methods, both static and dynamic. l n. ' (' 1 Some of the loads that are considered on each component where applicable are as follows: I
- 1. Control Rod Scram (equivalent static load)
- 2. Differential Pressure
- 3. Spring Preloads
- 4. Coolant Flow Forces (static)
- 5. Temperature Gradients
- 6. Differences in thennal expansion
- a. Due to temperature differences
- b. Due to expansion of different materials
- 7. Interference between components
- 8. Vibration (mechanically or hydraulically induced)
- 9. All operational transients listed in Table 5.2-2
'J 4.2-58 i
- 10. Pump Overspeed i 11. Seismic Loads (operation basis earthquake'and design basis earthquake) i The main objective of the analysis is to satisfy allowable stress {
limits, to assure an adequate design margin, and to establish deformation limits which are concerned primarily with the functioning of the components. The stress limits are established not only_ to assure l that peak stresses will not reach unacceptable values, but also limit the amplitude of the oscillatory stress component in consideration of fatigue characteristics of the materials. Standard methods of strength of materials are used to establish the stresses and deflections of these components. The dynamic behavior of the reactivity l control components has been studied using experimental test data and experience from operating reactors. l l l The design of reactivity component rods provides a sufficient cold ! void volume within the burnable poison and source rods to limit the internal pressures to a value which satisfies the criteria in Section 4.2.3.1. The void volume for the helium in the burnable poison rods is obtained through the use of glass in tubular form which provides a central void along the length of the rods. Helium j gas is not released by the neutron absorber rod material, thus the absorber rod only sustains an external pressure during operating I conditions. The internal pressure of source rods continues to increase from ambient itil end of life at which time the internal pressure never exceeds that allowed by the criteria in Section l 4.2.3.1. Based on available data for properties of the borcsilicate glass and on nuclear and thermal calculations for the burnable poison rods, gross swelling or cracking of the glass tubing is not expected during operation. Some minor creep of the glass at the hot spot on the inner surface of the tube could occur but would continue only until the glass came in contact with the inner liner. The O 4.2-59 l l
J wall thickness 'of the inner liner is sized to provide adequate
; support in the event of slumping and to collapse locally before rupture of the exterior cladding if unexpect.d large volume changes l
~due to swelling or cracking should occur. The top of the inner liner is open to allow coranunication to the central void by tne helium which diffuses out of the glass, t
. . 'l Sufficient diametral and end clearances have been provided in the neutron absorber, burnable poison, and source rods to accommodate the relative thermal expansions between the enclosed material and the surrounding clad and end plugs. There is no bending or warping induced in the rods although the clearance offered by the guide thimble would permit a postulated warpage to occur without restraint on the rods. Bending, therefore, is not considered in the analysis of the rods. The radial and axial temperature profiles have been determined by considering gap conductance, thermal expansion, and j neutron and/or gamma heating of the contained material as well !
as garena heating of the clad. The maximum neutron absorber material temperature was found to be less than 850 F which occurs axially at only the highest flux region. The maximum borosilicate glass temperature was calculated to be about 1200 F and takes place following the initial rise to power. The glass temperature then decreases rapidly for the following reasons: (1) reduction in power generation due to B 10 depletion; (2) better gap conductance as the helium produced diffuses to the gap; and (3) external gap reduction due l to borosilicate glass creep. Rod, guide thimble, and dashpot flow l analysis performed indicates that the flow is sufficient to prevent coolant boiling and maintain clad temperatures at which the clad material has adequate strength to resist coolant operating pressures and rod internal pressures. Analysis on the full length rod cluster control spider indicate: the spider is structurally adequate to withstand the various operating loads including the higher loads which occur during the drive mechanism O 4.2-60 , ?
., e - . ~ . p - , . - .-
l stepping action and rod drop. Experimental verification of the
)
p spider structural apability is planned (see Section 1.5). Al- l though the partial length rod cluster control is structurally identical i to the full length rod cluster control spider it is not subjected ; l to the loads of rod drop. The roller nut mechanism moves smoothly l with negligible acceleration. Therefore, the spider has a large l l structural margin when used with the partial length rod cluster l drive mechanisms. l l The materials selected are considered to be the best available j from the standpoint of resistance to irradiction damage and compatibility
- to the *;eactor environment. The materials selectea partially dictate l
!. the reactor environment (e.g., Cl~ control in the coolant). The current design type reactivity controls have been in service for as much as six years with no apparent degradation of construction i materials. 1 L l With regard to the materials of construction exhibiting satisfactory ; l A resistance to adverse property changes in a radioactive environment, I it should be noted that work or breeder reactors in current design, similar materials are being applied. At high fluences the austenitic materials increase in strength with a corresponding decreased ductility (as measured by tensile tests) but energy absorption (as measured l by impact tests) remais quite high. Corrosion of the materials exposed to the coolant is quite low and proper control of Cl and 02 in the coolant will prevent the occurrence of stress corrosion. The 17-4 PH parts are all aged at the highest standard aging temperature nf 1100 F to avoid stress corrosion problems exhibited by aging at lower temperatures. l Analysis of the full and partial length rod cluster control assemblies show that if the drive mechanism housing ruptures the rod cluster control assembly will be ejected from the core by the pressure differential of the operating pressure and ambient pressure across . 4 1 4.2-61 1 L _ ,e - ~ -n g - 9
l the drive rod assembly. The ejection is also predicted on the failure j i of the drive mechanism to retain the drive rod / rod cluster control ! assembly pcsition. It should be pointed out that a drive mechanism housing rupture will cause the ejection of only one rod cluster control 1 assembly with the other assemblies remaining in the core. Analysis l also showed that a pressure drop in excess of 4000 psi must occur i across a two-fingered vane to break the vane / spider body joint causing ejection of two neutron absorber rods from the core. Since the greatest pressure nf the primary system coolant is only 2250 psi, i a pressure drop 4 excess of 4000 psi could not be expected to occur. I Thus, the ejection of the neutron absorber rods is not possible. l Ejection of a burnable poison or thimble plug assembly is conceivable based on the postulation that the hold down bar fails and that the base plate and burnable poison rods are severely deformed. In the unlikely event that failure of the hold down bar occurs, the upward displacement of the burnable poison assembly only permits the base plate to contact the upper core plate. Since this displacement is l small, the major portion of the borosilicate glass tubing remains positioned within the core. In the case of the thimble plug assembly, l j the thimble plugs will partially remain in the fuel assembly guide ' thimbles thus maintaining a majority of the desired flow impedance. Further displacement or complete ejection would necessitate the square base plate and burnable poison rods be forced, thus plastica 11y deformed, l to fit up through a smaller diameter hole. It is expected that this l condition requires a substantially higher force or pressure drop than that of the hold down bar failure. , i l Experience with control rods, burnable poison rods, and source rods are dise,ussed in Reference 9. The mechanical design of the reactivity control components provides for the protection of the active elements to prevent the loss of l F
- O -
DECEMBER,1973 4.2-62 f r
control capability and furstional failure of critical components. The components have been ruised for potential failure and conse-quences of a functional failure of critical parts. The results of the review are sunnarized below. Full and Part length Rod Clurter Control Assembly
- 1. The Lub: absorbing material is sealed from contact with the l primary coolant and the fuel assembly and guidance surfaces by a high quality stainless steel clad. Potential lbss of absorber j mass or reduction in reactivity control material due to mechanical I or chemical erosion or wear is tnerefore reliably prevented.
l
- 2. A breach of the cladding for any postulated reason does not I result in serious consequences. The absorber material silver-indium-cadmium is relatively inert and would still remain remote from high coolant velocity regions. Rapid loss of material resulting in significant loss of reactivity control material would not occur.
O 3. The individually clad absorber rods are doubly secured to the retaining spider vane by a threaded joint and a welded lock I pin. No failure of this joint has ever been experienced in functional testing or in years of actual service in operating plants such as San Onofre, Connecticut Yankee, Zorits Beznau No. 1, Robert Emmett Ginna, etc. It should also be noted that in several instances of control rod jamming caused by foreign particles, the individual rods at the site of the jam have borne the full capacity of the control rod drive mechanism and higher impact loads to dislodge l the jam without failure. The conclusion to be drawn from this experience is that this joint is extremely insensitive to potential mechanical damage. A failure of the joint would result in the insertion of the individual rod into the core. This results in reduced reactivity which is a fail safe condition. : O . 4.2-63 i
4 The spider finger braze joint by which the individual rods ,m are fastened to the vanes has also experienced the service d described above and been subjected to the same jam freeing procedures also without failure. A failure of this joint would also result in insertion of the individual rod into the core.
$. The radial vanes are attached to the spider body, again by a brazed joint. The joints are designed to a theoretical strength in excess of that of the conponents joined.
It is a feature of the design that the guidance of the rod cluster control is accomplished by the inner fingers of these vanes. They are therefore the most susceptible to mechanical damage. Since these vanes carry two rods, failure of the vane-to-hub joint such as the isolated incidents at Connecticut-Yankee does not prevent the free insertion of the rod pair.b93 Neither does such a failure interfere with the continuous p free operation of the drive line, also as experienced at Connecticut-() Yankee.b93 Failure of the vane-to-hub joint of a single rod vane could potentially result in failure of the separated vane and rod to insert. This could uccur only at withdrawal elevations where the spider is above the continuous guidance section of the guide tube (in the upper internals). A rotation of the disconnected vane could cause it to hang on one of the guide cards in the intermediate guide tube. Such an occurrence would be evident from the failure of the rod cluster control to insert below a certain elevation but with free motion above this point. This possibility is considered extremely remote because the single rod vanes are subjected to only vertical loads and very light lateral reactions from the rods. The consequences of such a failure are not considered critical since only one G 4.2-64
drive line of the reactivity control system would be involved. l This condition is readily observed and can be cleared at shutdown. j
- 6. The spider hub being of single unit cylindrical construction '
is very rugged and of extremely low potential for damage. It is difficult to postulate any condition to cause failure. Should some unforeseen event cause fracture of the hub above ' the vanes, the 1ower portion with the' vanes and rods attached would insert by gravity into the core causing reactivity decrease. The rod could then not be removed by the drive line, again a fail safe condition. Fracture below the vanes cannot be postulated since all loads, including scram impact, are taken above the vane elevation.
- 7. The rod cluster control rods are provided a clear channel for insertion by the guide thin 61es of the fuel assemblies.
All fuel rod failures are protected against by providing this physical barrier between the fuel rod and the intended insertion channel. Distortion of the fuel rods by bending , cannot apply sufficient force to damage or significantly distort the guide thinble. Fuel rod distortion by swelling, though precluded by design, would be terminated by fracture before contact with the guide thinble occurs. If such were not the case, it would be expected that a force reaction at the point of contact would cause a slight deflection of the guide thicble. The radius of curvature of the deflected shape of the guide thimbles would be sufficiently large to have a negligible influence on rod cluster control insertion. Burnable Poison Assenblies The burnable poison assenblies are static temporary reactivity control elements. The axial position is assured by the hold down assembly which bears against the upper core plate. Their lateral position is maintained by the guide thinbles of the fuel assemblies. 4.2-65
The individual rods are shouldered against the underside of the retainer plate and securely fastened at the top by a threaded nut which is then locked in place by a welded pin. The square dimension of the retainer plate is larger than the diameter of the flow holes through the core plate. Failure of the hold down bar or spring pack therefore does not result in ejection of the burnable poison rods from the core. The only incident that could potentially result in ejection of the burnable poison rods is a multiple fracture of the retainer plate. This is not considered credible because of the light loads borne by this component. During normal operation the loads borne by the plate are approximately 5 lb/ rod or a total of 100 lb. distributed at the points of attachment. Even a multiple fracture i 1 of the retainer plate would result in jaming of the plate segmentsi i against the upper core plate, again preventing ejection. Excessive reactivity increase due to burnable poison ejection is therefore prevented. The same type of stainless steel clad used on red cluster controls
- is also used on the burnable poison rods. In this application i there is even less susceptibility to mechanical damage since these are static assemblies. The guide thimbles of the fuel assembly f afford the same protection from damage due to fuel rod failures i as that described for the rod cluster control rods.
l The consequences of clad breach are also similarly small. The poison material is borosilicate glass which is maintained in position by a central hollow tube. In the event of a hole developing in the clad for any postulated reason the expected consequence is l only the loss of the helium produced by the absorption process 1 ( into the primary coolant. The glass is chemically inert and remains remote from high ccolant velocities, therefore significant loss
- 1. of poison material resulting in reactivity increase is not expected.
l Rods of this design have performed very well in actual service with no failures observed through full life of one fuel cycle. 4 DECEMBER, 1973 4.2-66 J f
l l Drive Rod Assemblies
- {
i All postulated failures of the drive rod assemblies either by fracture or uncoupling lead to the fail safe condition. If the drive rod assembly fractures at any elevation, that portion remaining j coupled falls with, and is guided by the rod cluster control assembly. This always results in reactivity decrease for full length control rods. Such an occurrence on a part length rod could result in reactivity increase but is not concidered an excessive reactivity increase. 4.2.3.3.2 Control Rod Drive Mechanism Material Selection l All pressure-containing materials comply with Section III of the ASME pressure vessel code, and with the exception of the needle vent valve, will be fabricated from austenitic (304) stainless steel or CF-8 stainless steel. The vent valve is a modified austenitic stainless steel cap screw. Magnetic pole pieces ara fabricated from 410 stainless steel. All non-magnetic parts,. except pins and springs, are fabricated from 304 stainless steel. Haynes 25 is used to fabricate link pins. Springs are made from Inconel X. Latch arm tips are l clad with Stellite 6 to provide improved wearability. Hard chrome l plate and Stellite 6 are used selectively for bearing and wear surfaces. l l l At the start of the development program, a survey was made to determine whether a material better than 410 stainless steel was available for the magnetic pole pieces. Ideal material requirements are l as follows: l } 1. High magnetic saturation value
- 2. High permeability 4.2-67
- 3. Low coercive force !
4 High resistivity (j~~] \ l s l
- 5. High curie temperature I
- 6. Corrosion resistant i
- 7. High impact strength l
- 8. Non-oriented )
l
- 9. High machinability
- 10. Radiation damage l After a comprehensive material trade-off study was made it was decided j that the 410 stainless steel was satisfactory for this application. l
\
The cast coil housings require a magnetic material. Both low-(~} carbon cast steel and ductile iron have been successfully tested for this application. The choice, made on the basis of cost, indicates that ductile iron will be specified on the control rod i drive mechanism. The finished housings are zinc plated to provide i corrosion resistance. l l l Coils are wound on bobbins of molded Dow Corning 302 material, with double glass-insulated copper wire. Coils are than vacuum impregnated with silicon varnish. A wrapping of mica sheet is secured to the coil outer surface. The result is a'well-insulated coil capable 1 of sustained operation at 200 degrees centigrade. The drive shaft assembly utilizes a 410 stainless steel drive rod. The coupling is machined from 403 stainless steel. Other parts are i 304 stainless steel with the exception of the springs which are Inconel-X and the locking button which is Haynes 25. j N..) 4.2-68 DECEMBER,1973
Radiation Damage p As required by the equipment specification, the control rod drive mechanisms are designed to meet a radiation requirement of 10 RADS /HR. Materials have been selected to meet this requirement. The above radiation level which amounts to 1.753 x 10 6RADS in twenty years will not limit control rod drive mechanism life. Control rod drive mechanism's at Yankee Rowe which have been in operation since 1960 have not experienced problems due to radiation. Positioning Requirements The mechanism has a step length of 5/8 inches which determines the positioning capabilities of the control rod drive mechanism. (Note: Positioning requirements are determined by reactor physics.) Evaluation of Materials Adequacy The ability of the pressure housing components to perfonn throughout I O the design lifetime as defined in the equipment specification is confirmed by the stress analysis report required by the ASME boiler and pressure vessel code, Section III. Internal components subjected to wear will withstand a minimum of 3,000,000 steps without refurbishment as confirmed by life tests. Results of Dimensional and Tolerance Analysis 1 With respect to the control rod drive mechanism systems as a whole, critical clearances are present in the following areas:
- 1. Latch assembly (Diametral clearances)
- 2. Latch arm-drive rod clearances O
4.2-69 DECEMBER,1973 l
l i l 3. Coil stack assembly-thermal clearances O 4. Coil fit in coil housing The following write-up defines clearances that are designed to provide reliable operation in the control rod drive mechanism in
. these four cri tical' areas. These clearances have been proven by life tests and actual field performance at operating plants.
- 1. Latch Assembly - Thermal Clearances l
l The magnetic jack has several clearances where parts made of 410 stainless steel fit over parts made from 304 stainless steel. Differential thermal expansion is therefore important. Minimum clearances of these parts at 68 F is 0.011 inches. At the maximum design temperature
.of 650 F minimum clearance is 0.0045 inches and at the maximum expected operating temperatures of 550 F is 0.0057 inches. -
- 2. Latch Arm - Drive Rod Clearances The control rod drive mechanism incorporates a load transfer action.
The movable or stationary gripper latch are not under load during engagement, as previously explained, due to load transfer action. Figure 4.2-20 shows latch clearance variation with the drive rod as a result of minimum and maximum temperatures. Figure 4.2-21 shows clearance variations over the design temperature range. i
- 3. Coil Stack Assembly - Thermal Clearances l 1
The assembly clearance of the coil stack assembly over the latch j housing was selected so that the assembly could be removed under all anticipated conditions of thermal expansion. At 70 F the inside diameter of the coil stack is 7.308/7.298 inches. The outside diameter of the latch housing is 7.260/7.270 inches. 4.2-70
l Themal expansion of the mechanism dua to operating temperature of the control rod drive mechanism results in minimum inside 1 h w/ diameter of the coil stack being 7.310 inches at 222*F and the maximum latch housing diameter being 7.302 inches at 532 F. Under the extreme tolerance conditions listed above it is necessary to allow time for a 70*F coil housing to heat during a replacement operation. I Four coil s',ack assembliet were removed from four hot control rod drive mechanism mounted on 11.035 inch centers on a 550 F test loop, allowed to cool, and then replaced without incident as a test to prove the preceding.
- 4. Coil Fit in Coil Housing Control rod drive mechanism and coil housing clearances are selected so that coil heat up results in a close or tight fit. This is l done to faciiitate thermal transfer and coil cooling in a hot control rod drive mechanism.
4.2.3.4 Tests. Verification and Inspections 4.2.3.4.1 Reactivity Control Components Tests and inspections are performed on each reactivity control component to verify the mechanical characteristics. In the case of the full length rod cluster control assembly, prototype testing has been conducted and both manufacturing test / inspections and functional testing at the plant site are performed. During the component manufacturing phase, the following requirements apply to the reactivity control components to assure the proper functioning during reactor operation: (m 4.2-71
1 I I All materials are procured to specifications to attain the 1. desired standard of quality. [] v
'2. A spider from each braze lot is proof tested by applying a 4500 pound load to the spider body, so that approximately 187.5 lbs.
is applied to each of the 24 fingers. This proof load provides j a bending moment at the spider body approximately equivalent to the load caused by the acceleration imposed by the CRDM. All spiders are tested in this manner, even though the spiders used i with the part length rods experience much lower loads.
- 3. All clad /end plug welds are checked for integrity by visual inspection, X-ray, and are helium leak checked. All the seal welds in the neutron absorber rods, burnable poison rods and I source rods are checked in this manner.
- 4. To assure proper fitup with the fuel assembly, the rod cluster control, burnable poison and source assemblies are installed
(] in the fuel assembly without restriction or binding in the kl dry condition with a force not to exceed 25 pounds. Also a straightness of .012/ft. is required on the entire inserted length of each rod assembly. The full length rod cluster control assemblies are functionally tested, following core loading but prior to criticality to demon-strate reliable operation of the assemblies. Each assembly l is operated (and tripped) one time at no flow / cold conditions 1 and one time at full flow / hot conditions. In addition, selected I assemblies, ainounting to about 15 to 20 percent of the total assemblies are operated at no-flow / operating temperature conditions and full flow / ambient conditions. Also the slowest rod and the fastest rod are tripped 10 times at no-flow / ambient conditions l h LJ 4.2-72 DECEMBER, 1973
_.._... _ _ _ _ m. . . . _ . _ _ _ _ _ . . _ _ - . _ _ _ _ . . _ . . . _ . _ _ _ . _ . l l and at full flow / operating temperature conditions. Thus each assembly is tested a minimum of 2 times or up to 14 times maximum to insure that the assemblies are properly functioning. 4.2.3.4.2 Control Rod Drive Mechanisms Quality assurance procedures during production of control rod drive mechanisms include material selection, process control, mechanism component tests during production and hydrotests. After all manufacturing procedures had been developed, several' l prototype control rod drive mechanisms and drive rod assemblies i were life tested with the entire drive line under environmental f conditions of temperature, pmssure and flow. All acceptance tests were of duration equal to or greater than service required for the plant operation. All drive rod assemblies tested in this manner i have shown minimal wear damage. t These tests include verification that the trip time achieved by ; the full length control rod drive mechanisms meet the design requirement of 1.8 seconds from start of rod cluster control assembly motion I to dashpot entry. This trip time mquirement will be confirmed for each control rod drive mechanism prior to initial reactor operation and at periodic intervals after initial reactor operation. In l addition, a Technical Specification has been set to ensure that the trip time requirement is met. It is expected that all control rod drive mechanisms will meet I specified operating requirements for the duration of plant life with normal refurbishment. However, a Technical Specification pertaining to an inoperable n>d cluster control assembly has been set. If a rod cluster control assembly cannot be moved by its mechanism, adjustments in the boron concentration ensure that adequate shutdown l l lO l 1 4.2-73
~ :: -_ -_ -,
margin would b3 achieved following a trip. Thus, inability to ! move one rod cluster control assembly can be tolerated. More than e one inoperable rod cluster control assembly could be tolerated, but would impose additional demands on the plant operator. Therefore , I the number of inoperable rod cluster control assemblies has been i limited to one. In order to demonstrate continuous free movement of the full length l rod cluster control assemblies and to ensure acceptable core power distributions during operation, partial-movement checks are performed on every full length rod cluster control assembly every two weeks during reactor critica' operation. In addition, periodic drop l tests of the full length rod cluster control assemblies are performed at each refueling shutdown to demonstrate continued ability to meet trip time requirements, to ensure core subcriticality after reactor trip, and to limit potential reactivity insertions from a hypothetical rod cluster control assembly ejection. During these tests the acceptable drop time of each assembly is not greater than 1.8 seconds, at full flow and operating temperature, from the beginning of motion to dashpot entry. To confirm the mechanical adequacy of the fuel assembly and full length rod cluster control assembly, functional test programs have been conducted on a full scale control rod. The prototype assembly . was tested under simulated conditions of reactor temperature, pressure, I and flow for approxiniately 1000 hours. The prototype mechanism accumulated about 3,000,000 steps and 600 scrams. At the end of the test the control rod drive mechanism was still operating satisfactorily. A correlation was developed to predict the amplitude of flow excited vibration of individual fuel rods and fuel assemblies. Inspection i of the drive line components did not reveal significant fretting. l The control rod free fall time against 125 percent of nominal flow l was less than 1.5 seconds to the dashpot; about 10 ft. of travel. 4 4.2-74 1
l 1 i Actual experience on tne. Ginna, Mihama No.1 Point Beach No. I and H. B. Robinson plants indicates excellent performance of control ( rod drive mechanisms. There has been excellent perforrance from Westinghouse part length control rod drive mechanisms as demonstrated on operating plants such as Benzau No.1, Point Beach No.1 H. B. Robinson, Ginna, and Mihama No. 1. All units are production tested prior to shipment to confirm ability i 1 of control rod drive mechanisms to meet design specification-operational requirements. Periodic tests are also conducted during plant operation I to confirm brake core operation. During refueling, tests are also conducted to confirm condition of stator windings, l 4.2.3.5 Instrumentation Applications j
/h#
(] Instrumentation for determining reactor coolant average temperature
)
(T,yg) is provided to create demand signals for moving groups of l full length rod cluster control assemblies to provide load follow (determined as a function of turbine impulse pressure) during normal operation and to counteract operational transients. The hot and cold leg resistance temperature detectors (RTD's) are described in Section 7.2 in the reactor coolant bypass loops. The location of the RTD's in each loop is shown on the flow diagrams in Chapter
- 5. The Reactor Control System which controls the reactor coolant average temperature by regulation of control rod bank posicion is described in Section 7.3.
Rod position indication instrumentation is provided to sense the actual position of each control rod - full length as wel? as part length - so that the actual position of the individual rod may be displayed to the operator. Signals are also supplied by this system as input to the rod deviation comparator. The rod position p) indication system is described in Chapter 7. I 4.2-75 l
The reactor makeup control system, whose functions are to permit
/G adjustment of the reactor coolant boron concentration for reactivity U control (as well to maintain the desired operating fluid inventory in the volume control tank), consists of a group of instruments arranged to provide a manually preselected makeup composition that
)
is borated or diluted as required to the charging pump suction header or the volume control tank. This system, as well as other systems including boron sampling provisions that are part of the Chemical and volume Control System, are described in Section 9.3. When the reactor is critical, the normal indication of reactivity status in the core is the position of the control bank in relation to reactor power (as indicated by the Reactor Coolant System loop AT) and coolant average temperature. These parameters are used i to calculate insertion limits for the control banks to give warning to the operator of excessive rod insertion. Monitoring of the neutron flux for various phases of reactor power operation as well as of core loading, shutdown, startup, and refueling is by means - j ('~] of the Nuclear Instrumentation System. The monitoring functions
'V and readout and indication characteristics for the following means !
of monitoring reactivity are included in the discussion on safety related display instrumentation in Section 7.5:
- 1. Nuclear Instrumentation System
- 2. Temperature Indicators
- a. T average (Measured)
- b. AT (Measured)
- c. Auctioneered T average l
- d. T reference
- 3. Demand Position of Rod Clusten Control Assembly Group
- 4. Actual Rod Position Indicator.
(3
)
4.2-76 i 1
4.
2.4 REFERENCES
O 1. J. A. Christensen, R. J. Allio and A. Biancheria, " Melting Point of Irradiated UO ." 2 WCAP-6065, February,1965.
- 2. J. Cohen, " Development and Properties of Silver Base Alloys as Control Rod Materials for Pressurized Water Reactors." WAPD-214, December, 1959.
3.- Supplemental information on fuel design transmitted from R. Salvatori, Westinghouse NES, to D. Knuth, AEC, as attachments to letters NS-SL-518 (12/22/73), NS-SL-521 (12/29/72), NS-SL-524 (12/29/72) and NS-SL-543 (1/12/73), (Westinghouse Proprietary); and supplemental information on fuel design transmitted from R. Salvatori, Westinghouse NES, to D. Knuth, AEC, as attachments to letters NS-SL-527 (1/2/73) and NS-SL-544 (1/12/73).
- 4. J. Weisman, P. E. MacDonald, A. J. Miller and H. M. Ferrari, q " Fission Gas Relesse from 00 2 Fuel Rods with Time Varying L/ Power Histories," Trans. Am. Nucl . Soc. , (12), (1969) pp 900-901.
- 5. C. M. Friedrich and W. H. Guilinger, "CYGRO-2, A Fortran IV Computer Program for Stress Analysis of the Growth of Cylin-drical Fuel Elemer.ts with Fission Gas Bubbles," WAPD-TM-547, November , .1966.
- 6. C. J. Kubit (Ed.), " Safety Related Research and Development for Westinghouse Pressurized Water Reactor - Program Summaries, Spring-Fall 1973 ." WCAP-8204, October,1973.
- 7. W. J. Dollard, " Nuclear Fuel Division Quality Assurance Program Plan ." WCAP-7800, Revision 2, September ,1973.
- 8. S. Kraus, " Neutron Shielding Pads," WCAP-7870, May,1972.
O - 4.2-77 l-
- 9. J. Skaitka, " Operational Experience - Westinghouse Cores,"
WCAP-8183, October,1973. C\
'7' 10. J. M. Hellman (Ed.), " Fuel Densification Experimental Results and Model For Reactor Operation," WCAP-8218, October, 1973 (llestinghouse Proprietary); and WCAP 8219, October, 1973.
e v 1 l I l e DECEMBER,1973
'V 4.2-78 i
t
TABLE 4.2-1 l MAXIMUM DEFLECTIONS ALLOWED FOR REACTOR INTERNAL SUPPORT STRUCTURES-l 3 Loop Plant 1 i No-Loss-of Allowable Function ; Component Deflections (ini . Deflections (in) ! 1 Upper 8arrel j radial inward 4.38 8.77 l radial outward 0.5 1.0 Upper Package 0.10 0.15 Rod Cluster Guide Tubes 1.00 1.75 4 Loop Plant
' l No-Loss-of Allowable Fun, tion Component- Deflections (in) Def_ lections (in)
Upper Barrel ! radial inward 4.1 8.2 ; radial outward 0.5 1.0- ! Upper Package 0.10 0.15 Rod Cluster Guide Tubes 1.00 1.75 4.2-79 or -
r 4 8.426 REF > 4-16 SPACES AT 0.496 = 7.93 6 -> p 33 y 0.245-> 4-
+ 0.122
- 0. 530 -> 4 y
}
8 88 8 8 0.05. + g.- O-> oo 496 O Og bO g,oeo,8 O g O n.273 ru n ASS v AND CONTROL R0D Oo > o O goooooO riTcH g* 1.488 4- 0.068 0 g g oos 4- 2 976 -> o og geo e og 2 0 O O o gooooog 6Uiot /gooooog ooo 88 'o o e' 8 00000000000000000 0000000000000000b /t 8.466 00000000000000000 0000000000000000g ( .040 TYP O g '9 O O ooo O WP 8% '#8 8 8 g% O ' pg gooo - og O O O g o= --o g goo < o og y O 8
]
y 'g 8 8o o o o o % Nmumm1. h O o oO SHEATH gh N FUEL ASSEMBLY 000 00000000000Q OOOOOOOOOOOOOCQ2Q WITHOUT R0D CONTROL CLUSTER CLUSTER CONTROL ELEMENT
- FUEL RODS 264 REQ'D 00 = 0.3M CLAD THICKNESS = 0.0225 CLAD WTERIAL - ZlRC-4 Figure 4.2-1. Fuel Assembly Cross Section 17 x 17 4.2-80 l
1 w %, (
\
Top Nozzle 0 g .875 DIA REF ( :-ver 4 y=b ^
^
--8 @ t L$(/ ,'/)7
< / i r.
( , 7.1 62 REF
.* . J w --v 1 8.404 SQ C -, + -
REF
\.
v __ . -v.s,
$d' $${, y TOP YlEW o.551 DI A REF Pad i CONTROL R00 1
151.
'g _ R00 CLUSTER CONTROL ASSEMBLY e~" -
n [ FUEL R0D T r --
-v f[ m i == t
{GRIDASSEMBLY (( f=== rm= h Y cL_.f === r ll ,1 === f ll j == c1 = - - -#+4:A L i i 3 === L ii _g == w _' y Mk ' (D === r' l j._.,j === r- j w' === [ f, C7 == k !! . w === t. !! 3. == j = = _r c ] === r- j ===
+ !! ]f
' \
(, == L 1 I l I _T_ === L._ _%. ===
\ \ T~ r- _
>\
'{
Q $ 1Q = ?k f === L l!
/ /
i == ? M. === t . J == M, === Triple n-8 * ( ) === r I I J === r I I J == 1 g c_2 == t i i -w === t ii -s == Leaf E ._ , c > == r i! i==r' === Spring ll .. ( __ 3 , == b____./ ! 3 == L i! M. = = {x ( === r !! ,1 == f ,1 ==
- _h =-I 1 I T == f % ==
Clam _ __/ Enc 'l os e
// /
Top Plate 4- 154. 0 R E F --*->- < 135.4 REF h < g < 159.? REF - Adapter Plate 4. t~~ . . . _ . . ..
#~
0.875 DI A REF . < (( Ak . ~, P
~
00000 JL ((0000 O [,~ Aki iau. , EN U"
*C8N8E8 00000000
' - Bottom Nozzle 8.u24 sq a @oQwo REF w , ,
6.750 REF V _ . _ ._ Bearing Plate BOTTOM VIEW Angle Leg Perforated Plat-53 REF - > ll l ll ll u
!! !==t. ! ./ ! == L. ! L__ / == t. !i ! == p , . . . . .wh
! l J === r ! W ==r !! 2 == f !! 1 =t l L ;--
!! T. = = t. !! T. = t. !! T. == L, i! T == t to I e p a
! l ,J = = r !! 1 == r !! _ s ==: r l! j = = r*~ ~ D I l
!! T. == L. I I T :== L 1 I T === L iI _T == - l
!! _ J == r i i j==r' ii j = = r' ii j ==
!! T. == L // ' T. == L. ll ' T, == L, ii ~T == -D Plenum
!I J == f I I J == r i
) =, -ij j ==
~ ~ -, -L _=..-.=_f r . _
b% j == r i j == 1 === ? T. === L _ J == r
!l i I ii
_'% === L j == r ' ll I I ii j = = r' _T === t j == r^ I ii li 'T == j == r p 31
!! T. == L. !! ~T. == L, !! ~T. == L. ll T. == t 3 i
_I _J === r I I J === r I I J ===r ;i j == ') l t n
+ j[.
! T. == t. li 's == 't ii ~T == t ii w == j F-
! . ' == ? !! j == r' ll ,J == r' ll j == 3 I T ===1 I I .% == I i1 3 ===I iI .g === '
L.r' I ll ll ll 2.7 REF < > 39.2 REF h < 83.0 REF h 4 56.8 REF M < 30.6 REF w 4--- - 6. 2 R E F y c:9o 12To676 -C3 ( Figure 4.2-2. Fuel Assembly Outline 17x17 2-81 AMENDMENT 1
i h / g ! g END PLUG f 9 N s i N s i N s i ' u A g PLENUM A I S PR ING l
- ; c:::P F s s 6.3" s s N N (TYP) '
s t g;;;; s x h d s<
) $M02 PE LLETS
- k X$s ,
1 N N l
% I N N N N -
s N 151.6" N N FUEL-CLAD GAP (REF) 144.0"
'A (TYP) k i N
, 4 N ZlRC ALOY CLAD
< s
% N N N N N N
N
' N y
N N N N N N N h b
) (
y % 1 SPECIFIC DIMENSIONS DEPEND ON DESIGN VARI ABLES SUCH AS PRE-PRESSURlZATION. POWER HISTORY. AND DISCHARGE BURNUP Figure 4.2-3. Fuel Rod Schematic l l ( 4.2-82 AMENDMENT 1 1
d% e
\
l (~ ALIGNING Pih UPPER CORE PLATE GUIDE PIN l I i CORE BARREL BAFFEL ( x11 1 111111111111111111,,,, ,,, , ,,,_ , Ld
,_ ,J MM L"
K1 Y
,"} ' "
m u
\
u --- l
~- _
w l l o \ o o o o o I 2 0
% o o o o o o !
g . . . . o o O ,
^ ^
v N0ZZLE O O NEUTRON SHIELD Pl a g, - - - - - - r - . in
- - = = g_ _= _
gy - - . ( UPPER CORE BARREL m
., LOWER @
t L FLANGE i i { Figure 4.2-6A. Three Loop Lower Ce
\> p*
4.2-35 (
CORE SUPPORT COL M /4WlG#lE RADIAL SUPPORT KEY
^'#'O ((] .
;1~ [:~;. n% cn i LOWER CORE PLATE CORE SUPPORT ' - *
~
i asams - m _mmmmama CORE SUPPORT COLM ;
,/ LOCKING NUT i
/ h
/ '^
E/ IP i V///AL ii i r /
/ .
- < BUTT TYPE COL M
~w i ;is i
. I y .x li a i l ',',','s t ENERGY ABSORBER ASS'Y
! i s , o\
, 7, - - 7 7 p o e o o o . ,l _,
a fSECONDARY CORE SUPPORT > a e o o o o o l s !
- - - r x
n :.
. 1
=i 7< INSTRUMENTATION GUIDE COL M (CRUCIFORM) l l \ -t
=1 1 LOWER TIE PLATE !
! I -9 ; "EL j l
- ,a 7 UPPER TIE PLATE -
=_=__- = = = ,
, [
t ! IRRADIATION SPECIMEN BASKET ! l )RE BARREL I , l l D Support AssernWy (Core Barrel Assembly)
?
. ~3{s .
i i a f i i
.,e. c : ,a. .m 1, r
..i...-
.., . .m , s ._.-., ., s 7-y,Lm"
'n.,,, g,1-. --...g......,,.<.....,..,I W. </
i .JI E e p_ 3 n .9 fj& f y, yt - e aa
- a' !
I Mkj
. --=.
F .. t V T t
- s. L .l - .lT -_.;-
. k- ~w .t sgh r g;g, !
{c ~
~
.uim, j e
i s -
-+
- C- _
o, , c.. . , & ', 1 . :.yw(.E ; N
~
ja ~ E I N _. ._ na , _ hy - X - x
\
1' c 1 ;a';a' m.se r
; =
g o .. i q
> %m . : ;
wmrcet \ .- p.. a. a "" *
- ,- ecs nn suem e f ..s,.m:.,,
d [ #1TC gj , m=' s
-y_ s ..m,
_-Q N ' ' " " ' ' ' ' " ' ' %=. .. .a . y _
~
w - e=~ m - - . f '\ ,n. , Ctet hNT [ b#Eufa3s lettLB Pasit b estaciaitee $*ts+4e 945GT
% p-- ,
p-- ima :m .a in .-r j
, ait 7
l f FIGURE 4.2-6B FOUR LOOP LOWER CORE SUPPORT ASSEMBLY (CORE BARREL ASSEMBLY) , k
SkeeII8 em&S (9 4me TuteN0 COUPLE COLuh4R$ ._ s 1 1
,, - 4.2-3
[ f ; I , l o APERTURE : CARD I - u m.v m ioon f ~;_ (. " t"} vress cost etattg Vn 2 oppts
.q.
( -> l 88' Post
-- g
* . g ll
, J f-w
~
I CC " 1 11
~
l , _ _ , l, mmmm ,e 3 ( m m
~
; c. <
L l r
- h mm. mO
' S m /
a c. mi n a o 3 W W l 1K , _ T X 11 II 11 11 ll ll 11 3 IW 0 W *N 11 11 Il ll 11 il 11
-- al -
E C""*"3 m m O g om -
, i t l 4 - -
'R '
l H - Q- ~'C l (J [ g __ s '
-I -
I Qw 0 L g iK
- H E
~ 11 11 ll ll ll ll 11 ~~
1M W W 11 11 II 11 Il 11 ll ia . a" a ': u 7 j 3 _ t i mi 4m i cm
\
me
\
3 ( '
)'
e _- i a c. o,
,o u
L B
~
1 ;; - - g " ' ir s ; a c_ - - - t
. art -
3 L
- m 13 c m.t. =,,wr etart l
l FIG. 4.2 - 7A Three Loop UPPER CORE SUPPORT ASSEMBLY Ho so o% 03
=
tr;l!I !!!> - l ii! .i !; !. t! I i! ;I:i b ,rI :;!l i !!L
- eI-Y O
TL S R B E R OM T E P E A P PS L P US P USA E h R O C R E P P U
,I
'l
;l
,)
\
e
,I r f u t
i! c
,)
u r t S t r o _ p p u S e r o
,I C
r . e p - \ p .O U
-T p
o L o r u o F _ 8 7
,I -
;i ' 2
_ 4
!I
\ -
,) e r
_ u g i
!! F
- 1) b n
M _ U _ L O C _ T l R O P P U - S
- O
- Nb
O OUPPORT COLUMNS - GUIDE TUBE O ASSEMBLY LOCATIONS
====r~ g
,, f sN s o
/ S' i N o J -4' %
// ,,s 2, ,
4/ +t; ts , 1xx ,
\\ O 1
//
// s' o I'<,* ) .o ,fe l
/ ,
/(^):o '\
o .o 7 sg o // 'ss k\
/f oi b T~ > o_o
_ [i j- Il O ll
\ kq ' o'n ll o O a tt* # ( *4*
++
a 5
% //
/
\N 1 g\ 9 /ss //
s,b 'Rs // O
+ a o o % // \'
/,h
-::=i=='- O O
DEEP BEAM h WELDMENTS -
- HEAD AND VESSEL ALIGNMENT PIN LOCATION I
(3 \) Figure 4.2-8A. Three Loop Plan View of Upper Core Support Structure 4.2-89 I I
6164-2 0 L GNH TPN LOC AT ION SUPPORT COLUMNS
,,, \ L ATION s <. %
O
/ % .
o o Oo g i N O o - I o n r% y O s I O
% /
O Figure 4.2-88. Four Loop-Plan View of Upper Core Support Structure 4.2-90
mp s ~.sdP s I I I (~ 1 i l RCC AS$fMSLY RCC ASSERY
' ' M" INSERTED
_44 , g
.2 % . . . ..__.... ..
m y, _ <p , . . . . s -
. N % k s .......
w .
,p % .... ...
.y .
,+,g N\
N N \ k1 \ cRom onivt q ASSE%U
- i i
l f' -
.y
(,e l i 4.2 i
.- w l
APERTURE CARD Alco Avctbbb on 7 7 A;ur; Crrd lllll/, y
- ,,~;;-:~~~-,,,,,,,,,,,,,,___
/ -
/ '
. 4M :[
r_ ' ' _% .
.4 _
m
'% \ \ < \ ..__\
i E MEE, *'N w ass =, w .su., 9 9 0 arroc 7c,. FIGURE 4.2-9 FULL LENGTH R00 CLUSTER CONTROL AND DRIVE ROD ASSEMBLY WITH INTERFACING C0t1PONENTS I
U
\
9 ( 4 4 > > ' 4 > p~ semina
.e N-s ,
, ;- 3
- i. a o a wwwg
) 3 -
g
' ' Nori. .
~~ '
u W N r/A /// . l 4 k 4 4lF F
$PIDER BODY
( 4.2< l
- _m,
.~
5 NJ tpe0L2-g er, r$,. Ab; ta LF !: i ri Ep;t^S Y "'Id ; 161.0 150.574
=
i l l O, , l - ll - - i o n < - na -
-Of -
- a; - --
A7 - C =i - --4 ----: ! - -
; y -
; 3 SPRING RETAINER
_ 7 C C -
!! I 20 ', - 1 n _ sm
.88 TRAVEL y' coq - ' - 'l - -
Ay -
;3 C. ; I 1l I
ll, , - - Ay -
;3 O ;
- -H-M ::: . ', -
?I p. 3; D a o
. 361 O I A.- . 381 o t A .
_ 142.00 _ ABSORSER LINCIM AB1DRBER 80'- S I LVER. 15' IND IUN. 5' CADNIUM f ~ Figure 4.2-10 Full Length Rod Cluster l Control Assembly Outline i 1 12 1 J
.g4- - - -a - ,& ~ 4 e o am- a - -,.. _ e a _ s. , a,.a ,44n,,,
l
.2 W_
w -, e N R. 8 G tt 1 3N% 2- s l i I 8 - U i 3 I, ry .g M i (,) e = ! 3 l l E ' l til i T i N n$ 9 i $
.?
u. ISEW WIAE I l { l 4 Pir l N
. (a .
4.2-93 t i
- w- _ , , - - - , - - - - - - _ - , , - - - - - - - , - - , - _ - - - - - - - - - - - - - - - - - - - - - - - , - - - - , , - - - - - - - - - , - - - - . --- - -
i 1 F F 1 F p h e -. j,-\
~ .
; e -
==
3 4 4lF F t SPIDER BODY ( L 4.2-94 w ~a
~
f r i f
.t i
APEKr0RE ! CARD l AgaAvaliLDIO]" f Aperture CnN } l I 161.9 Ii 150.574 )
= I
?
O ' I
+l - - 1, F 'C i O' -
AY - llJ ll l = - i +' - - ;y -
,;3 r SPRING RETAINER 7
t e- -
- ia5 ou ; - ,m_
.88 TRAVEL { ,
k ' _ - , il < _ _io nL i - rw n ty CI J l l -
'o1
-4y ; O_
v i t i m V - ;: - . .t y . - :g; O , n n
.363 otA- .381 dia.
36.00 , ABSORBER LENGTH ABSORSER 80- SILVER. 15 IN0luu 5 CADluu 90 6 ' BOD Q-Figure 4.2-12 Part Length Rod Cluster Control Assembly Outline I ( 1
t
\
.c M
',t r _
-b-j s
-C _-i - ~
E :;_ ( r -- ,e ~ ~
~
Ed__ 4 - l .
-- - -- L[L 9,D F' k/,, ;
CfJIJJuu lt, 1 t J q -- IN I# _d: 7 - t ' 2 - N Ed r, s, m s: -C _ J
\
BORDSILICAIE 2 TsLAS$ TUBE
~
6
, , 0lSON
. _ . <A00 (s m, r
4.2
= APERTURE 142.0 Ref.
~
(;!ARD PolSON LENGIN 2
~
ae:.; Avel!ao!o on J1 [] .iporture Card
!l --
;3 1
II - l3 - i l I i STAINLESS STL 1 i n l
+ '6t!'
L igLE 990 Mca7G-O Figure 4.2-13 Burnable Poison Assembly (Conceptual)
- 95
- - . - - _ ]
A .m -4 -- ,e - - - . 4 s.A, .a +-_a-- ax- - -, - A > e. ..e am --_ -_, 1 l G 1 o 2: 5 (o3 e n n O h
- =
n z I e-LeJ E j r N "i e O o= *~ N -. m o g t P o r eo- g l mA - i
\,,,e o
- Y l l " 5
)L E m l a N
n w m h ch sh D
== i M @
_ i
& ..O u
- rs E a s O '
' $ T as -
=
- o E
- E 1
- 3 1 l
.CB l \ d
- l / .
Y i < g
-j b i x W j S v n
v o m N w e m 17 lI fk ! t 4.2-96
- . - > +m -,.
1 ( l i E
-\
MM g
-sC :I --
3
> w i= - - M
[I .M a m _e.__ 1 ty l I c
= = =
- _ _ )
vm
.y I
= , i
[ .'dJ dIJIJ h l l d i
-4 :1-(
- e
[ 2 J STAINLES$ f, STEEL l ( o 3s: DIA. 0.381 0.436 OlA DIA e o PRlWARY SCOURCE SECONDARY IMlWBLE SOURCE PLUG l l _a
i
~ "a '
= #ERTURE OARD
=
142.0 REF. r0 : 50= LE""
= mo Avallableon Apertura Card l [ M- I l II - .
_1l l 3 e : : ; ) n 1.5 NOM. 32.5 NOM. 4 4._.D CALIFORNIUM SPACER ~~ l - 0<XXXI OO^Od / -
!l ] !
% i w _ m !
u :' 1 b m
! (L I Tl l]
I: 88.0 SEC0x0ARr source
- 2 ANilWONY BERVLLIUM NOTE: ALL DINENSIONS ARE IN INCHES 4
a 0.381 N Y DIA U BURIABLE
\ BORDSILICATE Pols D3 CLA$$ TUBE 990 Figure 4.2-15A.
nooWofThree Loop Source Assembly (Conceptual) 4.2-97
~ 150c
~\
=
- V~ ~D lV\ } [
J r . ^^ __Q E L. 1 I _ - ~_ (a--
~ j
- r _
, , , . , m ,
?t V// '
II
, ,w n L4 -
F-mI g_ q _{ _ J a _ _ _ f a V' ,
'h-'.;J
- U J IJ <J f ; -
p SPRINGS 1 _
} -
J NOTE: 1 01 02
\
f ' '
#: }
0.381 , 0 38' Dja. "g ( PRIMARY SCOURCE x SURNABLE POISDN BOR0 $1LICAIE GLASS TUBE (U 4.2-98
61614-12 l 1 l oi 4PERTURE
*1 r ts n l
1 1.5 NOM. 32.5 NOM
= = =
CALIFORNIUM SPACER /= 5 4 M Luitio o n j
<qmnuto Cerd 1
" ^'IDCOO^
- ~! -: d ;
21 - l l^ j u . l~- d" - l j' 142.0 REF poison trNctu ; nl^
?? -
n l? - - l , bl ~~ i IJ I l IJ
.J ,D - i n lL DIMENSIONS ARE $N INCHES 04 D JLT8076 - :
Figure 4.2-15B. Four Loop-Primary Source Assembly (Conceptual)
, w w-l I l l
=
\
l
=
1 i I l i 4: nx : -+E l
~ . -T ~ C
-5 Q
[
\. , . . , , ,[ M -
i
~
~ ~ ~ ~
! ft ///W lt ll
,...m Ld
.F. "I r, mi,m.
- M H ~ -
,P lt I
_ _ _ t7 i
- = = = = 1/ / / r
~ '
j ~p h55
, ~J .J IJ 1J U
~
, m W
i -
. A
-aC O --
l ( if _ l N1 I ; 2 STAINLESS 3 STEEL g h h 0.381 0.436 0.381
,,, o,, +
e 7 o i,
~
,,coNo m wi. a SouRCI PLUG 'emt l
t a
~w 4.2-99 I
e- -
6164-11
~ l 50. 4
=
142.0 REF P0110N LENGTH
;) - =;J @ERTURE og3D 3 " g g gyCOltDIO O
- t,perturo Ccd 2
vvvvvvvvvvvvvv E nn !!) 88.0 SECOND ARY SOURCE _ ANilMONY BERYLLIUM 1 ALL DICENSIONS ARE IN INCHES 40 SILICATE Q$5 TUBE Four Loop - Figure 4.2-15C. Secondary Source Assembly (Conceptual) C:C'D 2TOO7 -!
w uw k
=
l
\
_K4I 1r ( 1 _ 1 f~ t g f
~ ~T ~ C
- 5 U
( l f(: a____a , +-h [I ." F' Fl
+
s "st M r
; w -
o In I
- - - - o i 1/// ,
I
$ _'JJ J Jhh5 "~
i
, e f _A u ,
I i
-eC _ O ; --
- i. -
1 ll l N{ 1 i 2 STAINLESS ' 3 STEEL l l n it n 0.381 0.436 0.381 { DIA / D!A DIA g p ' l U I SECONDARY THINBLE BURNABLE
$00RCE PLUG POISON 1
i f
't ..
%d 4.2-99,
6164-11
~ 150. 4 142.0 REF PolSON LENOTH
< ,s ~
, , , n.PERTUREmn
( 1 I "i I > {}[,jgp 3 gg, #8$U p cCrtuiO b" "f 2 [ "vD M)WO9!) 88.0 SECONDARY SOURCE __ ANilMONY BERVLLIUM
)
i 4 TE' ALL DIMEN510NS ARE IN INCHES BORDSILIEATE GLA15 10BI Four Loop - Figure 4.2-15C. Secondary Source Assembly (Conceptual) oc'D 2?oo7 - l
r i
\
l 1
)
i _0.434 on
. g d C
~
t ] ( , im i _s a is s,.-.. - , im T; uuuu1Mpt4 '.4 is n ,. ,i J u L m , - ,
~ _ m_
R n f
- - - - o
- - - - 1/// p '
4_, ._ia im p - J . ..s 55E h , - u-e - rw l s ; a is 1 1M
%_7
'l
_, _5 a is I a
' Figure 4.2-16 Thimble Plug Assembly (Conceptual)
O 4.2-100
l l ~ Q (r, 1.0 Q[\ 1 l ROD TRAVEL HOUSING - cY:) H
'h a fr CABLE CONNECTION
% 4*
t, c
,. s (o
Yh i; l t i f
/ " }' f .
' l l
ji f , ! ' (' ke QM f
- r_ ' f s 7,,.
- hI 5 g
/
l kA I L > l s i OPERATING Coll l p ,s; -
?
STACK ASSEMBLY q; pq^ 8N 1 , Y 0
..yfn
,..+ a
/
u .j N PRESSURE HOUSING "lY v1 j * , DRIVE R0D ASSEMBLY oiSCounECT R0 4 .g,- slg&nl1 }
~ e.w A /.
I y' h l s r r l t Figure 4.2-17 Full Length Control Rod Dr,1ve Mechanism 4.2-101
k
' i _
Ip '= l]d DRfvt ROD ASST G
*~
p L l l ,
/,
h: / , r , ROD TRAVEL MOUSileG i p tATCM MouSileG
-{M/ ' #
, 'O RING l,
s p.
, 'i -FLUX RileG i ~E ','A -LOCK wire 4
l'-
, ~ \-Colt STACK ASS'y N j j4 's-LIFT POLE V', /.
JMIFT COIL (C) p MyS *-
~
-LtFT RETURN SPRING l> -l
/
,pFLUM RsNG ytoCu wire WASLE GRIPfYR POLE 4 .
t~ . . .
.. C.0L , . ,
p '
; mova8Lt eatpPte spelm6 f' ~ LOCK PLumets as0watLa LATCH 1 4. ATC M Pim
, . ". [
<pL&TCn ames a
6 #pLemsATCM PmLeasst I 4""* j Ei 2/ M M n g iOCs aos
- Supp0RT Tung INE L ATCN
^
- jg q' Ill STAT 80maev gespeta pcLE
- m. .
--4W40ETUSE W -.m..
l Ik 4 - NATS yemeretaesswee r r O f,j m. ,nu E. n, LWR Alte LOCs Witt L _, g 1'g,'
-m LerY sitTpaa spasme j g' _{ Spam8 9ttaantR hh ,
L ATC u psu fp-pM # LaTCn ames 6 t
' '~
-LOCE PLumsta STAfemacy LaTCes Q Linn 9sm
$-:9w LATCMLams um. .
Nsueront tues stafi0anov LaTCn
- L ATCw STOP h,
S//
, e
/7 v gj Figure 4.2-18 Full Length Controi Rod Drive fiechanism Schematic 4.2-102
1
%** <=s l
l I
.)
CABLE CONNECTION (' . i
.f a L Il I 1.1 f y- .yg, m
T 7 ..
.---m N """
_- _m -M , u
.{g>E ..
tuxxusuuu n.Krqydv _L mm_ _-. rwxsxwxu: xm w:nxxnxwuxwxwx=ec we s cq { -
/
CONOSE AL JOINT s POSITION DE TECTING ColLS
, \ [
'xOtSCONNECT ROD (LvDT) RE TAINING RING MECHANIC AL BR AKE
!$o'uo
< % SECTION 0-B
- 0. A S._ S SECT 80N A. A k
,,..+
9;I 4.2-10
~~
APERTURE C.4RO Altic AValiablO C11 Aperturo Card STATOR MECHANICAL BRAKE PRESSURE HOUSING HE AD ADAPTER % 1 l W^ 1 [ e s l t ff \ k?Ik !
"MMM\\%\%W L
"'*WEMMG'fOY~ i @
~
m m, ,mxmmumummmmuummhumm, mmh. _ n nm- -
----- M
- q. ; . .
t -
.w.__. s W C SPLIT ARM ATURE ROLLER NUT BR AKE ARM PIVOT LEAD SCRED ASS'Y LER NUT CLUSTER j DISCONNECT ROD COUPLING LOCKING
[ BUTTON ab$ ad x MFlph?-
^^ ^ '"'"""]' " " ST&""" ' '"d '.
1
-~'
r
~
1 N - 3
'h h..kU_U_UEM L y;, --
. _ -M l l SEC780N C-C COUPLING DISCONNECT ROD NUT Figure 4.2-19 Part Length Control Rod Drive Mechanism
3 Discussion Two independent reactivity control systems are provided, namely control rods and soluble boron in the coolant. The control rod system can compensate for the reactivity effects of the fuel and water temperature changes accompanying power level changes over the range from full-load to no-load. In addition, the control rod system provides the minimum shutdown margin under Condition I events and is capable of making the core subcritical rapidly enough to prevent exceeding acceptable fuel damage limits assuming that the highest worth control rod is stuck out upon trip. The boron system can co;npensate for all xenon burnout reactivity changes and will maintain the reactor in the cold shutdown. Thus, backup and emergency shutdown provisions are provided by a mechanical and a chemical shim control system which satisfies GDC 26. Basis When fuel assenblies are in the pressure vessel and the vessel lid is not in place, k eff will be maintained at or below 0.90 with control rods and. soluble boron. Further, the fuel will be maintained sufficiently subcritical that removal of all Rod Cluster Control Assemblies will not result in criticality. Basis Fuel will be stored in storage facilities which either (a) prevent significant moderation of neutrons by excluding all sources of water or (b) provide mechanical isolation of individual fuel assemblies such that k is less than 0.90 for the fully flooded condition assuming eff cold clean water and no control rods are present, and such that the array has k g less than 0.99 for any partially moderated condition. O 4.3-7 l l
l- 4. 3.1. 6 Stability l Basis The core will be inherently stable to power oscillations of the fundamental mode. This satisfies GDC 12. 1 Discussion Oscillations of the total power output of the core, from whatever
- cause, are readily detected by the loop temperature sensors and by the nuclear instrumentation. The core is protected by these systems and a reactor trip would occur if power increased unacceptably, preserving the design margins to fuel design limits. The stability I of the turbine / steam generator / core systems and the reactor control system is such that total core power oscillations are not nonna11y possible. The redundancy of the protection circuits ensures an extremely l low probability of exceeding design power levels.
Basis Spatial power oscillations within the core, with a constant core power output, should they occur can be reliably and readily detected i and suppressed. Discussion The core is designed so that diametral and azimuthal oscillations due to spatial xenon effects are self-damping and no operator action or control action is required to suppress them. The stability to diametral oscillations is so great that this excitation is highly l
't 4.3-8 DECEMBER,1973
l improbable. Convergent azimuthal oscillations can be excited by prohibited motion of individual control rods. Such oscillations are'readily observable and alarmed, using the excore long ion chambers. Indications are also continuously available from incore thermocouples and loop temperature measurements. Moveable incore detectors can be activated to provide more detailed in',ormation. In all presently proposed cores these horizontal plane oscillations are self-damping by virture of reactivity feedback effects designed into the core. However, axial xenon spatial power oscillations may occur late in core life. The control bank, the part-length control rods and ex-core detectors are provided for control and monitoring of axial power distributions. Assurance that fuel design limits are not exceeded is provided by reactor overpower AT and overtem-perature AT trip functions which use the measured axial power imbalance as an input. 4.3.1.7 Anticipated Transients Without Trip O G The effects of anticipated transients with failure to trip are not considered in the Design Bases of the plant. Analysis has shown that the likelihood of such a hypothetical event is negligibly small. U Furthermore, analysis of the consequences of a hypothetical failure to trip following anticipated transients will be performed to show that no significant core damage would result and system peak pressures would be limited to acceptable values and no failure of the Reactor Coolant System would result. These analyses will 1 I be documented by October 1974 in accordance with the AEC policy outlined in WASH 1270 " Technical Report on Articipated Transients Without Scram for Water-Cooled Power Reactors," September,1973. 4.3-9 DECEMBER, 1973
4.
3.2 DESCRIPTION
\ .
4.3.2.1 Nuclear Design Description The reactor core consists of a specified number of fuel rods which are held in bundles by spacer grids and top and bottom fittings. The fuel rods are constructed of Zircaloy cylindrical tubes containing 00 2 fuel pellets. The bundles, known as fuel assemblies, are arranged in a pattern which approximates a right circular cylinder. Each fuel assembly contains a 17 x 17 rod array composed of 264 fuel rods, 24 rod cluster control (RCC) thimbles and an in-core instrumentation thimble. Figure 4.2-1 shows a cross sectional view of a 17 x 17 fuel assembly and the related RCC locations. Further details of the fuel assembly are given in Section 4.2.1. . The fuel rods within a given assembly have the same uranium enrichment in both the radial and axial planes. Fuel assemblies of three { \ different enrichments are used in the initial core loading to establish a favorable radial power distribution. Figure 4.3-1 shows the fuel loading pattern to be used in the first core. Two regions consisting of the two lower enrichments are interspersed so as to form a checkerboard pattern in the central portion of the core. The third region is arranged around the periphery of the core and contains the highest enrichment. The reference reloading pattern is placement of new fuel on the core periphery, with depleted fuel moved inward. The core will normally operate approximately one year between refueling, accumulating approximately 11,000 MWD /MTU per year. The preliminary enrichments for the first core are shown in Table 4.3-1. The core average enrichment is determined by the amount of f!ssionable material required to provide the desired core lifetime and energy requirements, namely a region average discharge burnup of 33,000 MWD /MTU. The physics of the burnout process is such that operation 4.3-10
-1 l
of the reactor depletes the amount of fuel available due to the absorption of neutrons by the U-235 atoms and their subsequent fission. ( ! The rate of U-235 depletion is directly proportional to the power level at which the reactor is operated. In addition, the fission process results in the formation of fission products, some of which readily absorb neutrons. These effects, depletion and the buildup of fission products, are partially offset by the buildup of plutonium, shown in Figure 4.3-2 for the 17 x 17 fuel assembly, which occurs due to the non-fission absorption of neutrons in U-238. Therefore, at the beginning of any cycle a reactivity reserve equal to the depletion of the fissionable fuel and the buildup of fission product poisons over the specified cycle life must be " built" into the reactor. This excess reactivity is controlled by removable neutron absorbing material in the form of boron dissolved in the primary coolant and burnable poison rods. The concentration of boric acid in the primary coolant is varied to provide control and to compensate for long-term reactivity require-ments. The concentration of the soluble neutron absorber is varied Os to compensate for reactivity changes due to fuel burnup, fission product poisoning including xenon and samarium, burnable poison de-pletion, and the cold-to-operating moderator temperature change. Using its normal makeup path, the Chemical and Volume Control System (CVCS) is capable of interting negative reactivity at a rate of approx-imately 30 pcm/ min when the reactor coolant boron concentration is 1000 ppm and approximately 35 pcm/ min when the reactor coolant boron concentration is 100 ppm. If the emergency boration path is used, the CVCS is capable of inserting negative reactivity at a rate of i approximately 65 pcm/ min when the reactor coolant concentration is 1000 ppm and approximately 75 pcm/ min when the reactor coolant boron concentration is 100 ppm. The peak burnout rate for xenon is 25 pcm/ min (Section 9.3.4.3.1 discusses the capability of the CVCS to counteract xenon decay). Rapid transient reactivity requirements and safety shutdown requirements are met with control rods. O 4.3-11
As the boron concentration is increased, the moderator temperature coefficient becomes less negative. The use of a soluble poison alone would result in a positive moderator coefficient at BOL for the first cycle. Therefore, burnable poison rods are used in the first core to reduce the soluble boron concentration sufficiently to insure that the moderator temperature coefficient is negative for power operating conditions. During operation the poison content in these rods is depleted thus adding positive reactivity to offset some of the negative reactivity from fuel depletion and fission product buildup. The depletion rate of the burnable poison rods is not critical since chemical shim is always available and flexible enough to cover any possible deviations in the expected burnable poison depletion rate. Figure 4.3-3 is a graph of a typical core depletion with and without burnable poison rods. Note that even at end-of-life conditions some residual poison remains in the burnable poison rods resulting in a net decrease in the first ' cycle lifetime. Upon completion of the first cycle all the burnable poison rods are normally removed because the moderator temperature coefficient in reload cores is sufficiently negative. b d In addition to reactivity control, the burnable poison rods are stra-tegically located to provide a favorable radial power distribution. Figure 4.3-4 shows the burnable poison distribution within a fuel assembly for the several burnable poison patterns used in a 17 x 17 array. The preliminary burnable poison loading pattern to be used in the core is shown in Figure 4.3-5. Tables 4.3-1 through 4.3-3 contain a summary of the reactor core design parameters, including reactivity coefficients, delayed neutron fraction and neutron lifetimes. Sufficient information is included to permit an independent calculation of the nuclear performance characteristics of the core. Table 4.3-1 through 4.3-3 lists parameters that are based upon prelim-inary engineerir.g calculations. Minor changes in these parameters are expected as the design progresses and specific operating requirements 4.3-11a L
._. . . . . . . . . . _ . - . . . - . . . - - - - . - . - - . . . . - - ~ . - . _ . - . . - . . . . . -
l ! l become known. For example, a need tohs' orten the first cycle in ! [ order to avoid an undesirable ~ refueling date would probably be accomplished ' by a modest reduction in fuel enrichment. Such changes are routine l in final design process, and need only be of concern in licensing reviews when their effect is to alter the nuclear design limits given in Table 4.3-2 or when they have impact on the conclusions reached in Chapter 15. ! l l l -i
.\.
1 O - u f I i l l l I' l i
~
I' !O l 4.3-11b i
4.3.2.2 Power Distributiens The accuracy of power distribution calculations has been confinned
\ through approximately one thousand flux maps during some twenty years of operation under conditions very similar to those expected for the plant described herein. Details of this confirmation are given in Reference 2 and in Section 4.3.2.2.6.
4.3.2.2.1 Definitions Power distributions are quantified in tenns of hot channel factors. These factors are a measure of the peak pellet power within the reactor core and the total energy produced in a coolant channel and are expressed in tenns of quantities related to the nuclear or thermal design namely: Power density is the thennal power produced per unit volume of thecore(KW/ liter). Linear power density is. the thermal power produced per unit length of active fuel (KW/FT). Since fuel assembly geometry is standardized this is the unit of power density most comonly used. For all practical purposes it differs from KW/ liter by a constant factor which includes geometry and the fraction of the total thermal power which is generated in the fuel rod. Average linear power density is the total thennal power produced in the fuel rods divided by the total active fuel length of all rods in the core. 4.3-12 O
1 Local heat flux is the heat flux at th; surface of the cladding (Btuft -2 hr ). For nominal rod parameters this differs from linear power density by a constant factor. (]~x Rod power or rod integral power is the length integrated linear power density in one rod (KW). Average rod power is the total thermal power produced in the fuel rods divided by the number of fuel rods (assuming all rods have equal length). The hot channel factors used in the discussion of power distributions in this section are defined as follows: Fq, Heat Flux Hot Channel Factor, is defined as the maximum local heat flux on the surface of a fuel rod divided by the average fuel rod heat flux, allowing for manufacturing tolerances on fuel pellets' and rods and including fuel densification effects. 1 N ( F . Nuclear Heat Flux Hot Channel Factor, is defined as the maximum local fuel rod linear power density divided by the average fuel rod linear power density, assuming nominal fuel pellet and rod parameters. Fh, Engineering Heat Flux Hot Channel Factor, is the allowance on 1 heat flux required for manufacturing tolerances. The engineering factor allows for local variations in enrichment, pellet density and diameter, surface area of the fuel rod ano eccentricity of the gap between pellet and clad. Combined statistically the net effect is a factor of 1.03 to be applied to fuel rod surface heat flux. F"H, a Nuclear Enthalpy Rise Hot Channel Factor, is defined as the ratio of the integral of linear power along the rod with the highest integrated power to the average rod power. 4.3-13 DECEMBER, 1973
!1 Manufacturing tolerances ~, hot channel power distribution and surrounding channel power distributions are treated explicity in the calculation
' (O i
.y of DNB ratio described in Section 4.4.
It is convenient for the purposes of discussion to define subfactors of Fg, however, design limits are set in terms of the total peaking
- l. factor.
i' l Fg = Total peaking factor or heat flux hot-channel factor l- Maximum KW/ft
, Average KW/ft l 1 i without densification effects F2 = F"gxF
= F"g x F"Z x F" U xF O where '
, F' and F are defined above.
F" = the measurement uncertainty associated with a full core flux map with moveable detectors. Fh ratio of peak power density to average power density in the horizontal plane of peak local power. F" = ratio of the power per unit core height in the horizontal plane of peak local power to the average value of power _ per unit core height. If the plane of peak local power coincides with the plane of maximum power per unit core height then F"Z is the Core Average Axial Peaking Factor. io i 4.3-14 DECEMBER, 1973 r -, -, g - - . - - - , ,
To include the allowances made for densification effects, which 7 are height dependent, the following quantities are defined. f 1 V S(Z) = the allowance made for densification effects at height Z in the core. See section 4.3.2.2.5. P(Z) = ratio of the power per unit core height in the hori-zontal plane at height Z to the average value of power per unit core height. 1 Then Fg = Total peaking factor Maximum Kw/ft
" Average Kw/ft Including densification allowa'nce N
Fq = max F p (Z) x P(Z) x S(Z) , x F"uxF 73 4.3.2.2.2 Radial Power Distributions U The power shape in horizontal sections of the core at full power is a function of the fuel and burnable poison loading patterns, the presence or absence of a single bank of full length control rods, and the presence or absence of the single bank of part length rods. Thus, at any time in the cycle any horizontal section of the core can be characterized as (a) unrodded, (b) with PL rods, (c) with group D control rods, or (d) with both PL rods and group D control rods. These four situations combined with burnup effects determine the radial power shapes which can exist in the core at full power. The effect on radial power shapes of power level, xenon, samarium and moderator density effects are considered also but these are quite small. The effect of non-uniform flow distribution is negligible. While radial power distributions in various planes of the core are often illustrated, the core radial enthalpy rise distribution as determined by the integral of power up each channel q is of greater interest. Figures 4.3-6 through 4.3-11 show representative V' radial power distributions for one eighth of the core for representative operating conditions. These conditions are (1) Hot Full Power 4.3-15
__ _ __ _ _ . _ ~ . _ ___ ___ . -. . . . _ _ _ _ .. . .___ k l l l (HFP) at Beginning of Life (B0L) - unrodded - no xenon, (2) HFP O at BOL - unrodded - equilibrium xenon, (3) HFP at BOL - Bank D ; l in - equilibrium xenon, (4) HFP at BOL - Bank D in, Part length l Rods in - equilibrium xenon (5) HFP at Middle of Life - unrodded - ) equilibrium xenon, and (6) HFP at End of Life - unrodded - equilibrium xenon. l Since the position of the hot channel varies from time to time a l single reference radial design power distribution is selected for DNB calculations. This reference power distribution is chosen con-servatively to concentrate power in one area of the core, minimizing j the benefits of flow redistribution. Assembly powers are normalized to core average power.
- 4.3.2.2.3 Assembly Power Distributions l
l For the purpose of illustration, assembly power distributions from l the BOL and E0L conditions corresponding to Figures 4.3-7 and 4.3-10, respectively, are given for the same assembly in Figures 4.3-v 12 and 4.3-13, respectively. l l l Since the detailed power distribution surrounding the hot channel
- j. Varies from time to time, a conservatively flat assembly power distribution is assumed in the DNB analysis, described in Section 4.4, with the rod of maximum integrated power artifically raised to the design value of F H.
Care is taken in the nuclear design of all fuel cycles and all operating conditions to ensure that a flatter assembly power distribution does not occur with limiting values of F AH* l l 4.3.2.2.4 Axial Power Distributions The shape of the power profile in the axial or vertical direction is largely under the control of the operator through the manual operation of the part length control rods and automatic motion of p full length rods responding to manual operation of the Chemical
.V 4.3-16
and Volume Control System. Nuclear effects which cause variations in the axial power shape include moderator density, Doppler effect on resonance absorption, spatial xenon and burnup. Automatically
. controlled variations in total power output and full length rod motion are also important in determining the axial power shape at any time. Signals are available to the operator from the excore ion chambers which are long ion chambers outside the reactor vessel running parallel to the axis of the core. Separate signals are taken from the top and bottom halves of the chambers. The difference between top and bottom signals from each of four pairs of detectors
, is displayed on the control panel and called the Flux Difference, al. Calculations of core average peaking factor for many plants and measurements from operating plants under many operating situations are associated with either al or axial offset in such a way that an upper bound can be placed on the peaking factor. For these correlations axial offset is defined as
* *b axial offset =t *t + *b >
O and $ and $ are the top and bottom detector readings. t b Representative axial power shapes from Reference 3 for BOL, MOL, and E0L conditions are shown in Figures 4.3-14 through 4.3-16. These figures cover a wide range of axial offset including values not permitted at full power. The radial power distributions shown in Figures 4.3-8 and 4.3-9 involving the partial insertion of control rops represent a synthesis of power shapes from the rodded and unrodded blanes. The applicability of the separability assumption upon which this procedure is based is assured through extensive three-dimensional calculations of possible rodded conditions. As an example, Figure 4.3-17 compares the axial power distribution for several assemblies at different distances from inserted control rods with the core average axial distribution. O 4.3-17
. The only significant difference from the average occurs in the low power peripheral assemblies, thus, confirming the validity of the separability assumption.
4.3.2.2.5 Local Power Peaking Fuel densification, which has been observed to occur under irradiation in several operating reactors, causes the fuel pellets to shrink both axially and radially. The pellet shrinkage combin. I with random hang-up of fuel pellets results in gaps in the fuel column when the pellets below the hung-up pellet settle in the fuel rod. The gaps vary in length and location in the fuel rod. Because of decreased neutron absorption in the vicinity of the gap, power peaking occurs in the adjacent fuel rods resulting in an increased power peaking factor. A quantitative measure of this local power peaking is given by the power spike factor S(Z) where Z is the axial location in the core. 1 The method used to compute the power spike factor is described in Reference 27 and is sununarized in Figure 4.3-18. The information flow outlined in Figure 4.3-18 is as follows:
- a. The probability that an axial gap of a certain size will occur at a given location in the core is determined from fuel per-formance data.
- b. The magnitude of the power spike caused by a single axial gap of a certain size is determined from nuclear calculations as shown in Figure 4.3-19 in which the effect of the change from the 15 x 15 to the 17 x 17 fuel rod array is illustrated. This curve is valid for uranium fuel enrichments up to 3.5 w/o.
- c. For each axial interval to be analyzed, axial gap occurrence probabilities and the single event power spikes are entered into the DRAW computer code. The code produces a curve of power spike vs. probability of exceeding power spike for each elevation in the core. The power census for a core is then 4.3-18 DECEMBER, 1973
l l 1 statistically combined with the pow;r spike probability
,, curve to obtain a power spike penalty for the core such j that less than one rod will exceed F at a 95% confidence level .
The power spike factor due to densification is assumed to be a local perturbation. Thus, densification affects F but q not F AH' 1 The magnitude of the increased power peaking increase from no effect at the bottom of the core to a few percent at the top of the core is shown in Figure 4.3-20, which is applicable to the 94.5% (geometric) dense pellets. Since there is little difference between the 15 x 15'and the 17 x 17 fuel rod bundle, the higher curve is used. 4.3.2.2.6 Limiting Power Distributions According to the ANS classification of plant conditions (See Section 15.0), Condition I occurrences are those which are expected frequently or regularly in the course of power operation, maintenance, or man-
) euvering of the plant. As such, Cor.dition I occurrences are accom-modated with margin between any plant parameter and the value of that parameter which would require either automatic or manual pro-tective action. Inasmuch as Condition I occurrences occur frequently or regularly, they must be considered from the point of view of affecting the consequences of fault conditions (Conditions II, III and IV). In this regard, analysis of each fault condition described is generally based on a conservative set of initial conditions corresponding to the most adverse set of conditions which can occur during Condition I operation.
The list of steady state and shutdown conditions, permissible devi-ations (such as one coolant loop out of service) and operational } transients is given in Section 15.1. Implicit in the definition of normal operation is proper and timely action by the reactor operator. That is, the operator follows recommended operating g V 4.3-19 DECEMBER, 1973
procedures for maintaining appropriate power distributions and takes any necessary remedial actions when alerted to do so by the plant instrumentation. Thus, as stated above, the worst or limiting power ' distribution which can occur during normal operation is to be considered as the starting point for analysis of ANS Conditions II, III and IV events. Improper procedural actions or errors by the operator are assumed in the design as occurrences of moderate frequency (ANS Condition l II). Some of the consequences which might result are listed in Section 15.2. Therefore, the limiting power shapes which result from such Condition II events, are those power shapes which deviate
- from the normal operating condition at the recommended axial offset band, e.g. due to lack of proper action by the operator during a xenon transient following a change in power level brought about by control rod motion. Power shapes which fall in this category are used for determination of the reactor protection system setpoints so as to maintain margin to overpower or DNB limits.
The means for maintaining power distributions within the required ! hot channel factor limits are described in the Technical Specifications. l A complete discussion of power distribution control in Westinghouse PWRs is included in Reference 4. Detailed information on the design constraints on local power density in a Westinghouse PWR, on the j defined operating procedures and on the measures taken to preclude exceeding design limits is presented in the Westinghouse Topical l Report on peaking factors (5) . The following paragraphs summarize these reports and describe the calculations used to establish the upper bound on peaking factors. The calculations used to establish the upper bound on peaking factors, Fq and FaH, include all of the nuclear effects which influence the radial and/or axial power distributions throughout core life for i various modes of operation including load follow, reduced power operation, and axial xenon transients. J ~O 4.3-20 a 4
Radial power distributions are calculated for the full power condition , and fuel and moderator temperature feedback effects are included . for the average enthalpy plane of the reactor. The steady state nuclear design calculations are done for normal flow with the same mass flow in each channel and flow redistribution effects neglected. l The effect of flow redistribution is calculated explicity where it is important in the DNB analysis of accidents. The effect of xenon on radial power distribution is small (compare Figures 4.3-6 l and 4.3-7) but is included as part of the normal design process. j Radial power distributions are relatively fixed and easily bounded with upper limits. i The core average axial profile, however, can experience significant changes which can occur rapidly as a result of rod motion and load changes and more slowly due to xenon distribution. For the study of points of closest approach to axial power distribution limits, several thousand cases en examined. Since the properties of the l nuclear design dictate what axial shapes can occur, boundaries on , the limits of interest can be set in terms of the parameters which ! are readily observed on the plant. Specifically, the nuclear design parameters which are significant to the axial power distribution l analysis are: l (a) core power level (b) core height (c) coolant temperature and flow (d) coolant temperature program as a function of reactor power (e) fuel cycle lifetimes (f) rod bank worths (g) rod bank overlaps (h) part length rod worth (1) part length rod length l r Normal operation of the plant assumes compliance with the following conditions: 4.3-21 DECEMBER, 1973
(1) Control' rods in a single bank move together with no individual rod insertion differing by more than 13 steps from the k bank demand position; (2) Control banks are sequenced with overlapping banks; (3) The control full length and part length bank insertion limits are not violated; (4) Axial power distributien procedures, which are aiven in terms of flux difference control and control bank position, are observed. The axial power distribution procedures referred to above are part of the required operating procedures which are followed in normal operation. Briefly they require control of the axial offset (flux difference x fractional power) at all power levels within a permissible operating band of a target value corresponding to the equilibrium full power valus. In the first cycle, the O-target value changes from about -10% to 0% linearly through the life of the cycle. This minimizes xenon transient effects on the axial power distribution. The target value is the same whether the part length rod bank is used to control the power distribution or not. When the part length rod bank is used, there is also a defined position requirement on the full length control rod banks which corresponds to a power dependent degree of insertion equal to the reactivity defect required to return to full power. Additional procedures cover insertion and removal of the part length rod bank from the core without violating power density limits. Calculations are performed for normal operation of the reactor including load following maneuvers. Beginning, middle and end -
- of cycle conditions are included in the calculations. Different histories of operation are assumed prior to calculating the effect of load follow transients on the axial power distribution. These O different histories assume base loaded operation and extensive 4.3-22
i load following with and without the use of part length rods. ! q Results of such calculations are shown in Figures 4.3-21 and 4.3-22 i V for the beginning of the first cycle and the end of the first cycle. l These calculated points have been synthesized from axial calculations l combined with radial factors appropriate for rodded and unrodded j planes in the first cycle. The calculated values have been increased ) by a factor of 1.05 for conservatism, a' factor of 1.03 for the : engineering factor FEand by the height dependent densification allowance given in Figure 4.3-20. , These figures are usd to detemine an upper bound on total peaking factor (F ) for nomal operation, and the value is used as input j to the Loss of Coolant Accident studies. Allowing for fuel densi-fication effect the average kw/ft at 3411 Nt is 5.45 kw/ft for i the four loop plant. For this application the upper limit on F including densification effects is 2.50 corresponding to a peak local power density of 13.9 kw/ft at 102% power. For the three I loop plant rated at 2652 MWt, the average linear power density i is 5.20 kw/ft and the peak local power at 102% is 13.3 kw/ft. q D To determine reactor protection system set points, with respect to power distributions, three categories of events are considered, namely rod control equipment malfunctions, operator errors of commission and operator errors of omission. The first category comprises uncontrolled rod withdrawal (with rods moving in the nomal bank sequence) for both full and part length banks. Also included are motions of the full length and part length banks below their insertion limits. Power distribu-tions were calculated throughout these occurrences assuming short tem corrective action, that is no transient xenon effects were considered to result from the malfunction. The event was assumed ' to occur from typical nomal operating situations which did include normal xenon transients. It was further assumed in determining the power distributions that total power level would be limited O i DECE!SER,1973 4.3-23 , 4 _ - e--- 7.-,._ y , 1
____ . _ _ _ _ m . _ _ __ __ . - . _ _ _ _.. ___ _ l l-by reactor trip to below 118%. Since the study is to determine protection limits with respect to power and axial offset, no credit was taken for trip set point reduction due to flux difference. The results are given in Figure 4.3-23 in units
'of kw/ft. The peak power density which can occur in such events, assuming reactor trip at or below 118%, is thus limited to 15.2 kw/ft for the four loop plant and 14.6 kw/ft for the three loop plant including uncertainties and densification j effects.
The second category assumes that the operator mis-positions ' the part length. rod bank or the full length rod bank within the insertion limits and creates short term and long term ! (xenon transient) conditions not included in normal operating conditions. The third category assumes that the operator l fails to take action to correct a flux difference violation and a xenon transient or oscillation is allowed to proceed l j . uncorrected. These results also appear on Figure 4.3-23 as , Fq multiplied by 102% power including an allowance for calorimetric error. The use made of the envelope requires that the boundary be valid, for very skew power distributions, at power levels ! less than full power. Thus, when credit is taken for trip set point reduction due to flux difference, no product of j Fqtimes the power level at which trip would occur exceeds 15.2 kw/ft (four loop) or 14.6 kw/ft (three loop), including uncertainties and densification effects. ; The collection of points resulting from operator errors of J omission and commission as distinguished from those points resulting from rod movement, is enveloped by the broken line portion at 14.55 kw/ft (four loop) and 13.88 kw/ft (three loop) on Figure 4.3-23. This can be viewed as defining an upper boundary on total peaking factor by reference only to the excore detector flux difference without any procedural requirement to maintain 4.3-24 DECEMBER, 1973
o al within a narrow band about a target value. -Therofere, this I envelope can, at the operator's option, be a va,3 reference for a comparison of LOCA peaking factor requirements in lieu of the target value procedure. As will be noted in the Tech- j l nical Specifications, this option has been utilized for power l operation. Analyses of possible operating power shapes for the reactor l described herein show that the appropriate hot channel factors Fqand F g for peak local power density and for DNB analysis l at full power are the values given in Table 4.3-2 and addressed in Technical Specification 16.3.10. l F can be increased with decreasing power as shown in the q Technical Specifications. Increasing F H with decreasing l power is permitted by the DNB' protection setpoints and allows l radial power shape changes with rod insertion to the insertion ) limits as described in Section 4.4.3.2. It has been determined that provided the above conditions 1 through 4 are observed, the Technical Specification limits are met. When a situation is possible in normal operation which could result in local power densities in excess of those assumed as the pre-condition for a subsequent hypothetical accident, but which would not itself cause fuel failure, administrative controls and alarms.are provided for returning the core to : 1 a safe condition. These alarms are described in detail in l Chapters 7 and 16. 1 4.3.2.2.7 Experimental Verification of Power Distribution Analysis This subject is discussed in depth in Reference 2. A summary of this report is given here. l O DECEMBER,1973 4.3-25 l
l In a measurement of peak local power density, Fg , with the moveable o detector system described in Section 7.7.1 and 4.4.5, the following
- k. uncertainties have to be considered:
(a) reproducibility of the measured signal j (b) errors in the calculated relationship between detector current and local flux ; (c) errors in the calculated relationship between detector flux and peak rod power some distance from the measurement thimble. The appropriate allowance for (a) above has been quantified by repetitive measurements made with several inter-calibrated detectors by using the common thimble features of the incore detector system. This system allows more than one detector to access any thimble. Errors in category (b) above are quantified to the extent possible, p by using the fluxes measured at one thimble location to predict fluxes at another location which is also measured. Local power distribution predictions are verified in critical experiments on arrays of rods with simulated guide thimbles, control rods, burnable poisons, etc. These critical experiments provide i quantification of errors of types (b) and (c) above, Reference 2 describes critical experiments performed at the Westinghouse Reactor Evaluation Center and measurement taken on two Westinghouse plants with incore systems of the same type as used in the plant described herein. The report concludes that the uncertainty associated with the peak nuclear heat flux factor, Fg is 4.58% at the 95. confidence level with only 5% of the measurements greater than the inferred value. This is the equivalent of a 2a limit on a normal distribution and is the uncertainty to be associated with a full core flux map with moveable detectors reduced with a reasonable set of input data incorporating the influence of burnup on the radial power distribution. The uncer-D tainty is usually rounded up to 5%. 4.3-26
l In comparing measured power distributions (or detector currents) against the calculations for the same situation it is not possible () to subtract out the detector reproducibility. Thus a comparison between measured and predicted power distributions has to include some measurement error. Such a comparison is given in Figure 4.3-24 for one of the maps used in Reference 2. Since the first publication of the report, hundreds of maps have been taken on these and other reactors. The results confirm the adequacy of the 5% uncertainty allowance on Fq . A similar analysis for the uncertainty in FAH (rod integral power) j mcasurements results in an allowance of 3.65% at the equivalent ; of a 2e confidence level. For historical reasons an 8% uncertainty factor is allowed in the nuclear design basis; that is, the predicted rod integrals at full power must not exceed the design FAH less j 8%. This 8% may be reduced in final design to 4% to allow a ! wider range of acceptable axial power distributions in the DNB analysis and still meet the design bases of Section 4.3.1.3. ' 1 : A V A recent measurement in the second cycle of a 121 assembly, 12 j foot, core is compared with a simplified one dimensional core ! average axial calculation in Figure 4.3-25. This calculation does not give explicit representation to the fuel grids. l The accumulated data on power distributions in actual operation is basically of three types: (1) Much of the data is obtained in steady state operation at constant power in the normal operating configuration; 1 (2) Data with unusual values of axial offset are obtained part of the excore detector calibration exercise which is performed monthly; (3) Special tests have been perfonned in load follow and other transient xenon conditions which have yielded i useful information on power distributions. 1 DECEMBER, 1973 4.3-27
These data are presented in detail in Reference 5. Figure 4.3-26 p contains a summary of measured values ofqF as a function of
\. axial offset for five plants from that report.
4.3.2.2.8 Testing A very extensive series of physics tests is performed on first cores. These tests and the criteria for satisfactory results are described in detail in Chapter 14. Since not all limiting situations can be created at beginning of life, the main purpose of the tests is to provide a check on the calculational methods used in the predictions for the conditions of the test. Tests performed at the beginning of each reload cycle are limited to verification of steady state power distributions, on the assumption that the reload fuel is supplied by the first core designer. , 4.3.2.2.9 Monitoring Instrumentation e The adequacy of instrument numbers, spatial deployment, required bq/ correlations between readings and peaking factors, calibration and errors are described in References 2, 4, and 5. The relevant conclusions are summarized here in Sections 4.3.2.2.7 and 4.4.5. Provided the limitations given in Section 4.3.2.2.6 on rod insertion and flux difference are observed, the excore detector system provides adequate monitoring of power distributions. Further details of specific limits on the observed rod positions and flux difference are given in the Technical Specifications, Section 16.3.10 together with a discussion of their bases. l Limits for alarms, reactor trip, etc. are given in the Technical Specification, Section 16.2.3. Descriptions of the systems provide.1 are given in Section 7.7. 1 4.3-28 l l
_ . _ . . ~ . - _ . . _ _ - ._ - _ . _ _ _ . . _ - _ _ _ _ _ . _ _ _ . _ . _ ___ 4.3.2.3 Reactivity Coefficients I i
)
The kinetic characteristics of the reactor core determine the response of the core to changing plant conditions or to operator adjustments made during normal operation, as well as the core response during abnormal or accidental transients. These kinetic characteristics are quantified in reactivity coefficients. The reactivity coefficients reflect the changes in the neutron multiplication due to varying plant conditions such as power, moderator or fuel temperatures, or less significantly due to a change in pressure or void conditions. Since reactivity l coefficients change during the life of the core, ranges of coefficients are employed in transient analysis to determine the response of the plant throughout life. The results of such simulations and the reactivity coefficients used are presented in Chapter 15. The analytical methods and calculational l models used in calculating the reactivity coefficitits are l given in Section 4.3.3. These models have been confirmed l through extensive testing abnormal or accidental transients. These kinetic characteristics are quantified in reactivity coefficients. The reactivity coefficients reflect the changes ! in the neutron multiplication due to varying plant conditions such as power, moderator or fuel temperatures, or less significantly l due to a change in pressure or void conditions. Since reactivity l coefficients change during the life of the core, ranges of I coefficients are employed in transient analysis to determine the response of the plant throughout life. The results' of i l such simulations and the reactivity coefficients used are l presented in Chapter 15. The analytical mtthods and calculational' ! models used in calculating the reactivity coefficients are given in Section 4.3.3. These models have been confirmed through extensive testing of more than thirty cores similar to the plant described herein; results of these tests are discussed in Section 4.3.3. Quantitative information for 1 l l O 1 4.3-29 1 i
-- ., e , . , . _ ,
calculated reactivity coefficients, including fuel-Doppler coefficient, moderator coefficients (density, temperature, pressure, void) and power coefficient is given in the following sections. 4.3.2.3.1 Fuel Temperature (Doppler) Coefficient The fuel temperature (Doppler) coefficient is defined as the change in reactivity per degree change in effective fuel temperature and is primarily a measure of the Doppler broadening of U-238 and Pu-240 resonance absorption peaks. Doppler broadening of other isotopes such as U-236, Np-237 etc. are also considered but their contributions to the Doppler effect is small. An
)
increase in fuel temperature increases the effective resonance absorption cross sections of the fuel and produces a correspon' ding reduction in reactivity. The fuel temperature coefficient is calculated by performing two-group X-Y calculations using an updated version of the TURTLE code. Moderator temperature is held constant and the power level is varied. Spatial variation of fuel temperature is taken into account by calculating the effective fuel temperature as a function of power density as discussed in Section 4.3.3.1. The Doppler temperature coefficient is shown in Figure 4.3-27 as a function of the effective fuel temperature (at beginning-of-life and end-of-life conditions). The effective fuel temperature is lower than the volume averaged fuel temperature since the neutron flux distribution is non-uniform through the pellet and gives preferential weight to the surface temperature. The Doppler-only contribution to the power coefficient, defined later, is shown in Figure 4.3-28 as a function of relative O 4.3-30 DECEMBER. 1973
l core power. The integral of the differential curve on Figure 4.3-28 is the Doppler contribution to the power defect and V is shown in Figure 4.3-29 as a function of relative power. The Doppler coefficient becomes more negative as a function - of life as the Pu-240 content increases, thus increasing the l Pu-240 resonance absorption but less negative as the fuel temperature changes with burnup as described in section 4.3.3.1. The upper and lower limits of Doppler coefficient used in l accident analyses are given in Section 15. 4.3.2.3.2 Moderat:- Coefficients l The tr0derator coefficient is a measure of the change in reactivity due to a change in specific coolant parameters such as density, temperature, pressure or void. The coefficients so obtained I are moderator density, temperature, pressure and void coefficients. Moderator Density and Temperature Coefficients The moderator temperature (density) coefficient is defined { as the change in reactivity per degree change in the moderator temperature. Generally, the effect of the changes in moderator l density as well as the temperature are considered together. A decrease in moderator density means less moderation which results in a negative moderator coefficient. An increase in coolant temperature, keeping the density constant, leads to a hardened neutron spectrum and results in an increase ' in resonance absorption in U-238, Pu-240 and other isotopes.
)
The hardened spectrum also causes a decrease in the fission to capture ratio in U-235 and Pu-239. Both of these effects make the moderator coefficient more negative. Since water density changes more rapidly with temperature as temperature increases, the moderator temperature (density) coefficient become more negative with increasing temperature. l l O DECEMBER,1973
. L/ 4.3-31
- . - _ _ . _ . _ . .. - - . . - - ~ . _
The soluble boron used in the reactor as a means of reactivity control also has an effect on moderator density coefficient O since the soluble boron poison density as well as the water density is decreased when the coolant temperature rises. A l decrease in the soluble poison concentration introduces a positive component in the moderator coefficient. l l Thus, if the concentration of soluble poison is large enough, l the net value of the coefficient may be positive. With the burnable poison rods present, however, the initial hot boron I concentration is sufficiently low that the moderator temperature l coefficient is negative at operating temperatures. The effect of control rods is to make the moderator coefficient more negative by reducing the required soluble boron concentration and by increasing the " leakage" of the core. With burnup, the moderator coefficient becomes more negative primarily as a result of boric acid dilution but also to a significant extent from the effects of the buildup of plutonium and fission products. The moderator coefficient is calculated for the various plant conditions discussed above by performing two-group X-Y calculations, varying the moderator temperature (and density) by about +_ 5'F about each of the mean temperatures. The moderator coefficient i is shown as a function of core temperature and boron concentration for the unrodded and rodded core in Figures 4.3-30 through 4.3-32. The temperature range covered is from cold (68 F) to about 600 F. The contribution due to Doppler coefficient
]
(because of change in moderator temperature) has been subtracted from these results. Figure 4.3-33 shows the hot, full power moderator temperature coefficient plotted as a function of first cycle lifetime for the just critical boron concentration condition based on the design boron letdown condition. O 4.3-32 l i
- r
The modarator coefficients presented here are calculated on a=corewise basis, since they are used to describe the core behavior in normal ' and accident situations when'the moderator temperature changes can : be considered to affect the whole core. 1 Moderator Pressure Coefficient The moderator pressure coefficient relates the change in moderator j density, resulting from a reactor coolant pressure change, to the ) corresponding effect on neutron production. This coefficient is ! j of much less significance in comparison with the moderator temperature l l coefficient. A change of 50 psi in pressure has approximately the ' i same effect on reactivity as a half-degree change in moderator temperature. This coefficient can be determined from the moderator ' temperature coefficient by relating change in pressure to the cor-responding change in density. The moderator pressure coefficient is negative over a portion of the moderator temperature range at beginning of life (-0.004 pcm/ psi, BOL) but is always positive at operating conditions and becomes more positive during life (+0.3 pcm/ psi,EOL). Moderator Void Coefficient The moderator void coefficient relates the change in neutron multi-plication to the presence of voids in the moderator. In a PWR this coefficient is not very significant because of the low void content in the coolant. The core void content is less than one-half of one percent and is due to local or statistical boiling. The void coefficient at BOL varies from 50 pcm/% void at low temperatures to -250 pcm/% void at operating temperatures. The negative void coefficient at operating temperature becomes more negative with fuel burnup. 4.3.2.3.3 Power Coefficient The combined effect of moderator temperature and fuel temperature change as the core power level changes is called the total power J DECEMBER,1973 4.3-33
i coefficient and is expressed in terms of reactivity change per percent power change. The power coefficient at BOL and E0L condi- I . tions is given in Figure 4.3-34. ! It becomes more negative with burnup reflecting the combined effect of moderator and fuel temperature coefficients with burnup. The : power defect (integral reactivity effect) at BOL and E0L is given in Figure 4.3-35. ' 4.3.2.3.4 Comparison of Calculated and Experimental Reactivity Coefficients l I Section 4.3.3 describes the comparison of calculated and experimental l reactivity coefficients in detail. Based on the data presented there, the accuracy of the current analytical model is: 1 2% op for Coppler and power defect 1 2 pcm/*F for the moderator coefficient Experimental evaluation of the calculated coefficients will be done during- the physics startup tests described in Chapter 14. 4.3.2.3.5 Reactivity Coefficients Used in Transient Analysis i i l Table 4.3-2 gives the representative ranges for the reactivity coefficients used in transient analysis. The exact values of the l coefficient used in the analysis depend on whether the transient of interest is examined at the beginning or end of life, whether "
)
mostnegativeorthemostpositive(leastnegative) coefficients ! are appropriate, and whether spatial nonuniformity must be considered in the analysis. Conservative values of coefficients, considering l
-various aspects of analysis are used in the transient analysis. j This is completely described in Chapter 15.
The values listed in Table 4.3-2 and illustrated in Figures 4.3-27 through 4.3-35 apply to the core described in Table 4.3-1. The 1 O DECEMBER, 1973 4.3-34 9 I
coefficients appropriate for use in subsequent cycles depends on the , core's operating history, the number and enrichment of fresh fuel assemblies, the loading pattern of burned and fresh fuel, the number and location of any burnable poison rods, etc. The need for a reevaluation of any accident in a subsequent cycle is contingent upon whether or not the coefficients for that cycle fall within the 1 identified range used in the analysis presented in Chapter 15. Control rod requirements are given in Table 4.3-3 for the core described and for a hypothetical equilibrium cycle since these are markedly different. These latter numbers are provided for information only and their validity in a particular cycle would be an unexpected coincidence. 4.3.2.4 Control Requirements To insure the shutdown margin stated in the Technical Specifications under conditions where a cooldown'to ambient temperature is required, concentrated soluble boron is added to the coolant. Boron concentrations for several core conditions are listed in Table 4.3-2. For all core condition's including refueling, the boron concentration is well below the solubility limit. The rod cluster control assemblies are employed to bring the reactor to the hot shutdown condition. The minimum required shutdown margin is given in Technical Specification 16.3.10. The ability to accomplish the shutdown for hot conditions is demon-strated in Table 4.3-3 by comparing the difference between the Rod Cluster Control Assembly reactivity available with an allowance for the worst stuck rod with that required for control and protection purposes. The shutdown margin includes an allowance of 10 percent for analytic uncertainties (see Section 4.3.2.4.9). The largest reactivity control requirement appears at the end-of-life (E0L) when the moderator temperatur<.e coefficient reaches its peak negative value as reflected in the larger power defect. The control rods are required to provide sufficient reactivity to account for the power defect from full power to zero power and O DECEMBER,1973 4.3-35
to provide the required shutdown margin. The reactivity addition resulting from power reduction consists of contributions from Doppler, variable average moderator temperature, flux redistribution, and reduction in void content as discussed below. 1 4.3.2.4.1 Doppler l The Doppler effect arises from the broadening of U-238 and Pu-240 j resonance peaks with an increase in effective pellet temperature. This effect is most noticeable over the range of zero power to full power cue to the large pellet temperature increase with power generation. 4.3.2.4.2 Variable Average Moderator Temperature When the core is shutdown to the hot, zero power condition, the average moderator temperature changes from the equilibrium full load value determined by the steam generator and turbine character-istics (steam pressure, heat transfer, tube fouling, etc.) to the { equilibrium no load value, which is based on the steam generator I shell side design pressure. The design change in temperature is conservatively increased by 4 F to account for the control dead band and measurement errors. Since the moderator coefficient is negative, there is a reactivity addition with power reduction. The moderator coefficient becomes more negative as the fuel depletes because the boron concentration j is reduced. This effect is the major contributor to the increased i requirement at end of life. 4.3.2.4.3 Redistribution During full power operation the coolant density decreases with core height, and this, together with partial insertion of control rods, results in less fuel depletion near the top of the core. Under steady state conditions, the relative power distribution will be 0 4.3-36 DECEMBER, 1973
slightly asymmetric towards the bottom of the core. On the other j hand, at hot zero power conditions, the coolant density is uniform l' up the core, and there is no flattening due to Doppler. The result l will be a flux distribution which at zero power can be skewed toward the top of the core. The reactivity insertion due to the skewed distribution is calculated with an allowance for the most adverse effects of xenon distribution and part-length rod position. 4.3.2.4.4 Void Content A small void content in the core is due to nucleate boiling at l full power. The void collapse coincident with power reduction
- makes a small reactivity contribution.
4.3.2.4.5 Rod Insertion Allowance At full power, the control bank is operated within a prescribed band of travel to compensate for small periodic changes in boron concentration, changes in temperature and very small changes in the xenon concentration not compensated for by a change in boron concentration. When the control bank reaches either limit of this band, a change in boron concentration is equired to compensate for additional reactivity changes. Since the insertion limit is set by a rod travel limit, a conservatively high calculation of the inserted worth is made which exceeds the normally inserted reactivity. 4.3.2.4.6 Burnup Excess reactivity of 10% ap to 25% ap (hot)isinstalledatthe beginning of each cycle to provide sufficient reactivity to compensate for fuel depletion and fission products throughout the cycle. This reactivity is controlled by the addition of soluble boron to the L l i C 4.3-37
l coolant and by burnable poison. The soluble boron cacentration - for several core configurations, the unit boron worth, and burnable poison worth are, given in Tables 4.3-1 and 4.3-2. Since the excess reactivity for burnup is controlled by soluble boron and/or burnable poison, it is not included in control rod requirements. 4.3.2.4.7 Xenon and Samarium Poisoning
- Changes in xenon and samarium concentrations in the core occur at a sufficiently slow rate, even following rapid power level changes, l
that the resulting reactivity change is controlled by changing the soluble boron concentration. 1 4.3.2.4.8 pH Effects Changes in reactivity due to a change in coolant pH, if any, are sufficiently small in magnitude and occur slowly enough to be con-trolled by the boron system. Further details are available in Reference 6. O 4.3.2.4.9 Experimental Confirmation Following a normal shutdown, the total core reactivity change during cooldown with a stuck rod has been measured on a 121 assembly, 10 ft, high core and a 121 assembly, 12 ft. high core. In each case, the core was allowed to cooldown until it reaches criticality simulat-ing the steamline break accident. For the ten foot core, the total reactivity change associated with the cooldown is overpredicted by ) about 0.3% ap with respect to the measured result. This represents an error of about 5% in the total reactivity change and is about half the uncertainty allowance for this quantity. For the 12 foot core, the difference between the measured and predicted reactivity change was an even smaller 0.2% ap. These measurements and others I demonstrate the ability of the methods described in Section 4.3.3 to accurately predict the total shutdown reactivity of the core. A U 4.3-38 DECEMBER, 1973
4.3.2.5 Control Core reactivity is controlled by means of a chemical poison dissolved in the coolant, Rod Cluster control assemblies, and burnable poison l rods as described below ' 1 l 4.3.2.5.1 Chemical Poison l Boron in solution as boric acid is used to control relatively slow reactivity changes associated with:
- 1. The moderator temperature defect in going from cold shutdown at ambient temperature to the hot operating temperature at p ro power,
- 2. The transient xenon and samarium poisoning, such as that follow-ing power changes or changes in Rod Cluster Control position, !
- 3. The excess reactivity required to compensate for the effects !
of fissile inventory depletion and buildup of long-life fission products. l
- 4. The burnable poison depletion.
l The boron concentrations for various core conditions are presented in Table 4.3-2. 4.3.2.5.2 Rod Cluster Control Assemblies Two types of Rod Cluster Control Assemblies are employed: full-l length assemblies and part-length assemblies. The nutubers of respective i full-length and part-length assemblies are shown in Table 4.3-
- 1. The full-length Rod Cluster Control Assemblies are used for shutdown and control purposes to offset fast reactivity changes associated with:
l l D DECEMBER,1973 4.3-39 1 L b
l
- 1. The required' shutdown margin in the hot zero power, stuck rods condi tion, ,
l O 2. The reactivity compensation as a result of an increase in power I I above hot zero power (power defect including Doppler, and moderator reactivity changes),
- 3. Unprogrammed fluctuations in boron concentration, coolant tempera-ture, or xenon concentration (with rods not exceeding the allowable ,
rod insertion limits),
- 4. Reactivity ramp rates resulting from load changes.
4 The allowed full length control bank reactivity insertion is limited at full power to maintain shutdown capability. As the. power leve'l is reduced, control rod reactivity requirements are also reduced l and more rod insertion is allowed. The control bank position is monitored and the operator is notified by an alarm if the limit is approached. The determination of the insertion limit uses conser- l1 vative xenon distributions and part-length rod locations. In addition, the Rod Cluster Control Assembly withdrawal pattern determined from these analyses is used in determining power distribution factors and in determining the maximum worth of an inserted Rod Cluster Control Assembly ejection accident. For further discussion, refer to the Technical Specifications on Rod Insertion Limits. l Power distribution, rod ejection and rod misalignment analyses are I based on the arrangement of the shutdown and control groups of the Rod Cluster Control Assemblies shown in Figure 4.3-36. All shutdown l Rod Cluster Control Assemblies are withdrawn before withdrawal of I the control banks is initiated. In going from zero to 100 percent power, control banks A, B, C and D are withdrawn sequentially. The limits of rod positions and further discussion on the basis for l rod insertion limits are provided in Figures 16.3-6 through 16.3-8 in the Technical Specifications, Section 16.3.10. Provisions are O 4.3-40 DECEMBER, 1973 i l I
_ . . .. ._ . ~ - . - . - . . - - . - - - - . - - - - - - . - . - _ made for changing the insertion limits, if necessary, on the basis of plant startup and final operating data. 4.3.2.5.3 Power Shaping With The Part Length Control Rod Bank l The function of the part-length rods is to control the axial power distribution during xenon transients and to limit axial xenon oscilla-tions should they occur. In steady state operation at constant power, insertion of the part-1 l length rods is not required. In special situations, however, where l power shaping may be required, part length rods may be inserted to their insertion limit during full power operation. I When the part length rods are in use, the axial position of the part-length rod bank depends on the full-length rod insertion, power level, xenon distribution, etc. The operator moves the part- l length rod bank in response to the indicated flux difference (axial l offset times power) to maintain the core within the technical specifica- l tion power distribution limits. Generally a significant portion of the neutron power generation takes place in the region containing the part-length rods. Hence, their effect on possible reactivity l insertion rates and power shapes is taken into account in the analyses. l When the part length rods are not in use the indicated flux difference is controlled by suitable positioning of the full length control rods through changes in boron concentration. The part length rods 1 do not nove following a reactor trip. i 4.3.2.5.4 Burnable Poison Rods The burnable poison rods provide partial control of the excess reactivity available during the first fuel cycle. In doing so, these rods prevent the moderator temperature coefficient from being positive at normal operating conditions. They perfonn this function by reducing the requirement for soluble poison in the moderator 4.3-41 l
P l at the beginning of the first fuel cycle as described previously. l . The burnable poison rod pattern in the core together with the number of rods per assembly is shown in Figure 4.3-5, while the arrangements within an assembly are displayed in Figure 4.3-4. The reactivity ' worth'of these rods is shown in Table 4.3-1. The boron in the _ rods is depleted with burnup but at a sufficiently slow rate so that the resulting critical concentration of soluble boron concentration - is such that the moderator temperature coefficient remains negative at all times for power operating conditions. 4.3.2.5.5 Peak Xenon Startup I Compensation for the peak xenon buildup is accomplished using the boron control system. Startup from the peak xenon condition is accomplished with a combination of rod motion and boron dilution. The boron dilution may be made at any time, including during the shutdown period, provided the shutdown margin is maintained. 4.3.2.5.6 Load Follow Control and Xenon Control O During load follow maneuvers, power changes are accomolished usina control rod motion and dilution or boration by the boron system as required. Control rod motion is limited by the control rod 1 insertion limits on both full length and part-length rods as provided in the Technical Specifications and discussed in sections 4.3.2.5.2 and 4.3.2.5.3 The power distribution is maintained within acceptable limits through the location of the part-length rod bank. Reactivity changes due to the changing xenon concentration can be controlled by rod motion and/or changes in the soluble boron concentration. 4.3.2.5.7 Burnup Control of the excess reactivity for burnup is accomplished using soluble boron and/or burnable poison. The boron concentration must be limited during operating conditions to insure the moderator O
't.) DECEMBER, 1973
l temperature coefficient is negative. Sufficient burnable poison is installed at the beginning of a cycle to give the desired cycle lifetime without exceeding the boron concentration limit. The practical minimum boron concentration is 10 ppm. l 4.3.2.6 Control Rod Patterns And Reactivity Worth The full-length Rod Cluster Control Assemblies are designated by function as the control groups and the shutdown groups. The terms
" group" and " bank" are used synonymously throughout this report to describe a particular grouping of control assemblies. The Rod Cluster Assembly pattern is displayed in Figure 4.3-36 which is not expected to change during the life of the plant. The control banks are labeled A, B, C and D and the shutdown banks are labeled SA, SB, etc. , as applicable. Each bank, although operated and controlled as a unit, is comprised of two subgroups. The axial l position of the full-length Rod Cluster Control Assemblies may' be controlled manually or automatically. These Rod Cluster Control Assemblies are all dropped into the core following actuation of reactor trip signals.
Two criteria have been employed for selection of the control groups. First the total reactivity wcrth must be adequate to meet the require-l ments specified in Table 4.3-3. Second, in view of the fact that these rods may be partially inserted at power operation, the total l power peaking factor should be low enough to ensure that the power capability requirements are met. Analyses indicate that the first requirement can be met either by a single group or by two or more banks whose total worth equals at least the required amount. The l axial power shape would be more peaked following movement of a single group of rods worth three to four percent ap; therefore, four banks (described as A, B, C, and 0 in Figure 4.3-36) each worth approximately one percent an have been selected. l l O 4.3-43 l l
The position of control banks for criticality under any reactor condition is determined by the concentration of boron in the coolant. On an approach to criticality boron is adjusted to ensure that I criticality will be achieved with control rods above the insertion limit set by shutdown and other considerations (See Technical Specifica-tions 16.3.10). Early in the cycle there may also be a withdrawal limit at low power to maintain a negative moderator temperature coefficient. Usual practice is to adjust boron to ensure that the rod position lies within the so-called maneuvering band, that is such that an escalation from zero power to full power does not require further adjustment of boron concentration. Ejected rod worths are given in Section 15.4.6 for several different conditions. Experimental confinnation of these worths can be found by reference to start-up test reports such as Reference 7. Allowable deviations due to misaligned control rods are discussed in Section 16.3.10.3.5. 1 i A representative calculation for two banks of control rods withdrawn simultaneously (rod withdrawal accident) is given in Figure 4.3 37 Calculation of control rod reactivity worth versus time following l
. reactor trip involves both control rod velocity and differential I reactivity worth. The rod position versus time of travel after rod release assumed is given in Figure 4.3-38. For nuclear design purposes, the reactivity worth versus rod position is calculated by a series of steady state calculations at various control rod positions assuming all rods out of the core as the initial position in order to minimize the initial reactivity insertion rate. Also, to be conservative, the rod of highest worth is assumed stuck out of the core and the flux distribution (and thus reactivity impor-tance) is assumed to be skewed to the bottom of the core. The result of these calculations is shown on Figure 4.3-39
. O v DECEMBER, 1973 4.3-44
The shutdown groups provide additional negative reactivity to assure an adequate shutdown margin. Shutdown margin is defined as the amount by which the core would be subcritical at hot shutdown if all Rod Cluster Control Assemblies are tripped, but assuming that the highest worth assembly remains fully withdrawn and no changes in xenon, boron or part length Rod Cluster Control Assembly positions take place. The loss of control rod worth due to the material irradiation is negligible since only bank D and part length rods may be in the core under normal operating conditions. The values given in Table 4.3-3 show that the available reactivity in withdrawn Rod Cluster Control Assemblies provides the design bases minimum shutdown margin allowing for the highest worth cluster to be at its fully withdrawn position. An allowance for uncertainty in the calculated worth of N-1 rods is made before determination of the shutdown margin. 4.3.2.7 Criticality Of Fuel Assemblies Criticality of fuel assemblies outside of the reactor is precluded by Mequate design of fuel transfer and fuel storage facilities and by administrative control procedures. This section identifies those criteria important to criticality safety analyses. New fuel is generally stored in fuel storage facilities with no water present but which are designed so as to prevent accidental criticality even if unborated water is present. In the analysis for the storage facilities, the fuel assemblies are assumed to be in their most reactive condition, namely fresh or undepleted and with no control rods or removable neutron absorbers present. Assemblies can not be closer together than the design separation provided by the storage facility except in special cases such as in fuel shipping containers where analyses are carried out to establish the acceptability of the design. The mechanical integrity of the fuel assembly is assumed and no credit is taken 4.3-45
for neutron absorption properties of the storage facility unless specifically included in the design. For full flooding with unborated water, the fuel assembly spacing of the facility provides essentially full nuclear isolation and k,ff for the array is no greater than eff f r the single most reactive fuel assembly. The criterion k for full flooding is k eff
# 0.90.
The fuel assembly (17 x 17 fuel rods) of standard design and 3.5 w/o enriched uranium oxide, without a control rod or burnable poison rods, fully flooded and reflected with cold clean water, has a k eff f about 0.85. Two such fuel assemblies spaced one inch apart with parallel axes 9.5 inches apart have a k of about 0.99. eff Three such fuel assemblies spaced one inch apart with parallel axes would be supercritical. An infinite number of dry fuel assemblies of this design would have a k,ff < 0.80. 4.3.2.8 Stability O 4.3.2.8.1 Introduction The stability of the PWR cores against xenon-induced spatial oscilla-tions and the control of such transients are discussed extensively in References 4, 8, 9 and 10. A summary of these reports is given in the following discussion and the design bases are given in Section 4.3.1.6. In a large reactor core, xenon-induced oscillations can take place with no corresponding change in the total power of the core. The oscillation may be caused by a power shift in the core which occurs rapidly by comparison with the xenon-iodine time constants. Such a power shift occurs in the axial direction when a plant load change is made by control rod motion and results in a change in the moderator density and fuel temperature distributions. Such a power shift O - 4.3-46 l
l l 4 could occur in the diametral plane of the core as a result of abnormal l control action. l l
. Due to the negative power coefficient of reactivity, PWR cores are inherently stable to oscillations in total power. Protection against total power instabilities is provided by the fontrol and Protection System as described in Section 7.7. Hence, the discussion i on the core stability will be limited here to xenon-induced spatial I
oscillations. 4.3.2.8.2 Stability Index I Power distributions, either in the axial direction or in the X-Y plane, can undergo oscillations due to perturbations introduced in the equilibrium distributions without changing the total core power. The xenon-induced oscillations are essentially limited to the first flux overtones in the current PWR's, and the stability of the core against xenon-induced oscillations can be determined in terms of the eigenvalues of the first flux overtones. Writing, either in the axial direction or in the X-Y plane, the eigenvalue c of the first flux hannonic as c = b + ic. (1) then b is defined as the stability index and T = 2n/c as the oscillation 1 period of the first harmonic. The time-dependence of the first hannonic 64 in the power distribution can now be represented as 64(t)=Aett , ,,bt cos et, (2) where A and a are constants. The stability index can also be obtained approximately by: b = f in An (3) O 4.3-47 DECEMBER, 1973
where A n ' ^n+1 are the successive peak amplitudes of the oscillation and T is the time period between the successive peaks. 4.3.2.8.3 Prediction Of The Core Stability The stability of the core described herein (i.e. with 17 x 17 fuel assemblies) against xenon-induced spatial oscillations is expected to be equal to or better than that of earlier designs. The prediction is based on a comparison of the parameters which are significant in detemining the stability of the core against the xenon-induced oscillations, namely (a) the overall core size is unchanged and spatial power distributions will be similar, (b) the moderator temperature coefficient is expected to be similar to or slightly more negative, and (c) the Doppler coefficient of reactivity is expected to be equal to or slightly more negative at full power. Analysis of both the axial and X-Y xenon transient tests, discussed in Section 4.3.2.8.5, shows that the calculational model is adequate for the prediction of core stability. 4.3.2.8.4 Stability Measurements
- 1. Axial Measurements Two axial xenon transient tests conducted in a PWR with a core height of 12 feet and 121 fuel assemblies is reported in Reference ll, and will be briefly discussed here. The tests were performed at approximately 10% and 50% of cycle life.
Both a free-running oscillation test and a controlled test were perfomed during the first test. The second test at mid-cycle consisted of a free-running oscillation test only. In each of the free-running oscillation tests, a perturbation was introduced to the equilibrium power distribution through O 4.3-48
an impulse motion of the Control Bank D and the subsequent g oscillation was monitored to measure the stability index and () the oscillation period. In the controlled test conducted early in the cycle, the part-length (P/L) rods were used to follow the oscillations to maintain an axial offset (AO) within the prescribed limits. The A0 of power was obtained from the excore ion chamber readings (which had been calibrated j against the incore flux maps) as a function of time for both 1
'ree-running tests as shown in Figure 4.3-40.
l The total core poer was maintained constant during these spatial xenon tests, and the stability index and the oscillation l 1 period were obtained from a least-square fit of the A0 data i l in the form of Eq. (2). The A0 of power is the quantity that ; properly represents the axial stability in the sense that ! it essentially eliminates any contribution from even order l narmonics including the fundamental mode. The conclusions l of the tests are: ,~ () a. The core was stable against induced axial xenon transients both at the core average burnups of 1550 MWD /MTU and 7700 MWD /MTU. The measured stability indices are -0.041 1 hr-I for the first test (Curve 1 of Figure 4.3-40) and !
-0.014 hr-I for the second test (Curve 2 of Figure 4.3-40). The corresponding oscillation periods are 32.4 hrs. and 27.2 hrs. , respectively.
- b. The P/L rods were sufficient to shape the axial power profile and to dampen the axial xenon oscillations effec-tively.
- c. The reactor core becomes less stable as fuel burnup progresses and the axial stability index was essentially zero at 12,000 MWD /T.
$q / DECEMBER,1973 4.3-49
l 2. Measurements in the X-Y Plane Two X-Y xenon oscillation tests were performed at a PWR plant with a core height of 12 feet and 157 fuel assemblies. This plant has the highest power output of any Westinghouse PWR l l currentlyinoperation(1972). The first test was conducted l at a core average burnup of 1540 MWD /MTU and the second at l ! a core average burnup of 12900 MWD /MTV. Both of the X-Y xenon tests show that the core was stable in the X-Y plane at both
- bu rnups. The second test shows that the core became more stable as the fuel burnup increased and all Westinghouse PWR's with 121 and 157 assemblies are expected to be stable throughout their burnup cycles.
l In each of the two X-Y tests, a perturbation was introduced to the equilibrium power distribution through an impulse motion of one RCC unit located along the diagonal axis. Following the p,erturbation, the uncontrolled oscillation was monitored using the moveable detector and thennocouple system and the l ! excore power range detectors. The quadrant tilt difference l l 1s the quantity that properly represents the diametral oscillation I in the X-Y plane of the reactor core in that the differences ! of the quadrant average powers over two symmetrically opposite quadrants essentially eliminates the contribution to the oscilla-tion from the azimuthal mode. The quadrant tilt difference j (QTD) data were fitted in the form of Eq. (2) through a least-square method. A stability index of -0.076 hr~I with a period l of 29.6 hours was obtained from the thennocouple data shown in Figure 4.3- 41. i l It was observed in the second X-Y xenon test that the PWR L core with 157 fuel assemblies had become more stable due to an increased fuel depletion and the stability index was not determined. A , U 4.3-50
4.3.2.8.5 Comparison of Calculations with Measurements The analysis of the axial xenon transient tests was performed in an axial-slab geometry using a flux synthesis technique. The direct simulation of the A0 data was carried out using the PANDA Code [13] The analysis of the X-Y xenon transient tests was performed in an X-Y geometry using a modified TURTLE bl43 code. Both the PANDA and TURTLE codes solve the two-group time-dependent neutron diffusion equation with time-dependent xenon and iodine concentrations. The fuel temperature and moderator density feedback is limited to a steady-state model. All the X-Y calculations were performed j in an average enthalpy plane. 1 The basic nuclear cross-sections used in this study were generated from a unit cell depletion program which has evolved from the codes LEOPARD [15] and CINDEREI63. The detailed experimental data during the tests including the reactor power level, enthalpy rise and the impulse motion of the control rod assembly, as well as the plant follow burnup data were closely simulated in the study. 4 O The results of the stability calculation for the axial tests are compared with the experimental data in Table 4.3-4. The calcula-tions show conservative results for both of the axial tests with a margin of approximately 0.01 hr -I in the stability index. An analytical simulation of the first X-Y xenon oscillation test shows a calculated stability index of -0.081 hr -I , in goor' .T ee-
-I ment with. the measured value of -0.076 hr . As indicated eurlier, the second X-Y xenon test showed that the core had become more stable compared to the first test and no evaluation of the stability index was attempted. This increase in the core stability in the X-Y plane due to increased fuel burnup is due mainly to the increased ma'gnitude of the negative moderator temperature coefficient.
Previous studies of the physics of xenon oscillations, including three-dimensional analysis, are reported in the series of topical O - 4.3-51
___ __ _ _ ._ ~ _. reports, References 8, 9 and 10. A more detailed description of the experimental results and analysis of the axial and X-Y xenon transient tests is presented in References 11 and Section 1 of Reference 12. 4.3.2.8.6 Stability Control and Drotection The excore detector system is utilized to provide indications of xenon-induced spatial oscillations. The readings from the excore j detectors are available to the operator and also fonn part of the ' protection system.
- 1. Axial Power Distribution For maintenance of proper axial power distributions, the operator is instructed to maintain an axial offset within a prescribed I operating band, based on the excore detector readings. Should )
the axial offset be permitted to move far enough outside this
]
band, the protection limit will be reached and the power will I be automatically cut back. L Part-length control rods are provided for control of the axial power shape. As reported for one of the axial xenon transient tests in Section 4.3.2.8.4, the part-length rods were suffi-cient to shape the axial power distribution and to dampen the axial xenon oscillations effectively.
- 2. Radial Power Distribution The core described herein is calculated to be stable against X-Y xenon induced oscillations at all times in life.
The X-Y stability of large PWR's will be further verified as part of the startup physics test program at a PWR core with 193 fuel assemblies. The measured X-Y stability of the PWR core with 157 assemblies and the good agreement between O 4.3-52
.. ~ . ___ _ _ _ _ _ _ __ _ . _ _ __ __ - _ _ _ _ _ . _ _ _
I the calculated and measured stability index for this core, as discussed in Sections 4.3.2.8.4 and 4.3.2.8.5, make it very unlikely that a sustained X-Y oscillation can occur in a core with 193 assemblies. As discussed in Section 4.3.2.8.2, the X-Y stability of the new model core (i.e., with 17 x 17 l fuel assemblies) is expected to be equal to or better than the earlier PWR cores. However, in the unlikely event that X-Y oscillations occur, back-up actions are possible and would be implemented, if necessary, to increase the natural stability of the core until tests demonstrate a suitable stability. This is based on the fact that several actions could be taken to make the moderator temperature coefficient more negative, which will increase the stability of the core in the X-Y plane. l l Provisions for protection against non-symmetric perturbations l in the X-Y power distribution that could result from equipment malfunctions are made in the protection system design. This includes control rod drop, rod misalignment and asymmetric loss of coolant flow. l A more detailed discussien of the power distribution control l in PWR cores is presented in Reference 4. 4.3.2.9 Vessel Irradiation i It is beyond the scope of this section to present methods and analyses used in determination of neutron and gamma flux attenuation between the core and the pressure vessel other than a brief review given below. A more complete discussion on the pressure vessel irradia-tion and surveillance program is given in Section 5.4.3.6. The primary shielding material that serves to attenuate high energy neutron and gama flux originating in the core consists primarily l l 4.3-53 I~
- 1. -, _. . _ - - - -
of the core baffle, core barrel, the neutron pads, and associated water annuli, all of which are within the region between the core and the pressure vessel. In general, few group neutron diffusion theory codes are used to determine flux and fission power density distributions within the active core and the accuracy of these analyses is verified by in-core measurements on operating reactors. Outside the active core, methods such as those which use multigroup space dependent slowing down codes described in Section 5.4.3.6.2 are used. Region-wise power sharing infonnation from the core calculations is often used as reference source data for the multigroup codes. The neutron flux distribution and spectrum in the various structural components varies significantly from the core to the pressure vessel. Representative values of the neutron flux distribution and spectrum are presented in Table 4.3-5. The values listed are based on equi-librium cycle reactor core parameters and power distributions, and thus, are suitable for long term nyt projections and for correlation with radiation damage estimates. As discussed in Section 5.4.3.6, the irradiation surveillance program utilizes actual test samples to verify the accuracy of the calculated fluxes at the vessel. 4.3.3 ANALYTICAL METHODS Calculations required in nuclear design consist of three distinct types, which are performed in sequence:
- 1. detennination of effective fuel temperatures
- 2. generation of macroscopic few-group parameters
- 3. space-dependent, few-group diffusion calculations These calculations are carried out by computer codes which can be executed individually, however, at Westinghouse most of the O
4.3-54 l
l codes required have been linked to fom an automated design sequence which minimizes design time, avoids errors in transcription of (/ data, and standardizes the design methods. 4.3.3.1 Fuel Temperature (Doppler) Calculations 3 Temperatures vary radially within the fuel rod, depending on the heat generation rate in the pellet, the conductivity of the materials in the pellet, gap, and clad; and the temperature of the coolant. The fuel temperatures for use in most nuclear design Docpler calcula- 1 tions are obtained from a simplified version of the Westinghouse fuel rod design model described in Section 4.2.1.3.1 which considers the effect of radial variation of pellet conductivity, expansion-coefficient and heat generation rate, elastic deflection of th5 clad, and a gap conductance wh'ich depends on the initial fill gas, the hot open gap dimension, and the fraction of the pellet over which the gap is closed. The fraction of the gap assumed closed represents an empirical adjustment used to produce good g() agreement with observed reactivity data at BOL. Further gap closure occurs with burnup and accounts for the decrease in Doppler defect with burnup which has been observed in operating plants. For detailed calculations of the Doppler coefficient, such as for use in xenon stability calculations, a more sophisticated temperature model is used which accounts for the effects of fuel swelling, fission gas release, and plastic clad deformation. Radial power distributions in the pellet as a function of burnup are obtained from LASER U73 calculations. The effective U-238 temperature for resonance absorption is obtained from the radial temperature distribution by applying a radially dependent weic hting function. The weighting function was determined from REPAD b0 Monte Carlo calculations of resonance escape probabilities in several steady state and transient temperature distributions. In f 4.3-55 DECEMBER, 1973
1 each case a flat pellet temperature was determined which produced the same resonance escape probability as the actual distribution. The weighting function was empirically determined from these results. The effective Pu-240 temperature for resonance absorption is determined by a convolution of the radial distribution of Pu-240 number densities from Laser burnup calculations and the radial weighting function. The resulting temperature is burnup dependent, but the difference between U-238 and Pu-240 temperatures, in tenns of reactivity effects, is small. The effective pellet temperature for pellet dimensional change is that value which produces the same outer pellet radius in a virgin pellet as that obtained from the temperature model. The ' effective clad temperature for dimensional change is its average value. I The temperature calculational model has been validated by plant Doppler defect data as shown in Table 4.3-6 and Doppler coefficient data as shown in Figure 4.3-42. Stability index measurements also provide a sensitive measure of the Doppler coefficient near full power (See Section 4.3.2.8). It can be seen that Doppler defect data is typically within 0.2% 6p of prediction. 4.3.3.2 Macroscopic Group Constants 1 l Macroscopic few-group constants and analogous microscopic cross I sections (needed for feedback and microscopic depletion calculations) are generated for fuel cells by a recent version of the LEOPARD [15] l and CINDER b63 Codes, which are linked internally and provide burnup dependent cross sections. Nonnally a simplified approximation l of the main fuel chains is used; however, where needed, a complete l solution for all the significant isotopes in the fuel chains from 4.3-56
l Th-232 to Cm-244 is availableEI93 Fast and thermal cross section library tapes contain microscopic cross sections taken for the most part from the ENDF/B[20] library, with a few exceptions where . other data provide better agreement with critical experiments, isotopic measurements, and plant critical boron values. The effect on the unit fuel cell of non-lattice components in the fuel assembly is obtained by supplying an appropriate volume fraction of these , materials in an extra region which is homoganized with the unit cell in the fast (MUFT) and thermal (S0F0CATE) flux calculations. In the thermal calculation, the fuel rod, clad, and moderator are homogenized by energy-dependex disadvantage factors derived from an analytical fit to integral transport theory results. l Group constants for burnable poison cells, guide thimbles, instrument thimbles, and interassembly gaps are generated in a manner analogous to the fuel cell calculation. Reflector group constants are taken from infinite medium LEOPARD calculations. Baffle group constants are calculated from an average of core and radial reflector micro-scopic group constants for stainless steel. O Group constants for control rods are calculated in a linked version I of the HAMMER [21] and AIM [22] codes to provide an improved treatment of self shielding in the broad resonance structure of these isotopes at epithermal energies than is available in LEOPARD. The Doppler i broadened cross sections of the control rod materials are represented as smooth cross sections in the 54-group LEOPARD fast group structure ) and in 30 thermal groups. The four-group constants in the rod cell and appropriate extra region are generated in the coupled space-energy transport HAMMER calculation. A corresponding AIM calculation of the homogenized rod cell with extra region is used to adjust the absorption cross sections of the rod cell to match the reaction rates in HA!9IER. These transport-equivalent group constants are reduced to two-group constants for use in space-dependent diffusion calculations. In discrete X-Y calculations l 4.3-57 i
l I only one mesh intierval per cell is used, and the rod group constants are further adjusted for use in this standard mesh by reaction rate matching the standard mesh unit assembly to a fine-mesh unit l assembly calculation. Validation of the cross section method is based on analysis of critical experiments as shown in Table 4.3-7, isotopic data as ! shown in Table 4.3-8, plant critical boron (C B) values at HZP, BOL, as shown in Table 4.3-9 and at HFP as a function of burnup as shown in Figures 4.3-43 through 4.3-45. Control rod worth measurements are shown in Table 4.3-10. Confirmatory critical experiments on burnable poisons are described in Reference 23, 4.3.3.3- Spatial Few-Group Diffusion Calculations Spatial few-group diffusion calculations consist primarily of two-group X-Y calculations using an updated version of the TURTLE code and two-group axial calculations using an updated version of the PANDA code. Discrete X-Y calculations (1 mesh per cell) are carried out to determine critical boron concentrations and power distributions in the X-Y plane. An axial average in the X-Y plane is obtained by synthesis from unrodded and rodded planes. Axial effects in unrodded depletion calculations are accounted for by the axial buckling, which varies with burnup and is determined by radial depletion calculations which are matched in reactivity to the analogous I R-Z depletion calculation. The moderator coefficient is evaluated by varying the inlet temperature in the same X-Y calculations used for power distribution and reactivity predictions. ; Validation of TURTLE reactivity calculations is associated with the validation of the group constants themselves, as discussed O V 2 i 4.3-58 l
-~ _ _ _ _ _ _ _ _ . . _ _ _ _ . _ _ _ . _ _ _ _ . _ _ . _ _ _ . . _ _ _ _ _ . - . _
in Section 4.3.3.2. Validation of the Doppler calculations is , associated with the fuel temperature validation discussed in Section 4.3.3.1. Validation of the moderator coefficient calculations , l is obtained by comparison with plant measurements at hot zero l power conditions as shown in Table 4.3-11. ) i l l 1 ! Axial calculations are used to determine differential control rod worth curves (reactivity versus rod insertion) and axial power shapes during steady state and transient xenon conditions (flyspeck curve). Group constants and the radial buckling used in the axial calculation are obtained from the PANDA radial calculation, in which group constants in annular rings representing the various material regions in the X-Y plane are homogenized by flux-volume weighting. Validation of the spatial codes for calculating power distributions involves the use of incore and excore detectors and is discussed ! in Section 4.3.2.2.7. l Based on comparison with measured data it is estimated that the ) accuracy of current alialytical methods is: 10.2% Ao for Doppler defect i 2 x 10-5 /*F for moderator coefficient i 50 ppm for critical boron concentration with depletion 13% for power distributions 1 0.2% ao for rod bank worth 4.3.4 CHANGES This document has been prepared to support the Westinghouse change over from 15 x 15 to 17 x 17 fuel rod array in reference plant designs. 1 Whereas the basic approach to nuclear design is unchanged with the conver-sion to 17 x 17 values of parameters such as rod worth, coefficients, etc., lO 4.3-59 DECEMBER, 1973
+
changes to a small extent. Values of these parameters are given ) in the tables of this section. , V 4.
3.5 REFERENCES
- 1. " Anticipated Transients Without Reactor Trip in Westinghouse Pressurized Water Reactors", Topical Report, WCAP-8096, April, 1973.
- 2. F. L. Langford and R. J. Nath, Jr. " Evaluation of Nuclear Hot
' Channel Factor Uncertainties WCAP-7308-L, April,1969 (West-1 inghouse Proprietary) and WCAP-7810, December,1971.
- 3. A. F. McFarlane, " Core Power Capability in Westinghouse PWRs,"
WCAP-7267-L, October,1969 (Westinghouse Proprietary) and I WCAP-7809 December, 1971.-
- 4. J. S. Moore " Power Distributon Control of Westinghouse Pressurizer Water Reactors," WCAP-7208, September.1968 (Westinghouse Proprietary) I and WCAP-7811, December,1971.
- 5. A. F. McFarlane, " Power Peaking Factors," WCAP-7912-L, March, 1971 (Westinghouse Proprietary) and WCAP-7912, March,1972. l1
- 6. J. O. Cermak et al, " Pressurized Water Reactor pH - Reactivity Effect" Final Report, WCAP-3696-8 (EURAEC-2074), October, 1968.
- 7. J. E. Outzs, " Plant Startup Test Report, H. B. Robinson Unit No. 2," WCAP-7844, January,1972.
- 8. C. G. Poncelet and A. M. Christie, " Xenon-Induced Spatial Instal!!ities in Large PWRs," WCAP 3680-20,(EURAEC-1974),
March, 1968. 4.3-60 DECEMBER, 1973
l I i
- 9. F. 8. Skogen and A. F. McFarlane, " Control Procedures for Xenon-Induced X-Y Instabilities in large PWRs," WCAP 3680-21, O
v (EURAEC-2111), February,1969.
- 10. F. B. Skogen and A. F. McFarlane " Xenon-Induced Spatial Instabili-ties in Three-Dimensions," WCAP-3680-22 (EURAEC-2116), September, 1969.
- 11. J. C. Lee, et al, " Axial Xenon Transient Tests at the Rochester i
Gas and Electric Reactor," WCAP-7964, June,1971.
- 12. C. J. Kubit, " Safety Related Research and Development for West-inghouse Pressurized Water Reactors, Program Summaries, Fall 1972," WCAP-8004, January,1973.
- 13. R. F. Barry, et al, "The PANDA Code," WCAP-7757, September,1971.
i
- 14. S. Altomare and R. F. Barry, "The TURTLE 24.0 Diffusion Depletion Code," WCAP-7758, September,1971. '
l
)
V(3
- 15. R. F. Barry, " LEOPARD - A Spectrum Dependent Non-Spatial Depletion Code for the IBM-7094," WCAP-3269-26, September,1963.
I 16 T. R. England, " CINDER - A One-Point Depletion and Fission Product Program," WAPD-TM-334, August, 1962.
- 17. C. G. Poncelet, " LASER - A Depletion Program for Lattice Calculations Based on MUFT and THERM 0S," WCAP-6073, April,1966. 1
- 18. J. E. 01hoeft, "The Doppler Effect for a Non-Uniform Temperature Distribution in Reactor Fuel Elements," WCAP-2048, July,1962.
- 19. R. J. Nodvik, et al, " Supplementary Report on Evaluation of Mass Spectrometric and Radiochemical Analyses of Yankee Core I Spent Fuel, Including Isotopes of Elements Thorium Through Curium." WCAP-6086, August,1969.
O - 4.3-61 DECEMBER, 1973
- 20. M. K. Drake, Ed. , " Data Formats and Procedure for the ENDF
! Neutron Cross Section Library," BNL-50274, ENDF-102, Vol. I,
,)
( 1970.
- 21. J. E. Suich and H. C. Honeck, "The HAMMER System, Heterogeneous Analysis by Multigroup Methods of Exponentials and Reactors,"
DP-1064, January,1967.
- 22. H. P. Flatt and D. C. Baller, " AIM-5, A Multigroup, One Dimensional 1 Diffusion Equation Code," NAA-SR-4694, March, 1960.
- 23. J. S. Moore, " Nuclear Design of Westinghouse Pressurized Water Reactors with Burnable Poison Rods," WCAP-7806, December.1971.
- 24. L. E. Strawbridge and R. F. Barry, " Criticality Calculations for Uniform Water-Moderated Lattices," Nucl . Sci . and Eng.
2_3, 3 58 (1965).
,~ 25. R. J. Nodvik, "Saxton Cor6 2, Fuel Perfonnance Evaluation, Part II," WCAP-3385-56, Pt. II, "Saxton Core II Fuel Performance Evaluation," July,1970.
- 26. R. D. Leamer, et al , "PU02 -U0 2 Fueled Critical Experiments,"
WCAP-3726-1, July, 1967.
- 27. J. M. Hellman, (Ed.), " Fuel Densification Experimental Results and Model for Reactor Application," WCAP-8218 (Westinghouse j Proprietary) October,1973, and WCAP-8219, October,1973.
n ('"') 4.3-62 DECEMBER,1973 l
. _ . - . -. .- ~. - - . - - --.---.- - -_- .-
l j. TABLE 4.3-1 Sheet 1 of 3 ; REACTOR CORE DESCRIPTION (Four Loc 9, First Cycle) I Active Core ; Equivalent Diameter, in. 132.7 f l
. Core Average Active Fuel Height, '
l
'I First Core, in. 143.7 Height-to-Diameter Ratio 1.09 Total Cross-Section Area, ft 2 96.06 H 0/U Molecular Ratio, Lattice (Cold) 3.43 2
Reflector Thickness and Composition Top - Water plus Steel, in. s10 Bottom - Water plus Steel, in. s10 I E Side - Water plus Steel, inc. $15 Fuel Assemblies Number 193 Rod Array 17 x 17 Rods per Assembly 264 Rod Pitch, in. 0.496 Overall Transverse Dimensions, in. 8.426 x 8.426 Fuel Weight (as U0 2
), lbs. 222,739 Zircaloy Weight, 1bs. 50,913 Number of Grids per Assembly 7-R type 1 l Composition of grids INC718 j l- Weight of Grids (Effective in Core) lbs 1992 )
Number of Guide Thimbles per Assembly 24 ; Composition of Guide Thimbles Zircaloy 4 Diameter of Guide Thimbles (upper 0.450 I.D. x 0.482 0.D. part),in. Diameter of Guide Thimbles (lower part),in. 0.397 I.D. x 0.4290 0.D. Diameter of Instrument Guide Thimbles, in. 0.450 I.D. x 0.4820 0.D. 4.3-63 DECEMBER, 1973
. . ...- - -. ~ . . _ . . . - - . . . - - - - . . - . _ . . - . _ . - . ~ . ,
TABLE 4.3-1 (Continued) Sheet 2 of 3 REACTOR CORE DESCRIPTION l (Four Loop, First Cycle) 1 Fuel Rods-
]
Number 50,952 Outside Diameter, in. 0.374 Diameter Gap, in. , 0.0065 Clad Thickness, In. 0.0225 l Clad Material Zircaloy-4 Fuel Pellets Material 002 Sintered Density (percent of Theoretical) 95 j Fuel Enrichments w/o Region l' 2.10 Region 2 : 2.60 Region 3 3.10 V Diameter, in. 0.3225 Length, in. 0.530 g,3 4,.* *t/p4/ Mass of U02 per Foot of Fuel Rod, lb/ft 0.364 l Rod Cluster Control Assemblies Neutron Absorber Ag-In-Cd 80%, 15%, 5% ! Composition - Diameter, in 0.341 Density, lbs/in.3 0.367 Cladding Material Type 304, Cold Worked Stainless Steel Clad Thickness, in. 0.0185 Number of Clusters Full Length 53 i Part Length 8 Number of Absorber Rods per Cluster 24 Full Length Assembly Weight (dry), lb. 157 4,3-64 DECEl6ER, 1973 6
. . . , . . -. . . . , . -- n
1 i TABLE 4.3-1 (Continued) Sheet 3 of 3 ( , ( ,) REACTOR CORE DESCRIPTION ~ (FourLoop,FirstCycle) 1 Burnable Poison Rods (First Core) Number 1520 l Material Borosilicate Glass Outside Diameter, in. 0.381 Inner Tube, 0.D. , in. 0.1805 Clad Material Stainless Steel l Inner Tube Material Stainless Steel Boron Loading (w/o B23 0 in glass rop-,, _12.5 I Weight of Boron - 10 per foot of rodrlb/ftc0.000419 - Initial Reactivity Worth, %Ap 7.63 (hot), s5.5.(cold) l Excess Reactivity l Maximum Fuel Assembly k, (Cold, Clean, Unborated Water) 1.39 Maximum Core Reactivity (Cold, Zero ('~' ) l Power, Beginning of Cycle) 1.222 l l l l ( \
'LJ 4.3-65 DECEMBER, 1973 l
l
l TABLE 4.3-2 , NUCLEAR DESIGN PARAMETERS l (Four Loop, First Cycle)
^
i Core Average Linear Power, kW/ft. including : densification effects 5.45 1 ; i
. Total Heat Flux Hot Channel Factor, Fg including densification effects 2.50 )
Nuclear Enthalpy Rise Hot Channel Factor, N F aH 1.55 Reactivity Coefficients Doppler Coefficient See Figures 4.3-27 and 4.3-28 Moderator Temperature Coefficient at ,p El OperatingConditions,pcm/'F[a] O to -40 Boron Coefficient in Primary Coolant, pcm/ ppm [a] -16 to -8 Rodded Moderator Density Coefficient 5 at Operating Conditions, pcm/gm/cc < +0.43 x 10 Delayed Neutron Fraction and Lifetime B BOL,(E0L) 0.0075,(0.0044) eff t*, BOL, (E0L), " sec 19.4 (18.1) l l Control Rod Worths l l' Rod Requirements See Table 4.3-3
- i. Maximum Bank Worth, pcm < 2000 Maximum Ejected Rod Worth See Chapter 15 1
[a]
.,_. Note: 1 pcm s (percent mille) 10-5 ap where op is calculated
( from two statepoint values of k,ff by in (k2 /*1)* D EEMBER, W 3 4.3-66 l
r I TABLE 4.3-2 (Continued) Sheet 2 of 2 NUCLEAR DESIGN PARAMETERS e (FourLoop,First. Cycle) i Boron Concentrations Refueling, k,ff = 0.9, Cold, Rod Cluster ' Control Assemblies In 1718 Zero Power, K,ff = 0.99, Cold, Rod Cluster Control Assemblies Out 1447 ; I Zero Power, keff = 0.99. Hot, Rod Cluster Control Assemblies ~Out 1420 Full Power, No Xenon, k,ff = 1.0, Hot, ' L: Rod Cluster Control Assemblies Out 1190 Full Power, Equilibrium Xenon, k,ff = 1.0, Hot, Rod Cluster Control. Assemblies Out 894 Reduction With Fuel Burnup i. First Cycle, ppm /GWD/MTU bb3- See Figure 4.3-3 Reload Cycle, ppm /GWD/MTU s 100 [b] l Gigawatt Day (GWD) = 1000 Megawatt Day (1000 MWD). During the first cycle, fixed burnable poison rods are present which significantly reduce the boron depletion rate compared to reload cycles. The values quoted are representative of averages only. l 4.3-67 DECEMBER,1973
~
b b _ b(, L TABLE 4.3-3 REACTIVITY REQUIREMENTS FOR R0D CLUSTER CONTROL ASSENLIES (Four Loop) - Reactivity Effects, Beginning of Life End of Life End of Life ' percent (First Cycle) (First Cycle) (Equilibrium Cycle) (Preliminary)
- 1. Control requirements Fuel temperature (Doppler), %Ap 1.37 1.08 1.30 Moderator temperature, %Ap .10 1.06 1.25 Void, %Ao .01 .05 .05 Redistribution, %ap .50 .85 .85 Rod Insertion Allowance, %Ap .50 .50 .50
- 2. Total Control, %ap 2.48 3.54 3.95 r
- 3. Estimated Rod Cluster Control Assembly Worth (53 Rods)
- a. All full length assemblies I i Y' inserted, %Ap 8.26 8.13 7.30
$ b. All but one (highest worth) ,
assemblies inserted, %Ap 6.96 6.80 6.20
- 4. Estimated Rod Cluster Control Assembly credit with 10 percent adjustment to accomodate uncertainties (3b - 10 percent),%Ao 6.26 6.12 5.58
- 5. Shutdown margin available (4-2),
%ap 3.98 2.58 1.63g,)
e b ili 9
]
[a]The design basis minimum shutdown is 1.6%. 4.3-68 - - . _ - n_ . _ _ - ..-._-._..___-:-.__-x__--.__- . , - - - . --.._.__. _ -------- ----____ _ _--- - - - - .-_-____ - - - --__.--._ -- _ . - - - . _ _ . _ - - _ _ _ _ - - _ - - _ _ _ _ _ .
TABLE 4.3-4 O t] AXIAL STABILITY INDEX PWR CORE WITH A 12-FT HEIGHT Burnup F Z B Stability Index (hr-I) (MWD /T) (ppm) Exp Calc 1550 1.34 1065 - 0.041 -0.032 7700 1.27 700 -0.014 - 0.006 1 Difference: +0.027 +0.026
,~ \
v i "% 4.3-69 DECEMBER, 1973
_ . - ~ . . _Sm l- - ,. TABLE 4.3-5 2 TYPICAL NEUTRON FLUX LEVELS (n/cm -sec) AT FULL POWER (Four Loop) E > l.0 MeV 5.53 Kev < E .625 ev i E- E < .625 ev i 1.0 Mev < 5.53 Key (ny)0-CORE CENTER 6.51 x 10 13 .1.12 x 10 I4' 8.50 x 10 13 3.00 x 10 13 CORE' OUTER RADIUS 3.23 x 10 13 5.74 x 10 13 13 i AT MIDHEIGHT 4.63 x~10 8.60 x 10 12 CORE TOP, ON AXIS 1.53 x 10 13 2.42 x 10 13 2.10 x 10 13 1.63 x 10 13 CORE BOTTOM, ON AXIS 2.36 x 10 I3 3.94 x 10 13 3.50 x 10 13 1.46 x 10 13
.o S PRESSURE VESSEL
= INNER WALL, 2.77 x 10 10 5.75 x 10 10 .
6.03 x 10 10 8.38 x 10 10 AZIMUTHAL PEAK, CORE MIDHEIGHT I i
~
. ._ _ _ _ ~ . _ . . __ . _ _ . . , , _ _ . . . . . . . . . - . _ , _ _ _ . . _ . ..
I TABLE 4.3-6
.I COMPARISON OF MEASURED AND CALCULATED D0PPLER DEFECTS 1 :
I Plant Fuel Type Core Burnup Measured (pcm)(I) Calculated (pcm) (MWD /MTU) I 1 l 1 Air-filled 1800 1700 1710 - 2 Air-filled 7700 1300 1440 3 Air and 8460 1200 1210 helium filled (1) 5 pcm = 10 x in k 3 /k 2 i l l l t
.l i
l.
.r . :
4.3-71 DECEMBER, 1973 .c
=
TABLE 4.3-7 BENCHMARK CRITICAL EXPERIMENTS-1 1
' Description of(I) No. of . LEOPARD k,ff Using Experiments- Experiments Experimental Bucklings l
l U0 2 Al Clad , 14 1.0012 SS Clad 19 0.9963 l Borated H 2O 7 0.9989 Total 40 0.9985 U-Metal Al Clad 41 0.9995 Unclad-. 20 0.9990 Total 61 0.9993 Grand Total 101 0.9990 l 4-- [ 1 l (1) Reported in Reference 24. 1 1 l l l l O 4.3-72 DECEPEER,1973 i 1
. - . . , . __ _ . _ . . _ , _. .,- _ , . _ . ~ - . . . _ . . . _ .
TABLE 4.3-8 SAXTON CORE II IS0 TOPICS R0D MY, AXIAL ZONE 6 1 Atom Ratio Measured (I) 2a Precision (%) LEOPARD Calculation U-234/U 4.65 x 10-5 29 4.60 x 10-5 U-235/U 5.74 x 10-3 0.9 5.73 x 10-3 U-236/U 3.55 x 10-4 15.6 3.74 x 10-4 U-238/U 0.99386 10.01 0.99385 Pu-238/Pu 1.32 x 10-3 2.3 1.222 x 10-3 Pu-239/Pu 0.73971 10.03 0.74497 Pu-240/Pu 0.19302 10.2 0.19102 Pu-241/Pu 6.014 x 10-2 0.3 5.74 x 10-2 Pu-242/Pu 5.81 x 10-3 0.9 5.38 x 10-3 Pu/U I2) 5.938 x 10-2 0.7 5.970 x 10-2 Np-237/U-238 1.14 x 10-4 11 5 0.86 x 10-4 Am-241/Pu-239 1.23 x 10-2 15 1.08 x 10-2 Cm-242/Pu-239 1.05 x 10~4 +10 1.11 x 10~4 Cm-244/Pu-239 1.09 x 10~4 120 0.98 x 10~4 J (1) Reporteo in Reference 25 (2) Weight ratio. O 4.3-73 DECEMBER, 1973
r t l TABLE 4.3-9 l CRITICAL BORON CONCENTRATIONS, HZP, BOL j 1 : I Plant Type Measured Calculated l 2-Loop, 121 Assemblies 10 foot core 1583 1589 2-Loop, 121 Assemblies ] 12 foot core 1625 1624 2-Loop,121 Assemblies 12 foot core 1517 1517 3-Loop, 157. Assemblies 12 foot core 1169 1161 O l I i. O 4.3-74 DECEMBER, 1973 y
.- .- _ _ . ..- _ - _ =._ - . - - . - .- - - . . . _ .. - .- _.- -
- j. TABLE 4.3-10
,r-C019ARISON OF MEASURED AND CALCULATED R0D WORTH i 1 l
j ' 2-Loop Plant 121 Assemblies, 10 foot core Measured (pcm) Calculated (pcm) Group 8 1885 1893 Group A 1530 1649 l Shutdown Group 3050 2917 ! 1 1 ESADA Critical II) , 0.69" Pitch, 2 w/o Pu0 , 8% Pu 240 , 2 9JontrolRods 1 6.21" rod separation 2250 2250 2.07" rod separation 4220 4160 1.38" rod' separation 4100 4010 1 l (1) Reported in Reference 26. i l l l O 4.3-75 DECE M ER, 1973 1
! TABLE 4.3-11 COMPARIS0N OF MEASURED AND CALCULATED MODERATOR j l- C0EFFICIENTS AT HZP, BOL j 1
~ ,
1 Plant Type /- Measured a I) iso Calculated ag3, Control Bank Configuration (pcm/ F) (pcm/*F) { I~ 3-Loop.157 Assemb1ies, i 12 foot core f D at 160 steps -0.50 -0.50 ) D in, C at 190 steps -3.01 -2.75 f D in, C at 28 steps -7.67 -7.02 ! B, C, and D in - -5.16 -4.45 ! 2-Loop, 121 Assemblies. 12 foot core ' D at 180 steps +0.85 +1.02 D in C at 180 steps -2.40 -1.90 C and D in, B at 165 steps -4.40 -5.58 B, C, and D in. A at 174 steps -8.70 -8.12 l (1)' Isothermal coefficients, which include the Doppler effect in the fuel. k 5 a = 10 1n / AT*F in l 1 l O 4.3-76 DECEMBER, 1973 1
TABLE 4.3-1 l REACTOR CORE D_ES_CRIPTION l
'. (Three Loop, First Cycle) I l
Active Core Equivalent Diameter, in. 119.7 Core Avg. Active Fuel Hgt. , First Core, in.143.7 I Height-to-Diameter Ratio 1.20 2 Total Cross-Section Area, ft 78.14 H 0/U Molecular Ratio, Lattice (Cold) 3.43 2 Reflector Thickness and Composition Top - Water plus Steel, in. s10 Bottom - Water plus Steel, in. s10 l Side - Water plus Steel, in, s15 Fuel Asseinblies Number - 157
^
Rod Array 17 x 17 x Rods per Assembly 264 Rod Pitch, in. 0.496 l Overall Transverse Dimensions, in. 8.426 x 8.426 Fuel Weight (as U0 2
),1bs. 181,205 Zircaloy Weight,1bs. 38,230 Number of Grids per Assembly 7-R type 1 Composition of grids INC718 Weight of Grids (Effective in Core) lbs 1336 Number of Guide Thimbles per Assembly 24 l
Composition of Guide Thimbles Zircaloy 4 Diameter of Guide Thimbles (upper 0.450 I.D. x 0.482 0.D. part),in. l Diameter of Guide Thimbles (lower part),in. 0.397 1.D. x 0.429 0.D. Diameter of Instrument Guide Thimbles, in. 0.450 1.D. x 0.482 0.D. O
) 4.3-77 DECEMBER,1973
)
3 E ! l-
' TABLE 4.3-1 (Continued) ' I ,p REACTOR CORE DESCRIPTION l
L l 1 (Three Loop, First Cycle) Il: i Fuel Rods _ .
^
Number 41,448 i Outside Diameter,. in. 0.374 , Diameter Gap, in. 0.0065 l Clad Thickness, In. 0.0225 l i' Clad Material Zircaloy-4 ' i l' ! l' Fuel Pellets i Material -UO Sintered 2 l I Density (percent of Theoretical) 95 Fuel Enrichments w/o l Region 1 2.10 Region 2 2.60 ) Region 3 3.10 Diameter, in. 0.3225 Length, in. 0.530 Mass of U02 per Foot of Fuel Rod, lb/ft 0.364 l Rod Cluster Control Assemblies Neutron Absorber Ag-In-Cd Composition 80%,15%, 5% - !: Diameter, in 0.341 Density,lbs/in.3 0.367 Cladding Material Type 304, Cold Worked Stainless Steel Clad Thickness, in. 0.0185 i Number of Clusters Full Length 48 ! Part length 5 l Number of Absorber Rods per Cluster 24 ] l Full Length Assembly Weight (dry), 1b. 157 l [ 4.3-78 DECEPBER,1973 1 l i
. . _ , ._ , - . i
i L l TABLE 4.3-1 (Continued)
), ] REACTOR CORE DESCRIPTION _
! (Three Loop, First Cycle) ! Burnable Poison Rods (First Core) '! i Number 1072 Material Borosilicate Glass f' Outside Diameter, in. 0.381 i Inner Tube, 0.D., in. 0.1805 ! Clad Material . Stainless Steel l Inner Tube Material Stainless Steel : ' 1 Boron Loading (w/o 23 B 0 in glass rod) 12.5 i Weight of Boron - 10 per foot of rod, lb/ft .000419 l < Initial. Reactivity Worth, %Ap 7.0 (hot),s5.5(cold) Excess Reactivity Maximum Fuel Assembly k ,(Cold, Clean, l UnboratedWater) <1.6 I Maximum Core Reactivity (Cold, Zero Power,BeginningofCycle) 1.25 l I I i
, i i
.4.3-79 i DECEMBER, 1973 I.
i i !
~!.
TABLE 4.3-2
'O V
NUCLEAR DESIGN PAP.AMETERS (Three Loop, First Cycle) L Core Average Linear Power, kW/ft, including densification effects 5.20 I l l Total Heat Flux Hot Channel Factor, including densification effects 2.50 1 l Nuclear Enthalpy Rise Hot Channel Factor, F H 1.55 Reactivity Coefficients Doppler Coefficient See Figures 4.3-27 and 4.3-28 l Moderator Temperature Coefficient at l Operating Conditions, pcm/*FE*3 0 to -40 Boron Coefficient in Primary Coolant, pcm/ ppm [a] -16 to -8
-(p) Rodded Moderator Density Coefficient i 5 j at Operating Conditions, pcm/gm/cc < +0.43 x 10 Delayed Neutron Fraction and Lifetime BeffBOL,(EOL) 0.0075 (0.0044) 1*, BOL, (EOL), u see 19.4 (18.1) l l
Control Rod Worths i Rod Requirements See Table 4.3-3 l Maximum Bank Worth, pcm < 2300 Maximum Ejected Rod Worth See Chapter 15 Laj Note: 1 pcm E (percentmille)10-5 op where ao is calculated from two statepoint values of k,ff by in (k2 /k)). ( V) 4.3-80 DECEMBER, 1973 l
l l TABLE 4.3-2 (Continued) , l NUCLEAR DESIGN PARAMETERS (Three Loop, First Cycle)
.. Boron Concentrations Refueling, k,ff <_0.9, Cold, Rod Cluster t Control Assemblies In 2000 Zero Power, K,ff = 0.99, Cold, Rod i
Cluster Control Assemblies Out 1455 ! Zero Power, k,ff = 0.99, Hot, Rod Cluster Control Assemblies Out 1430 Full Power, No Xenon, keff = 1.0, Hot, Rod Cluster Control Assemblies Out 1198 i Full Power, Equilibrium Xenon, k,ff = 1.0, I Hot, Rod Cluster Control Assemblies'Out 905 Reduction With Fuel Burnup First Cycle, ppm /GWD/MTU EN See Figure 4.3-3
- Reload Cycle, ppm /GWD/MTU s 100 l
[b] Gigawatt Day (GWD) = 1000 Megawatt Day (1000 MWD). During the first cycle, fixed burnable poison rods are present which significantly reduce the boron depletion rate compared to reload cycles. The values quoted ; are representative of averages only. l l 4.3-81 . DECEMBER, 1973 l
r} O b U .f* d TABLE 4.3-3 REACTIVITY REQUIREENTS FOR ROD CLUSTER CONTROL ASSEMBLIES (Three Loop) Reactivity Effects, ~Beginning of Life' End of Life End of Life percent (First Cycle) (First Cycle) (Equilibrium Cycle) (Preliminary) ,
- 1. Control requirements Fuel temperature (Doppler), %ap 1.30 1.20 1.30 Moderator temperature, %ap .5 1.17 1.25 Void, %ap .05 .05 .05 Redistribution, %ap .50 .85 .85 Rod Insertion Allowance, %ap .50 .50 .50
- 2. Total Control, %ap 2.85 3.77 3.95
- 3. Estimated Rod Cluster Control
- Assembly Worth (48 Rods) la a. All full length assemblies I lo inserted, %ap 9.88 9.57 8.50
- b. All but one (highest worth) assemblies inserted, %ap 7.85 7.81 6.64
- 4. Estimated Rod Cluster Control Assembly credit with 10 percent adjustment to accomodate uncertainties (3b - 10 percent),%ap 7.06 7.03 5.97
- 5. Shutdown margin available (4-2),
g %Ap 4.21 3.26 2.02[,3 R iE h [aj The design basis minimum shutdown is 1.77%. w"
O O O Table 4.3-5 TYPICAL NEUTRON FLUX LEVELS (n/cm -sec)__AT FULL POWER (Three Loop) E > 1.0 Mev 5.53 Kev < E .625 ev < E E < .625 ev -
< 1.0 Mev
< 5.53 Key (nv), i CORE CENTER 6.51 x 10 13 1.12 x 10 I4 8.50 x 10 I3 3.00 x 10 13 i ,
O CORE OUTER RADIUS I AT MIDHEIGHT 3.23 x 10 13 5.74 x 10 13 4.63 x 10 13 8.60 x 10 12
?
CORE TOP, 13 1.53 x 10 I3 13 ON AXIS 2.42 x 10 13 2.10 x 10 1.63 x'10 i R CNE mum, 13 I3 13 13 ON AXIS 2.36 x 10 3.94 x 10 3.50 x 10 1.46 x 10 PRESSURE VESSEL 10 10 10 b INNER WALL, 2.77 x 10 10 5.75 x 10 6.03 x 10 8.38 x 10 l M HE G T , I i
.i
6274-39 R P N M L K J H G F E D C B A 180o k l 2 GE NW E M RMN a MMM M M E E MM l - WMEM E E EMME ' s MM M E E E E M2 6 h h goo g 7 h M h M hM h h hM hM hk 95 & , $6R _ N 270o
#M M M 66 9 . M M M M 9 E E M iO M M M M M M M N O " M8 M M M E M ME i2 MMMM M ' M MMMM is E MM M M M MMM lu MMM M M MMM 15 h h kj h h h hh Oo
__ FIRST CORE RELOAD CORE l REGION I ONCE OR TWICE BURNED FUEL REGION 2 ONCE OR TWICE BURNED FUEL h REGION 3 FRESH FUEL l Figure 4.3-1 Fuel loading Armngement
, (Four Loop)
I 4.3-84 1 l
i l 6694-i O l, 9 0 8 - _ ,,g E E E g 7 -
-8 E 3 %
m 3 W 6
-\2 m
- t o .
o a " " - ~
-"% o 5 -
# - -16 m g / u-235 1
' d 1
=4
_ / - o l l eu2as, / { g3 [ u-238
-24 5 p i
/
M
~
k l
'S 2 -
-28 E I E 5o Pu241 I -
""~ -32 0
L ~~* ~ l l I I I I I
-36 0 4 8 12 16 20 24 28 32 36 40 BURf4UP,(GWD/MTU)
Figure 4.3-2 Production and Consumption of Higher Isotopes (Three and Four Loops) O 4.3-85
O O O k 2000 [ MOTE: HOT. FULL POWER, RODS OUT e g 1600 L
-N N WITHOUT BURMABLE PolSON
< N N %
5 1200 M N o
, N BURMUP DIFFERENCE y E O N
1565 MWD /MTU
= 5 N
", N
- 1400 - N 3 WITH BURMABLE P0lSONS N t N 5
o 1 I I I I I Ye 0 2000 16000 6000 8000 10000 12000 ii&000 16000 18000
. CORE AVERAGE BURMUP (MWD /MTU) i m
Figure 4.3-3 Baron Concentration versus First Cycle Burnup with and without Burnoble [o Poison Rods (4 Loop) 4
6274-4 O g 3 O E g g E E S g E O E O E E O E O E E E E E D E E O E O E O E E O E O E E S E E E E E O B E 20 BP'S 16 BP'S y O E O S O B E O E O E O E E O E O O O S S O E E O O E D E O E E O O O O ; O S O B O D O O 0 0 12 BP'S C0F,E CENTER 9 BP'S n g E E E g 8 O E O E E O O E O O O O O O O O O O 10 BP'S Figure 4. 3-4 Burnable Poison Rod Arrangement Within on Assembly (Four Loop) 4.3-87
6274-36 O R P N M L K J H G F E D C B A 180 l 10 10 10 2 9 12 20 I,l 12 9 3 9 20 16 16 20 9 f6 4 20 20 16 16 20 20 5 12 20 16 16 is 20 12 6 10 16 16 20 20 16 16 10 7 20 16 20 20 20 is 20 8 to is 16 20 20 16 16 to 90 270 9 20 is 20 20 20 16 20 10 10 16 16 20 20 16 16 10 ll 12 20 16 16 16 20 12 12 20 20 is is 20 20 13 9 20 16 16 20 9 14 9 20 12 9 12 l83 15 i0 10 10 0 NUMBER INDICATES NUMBER OF BURNABLE POISON RODS S INDICATES SOURCE ROD Figure 4. 3-5 Burnable Poison loading Pattern (Four Loop) 4.3-88
.. . - . . . - . . . - - - . . . . . . - - - - . - - . . - - . . . - - _ - . . - - ~ . . . - - . - . _ . . . . . . - . 6274-22 O 1 H G F E D C 8 A r ( l.063--- - - - -- - -- -- - --- - - - - -
-(
8 l 9 1.003 1.090 l 1.128 1.054 1.165 ; 10 i g, 1.152 1.192 1.152 1.179 l i . I 1.219 1.164 1.182 f.068 1.257 12 1.128 1.160 1.097 1.093 .931 .995 13 l 1.032 1.039 .996 .992 .840 .471
,q l
,5 .725 .M .669 .566 CALCULATED F $ g = 1.36 KEY:
- k. VALUE REPRESENTS ASSEMBLY RELATIVE POWER Figure 4. 3-6, Normalized Power Density Distribution Near Beginning of Life, Unrodded Core, Hor Full Power, No Xenon (FourLoop) 4.3-89 4
., - c_. , c _ m .. - . , , .
6274-24 I k H G F E D C B A-( l.097-- - - - - --- --- - -- --- --- ( 8 l 9 1.029 1.120 I 1.154 1.074 1.186 10 i i 1.167 1.210 1.163 1.187 l l.225 1.'171 1.190 1.065 1.241 1 12 t 1.127 1.163 1.094 1.088 .915 .969
,3 I
i l ,g 1.026 I.026 .991 .976 .823 .463 l l
,s .7i5 .779 .ee2 .560 cAtcuti1Eo ris . i.34 KEY:
{ VALUE REPRESENTS ASSEMBLY l RELATIVE POWER i Figure 4. 3-7. Normalized Power Density Distribution Near Beginni of Life ! Unrodded Core, Hot Full Power, Equilibrium Xenon (7our Loop) i i ' 4.3-90
6694-73 O 4 l i H G F E D C A B l 1 l q__ o,339 ~. __ _ . __ __ _ _
.__q j 8 :
; L -
9 0.957 1.074 i T 10 1.122 1.050 1.184 I lI l.100 1.177 1.163 1.166 1
\ l 12 i.001 1.122 1.188 1.025 1.028 !
13 1.105 1.173 1.128 1.092 0.885 0.965 l l L4 1.059 1.102 1.057 1.036 0.854 0.471 15 0.788 0.86: 0.726 0.607 CALCULATED F ( :1.37 i KEY: k VALUE REPRESENTS ASSEMBLY RELATIVE F0WER Figure 4.3-8 Normalized Power Density Distribution Near Beginning of 1.ife, Group D 30% Inserted, Hot Full Power, Equilibrium Xenon AMENDMENT 1 4.3-91
i 6274-26 O t H G F E D C B A ( 0.945- - - -- - - - - -- - - -- - -- -- - - -
-(
8 _ l 1.043 1.165 9 I l t 1.197 1.116 1.212 10 1 l gg 1.127 1.187 1.122 1.115 0.981 1.083 1.012 0.963 0.976 12 O ! 1.171 1.093 1.077 0.880 0.%8 13 1.116 l I 1.127 1.126 1.082 1.049 0.865 0.483 gg l
! e i
0.821 0.892 0.754 0.629 CALCULATED F"A H = 1,39 15 KEY:
- i. VALUE REPRESENTS ASSEMBLY
! RELATIVE POWER Figure 4.3-9. Normalized Power Density Distribution Near Beginning of Life, Group D 35% Inserted
- O Plus PL Inserted. Hot Full Power, Equilibrium Xenon (Four Loop) 4.3-92
. . _ . . - _ ~ . _ .~ __._ ._ . _ _._.. _ .. _ ... ._... _ _ _ .m.- _ - ...__._ _ ..._ . - . . . . . . _ _ _ _ _ . .
i 6274-23 j i i O l k I l 1 H G F E D C B A l l ( l.154- - - --- --- -- - -- - - -- ---
---(
8 1.189 1.152 9 I J 1.153 1.189 1.153 l 10 1 I 1 1.212 1.152 1.208 1.143
,i l l l
1.137 1.191 1.128 1.149 1.203 O i l.143 1.087 1.123 1.048 .993 .945 13 l 3
.972 1.039 .940 .962 .797 .483 gy
, 1 l
N
.707 .739 657 .542 CALCULATED F =1.30 15 Ay l
KEY: k VALUE REPRESENTS ASSEMBLY RELATIVE POWER Figure 4. 3-.10 Normalized Power Density Distribution Near Middle of Life, p) q, Unrodded Core, Hot Full Power, Equilibrium Xenon (Four Loop) 4.3-93
. . . . . - . - . ... . . . - . - . ~ _ _ . - . ~ - .
6274-27 l 1 [ H G F E D c 8 A ( l.054- - - - - - --- - -- - -- - -- --- - -
-( -
8 j g I.114 1.052 I i 10 1.Mi 1.113 1. M I
,, 1.123 1.054 1.122 1.074
, i l 12 1.060 1.132 1.069 1.144 1.186 lO i 1.134 1.086 1.132 1.055 1.070 1.006 13 l I l gg 1.006 1.111 .981 1.029 .859 .565 l l _ i l
.792 .806 .746 .608 l ,3 cAtcutATE0ria . i.25 KEY:
( VALUE REPRESENTS ASSEMBLY RELATIVE POWER l t Figure 4. 3-11. Normalized Power Density Distribution Near End of Life, l Unrodded Core, Hot Full Power, Equilibrium Xenon (Four Loop) 4.3-94
- . . . . . . . = _ . - - _. . .._. ._
6274-6 v i 1.02 l.02 1.03 1.02 1.05 1.08 1.02 1.06 1.12 f.03 1.08 1. 14 1. 17 1. 16
~
1.03 1. 10 1. 17 8.1 B l.03 1.08 1. 14 1. 14 1. 16 1. 18 1. 17 1.03 1.08 1. 14 1. 14 1. 16 1. 19 1. 17 l. 7 1.03 1. 10 1. 16 1.18 1. 19 1,20 [ i.03 1.09 1. 14 1.14 1.16 1. 19 1.17 1. 17 1.20 1. 18 k]/ 1.03 1.09 1.15 1.15 1.16 1. 19 1. 17 1. 17 1.20 1. 18 1. 18 1.04 1. I i 1. 18 1.18 l. 19 1. 19 1.20 1.20 1.04 1.09 1. 15 1.18 1. 17 1. 19 1. 17 1.17 1.19 1. 17 1. 17 1.20 1. 18 1.04 1.08 1. 14 1.18 1. 18 1. 16 1. 16 1. 18 1. 16 1. 16 1. 19 1. 19 1.04 1.06 1. 10 1.14 1. 16 1.16 1.16 1. 16 1. 16 1. 17 1.15 1. I I l.04 1.05 1.07 1.08 1. 10 1.12 1. 10 1. 10 8.12 1. 10 1. I i 1.12 1.1 I l.09 1.08 1.07 1.04 1.04 1.04 1.04 1.05 1.05 f.05 1.05 1.05 1.05 1.05 1.06 1.06 1.06 1.06 1.06 1.06 l Figure 4.3-12. Rodwise Power Distribution in a Typical Assembly (Assembly G-9) Near Beginning
'A of Life, Hot Full Power, Equilibrium
- Xenon. Unrodded Core (Four Loop) 4.3-95 l
t
i 6274-7 1
\,
0.98 0.98 0.99 0.98 1.00 1.03 , 0.99 1.02 1.07 1.00 1.03 1.09 1.ll 1.09 i 1.01 1.06 l. ll I. ll I.00 1.04 1.08 1.07 1.08 1.10 1.08 1.00 1.04 1.08 1.07 1.08 1.10 1.08 1.~0 8 j l.01 1.06 1.09 1.10 1.10 1.10 [ 1.00 1.04 1.08 1.07 1.08 1.10 1.08 1.08 1.10 1.08 l (, ) 1.00 1.04 1.08 1.07 1.08 1.10 1.08 1.08 1.10 1.08 1.08 1.01 1.06 1.11 1.11 1.10 1.10 1.10 1.10 t.00 1.03 1.09 l.1I l.09 1.II l.08 1.08 1.10 1.08 1.08 1.Ii 1.09 0.99 1.02 1.07 1. I i 1.Ii 1.07 1.07 1.09 1.07 1.07 1. I 1 1.1 I 0.98 1.00 1.03 1.07 1.09 1.08 1.08 1.08 1.08 1.09 1.07 1.03 0.98 0.99 1.00 1.02 1.03 1.06 1.04 1.04 1.06 1.04 1.04 1.06 1.03 1.02 1.00 0.98 0.98 0.98 0.98 0.99 1.00 1.01 1.00 1.00 1.01 1.00 1.00 1.01 1.00 0.99 0.98 0.98 0.98 l Figure 4.3-13. Rodwise Power Distribution in a Typical Assembly (G-9) Near End of Life, Hot Fuil Power, Equilibrium Xenon. Unrodded Core (FourLoop) 4.3-96
O O -O B. PL' S IN TOP C4 IN 30% AXl AL OFFSET = -20.5% 1.5 - C. PL' S IN BOTTOM C4 0UT AXIAL OFFSET = 9.8% f5 1.0 - I o A. PLS IN HIDDLE
$ Ct; IN 30% !
,A I AXI AL OFFSET = -30.57, y 0.5 fi PL PART LENGTH RODS C4 : CONTROLLING GROUP OF RODS t I I I I I I I I I I 0 ! 0 10 20 30 % 50 60 70 80 90 100 ; PERCENT OF ACTIVE CORE HEIGHT FRON BOTTOM m n > 2 j Figure 4.3-14. (Typical Axial Three and FourPbwer Loop) 9ppes Occuring at Start of Life j ; I _ _ _ _ J
O O O B. PL'S IN BOTTOM C4 IN 30% I.5 - AXI AL OFFSET = +18.9% 30,000 A. PL'S IN TOP C4 IM 30% , t AXI AL OFFSET s-10.7% w - t g 1 .0 20,000 g u &
- x
- E C. PL'S IN MlDOLE S y E C4 IM 10% b y AXI AL OFFSET = +35.3% a o
s% % w g i 0.5 %*- . , 10,000 i LEGEND: N NORMALIZED POWER
= = = FUEL DEPLETION N
\
N PL:PART LENGTH RODS g C4 = CONTROLLING GROUP OF RODS N% I I I I I I I I I O O O 10 20 30 40 50 60 70 80 90 100 PERCENT OF ACTIVE CORE HEIGHT FROM BOTTOM O Figure 4. 3-15. Typical Axial Power Shapes Occuring at Middle of Life ? (Three and Four Loop) r
O O O . 1.5 - A. PL'S IN TOP C. PL'S 25% WITHDRAWM - 30,000 C4 OUT C4 OUT AXI AL OFFSET =-21.7% AXIAL OFFSEr= +ll.4% 1.0 - - 20,000 2
~
f !
= # '% # 5 8 / -" '% ~
w O 8. PL'S IN HIDDLE $ i h C4 OUT d AXI AL OFFSET = 0.0% \ E 1 5 / \ lo,000 g 0.5 , , F LEGEND: \
. / NORMALIZED POWER
\ ,
' , / -- FUEL DEPLETION !
i
\
r
\
PL = PART LENGTH RODS 1 C4 = CONTROLLING GROUP OF RODS l l l l l 0 O 0 10 20 30 16 0 50 60 70 80 90 100 PERCENT ACTIVE CORE HEIGHT FROM BOTTOM S Figure 4. 3-16. Typical Axial Power Sha y (Three and Four Loop) pes Occuring at End of Life = i
-3 4
6274-53 3 O -
- s=
/,/
//
,/
// oG s' *5
/ Eit
/' / i$
oo
/ // _
a uo l lI l M 1
$.cE}
c3 1Il = a I a="z
- s' l's
'E '
l aGO $$ .2 I Wo5E5" QI l 2""5"d- tE L W: m: 5
- g-I t
i
*02WZf no rm a 5
m se i 8 eie
- 3. g-t1 ,-
1 1 W o JE a 1 u x c-l tO \ g 1 l l l l
]E?
e e3 s i .e u
' 212 kE en -
\\ \ O I
U t
\\
\s s
\
s b . Y E s a N N,
.9 u.
N sN N N N
\
l l Wm 9 o 8R
- _9 9 o
83M0d 3A11V'138 1VlXV \ 4 3-100
i 6937-89 O 1 FREQUENCY OF AXlALFREQUENCY SIZE FREQUENCY GAPS PER DISTRIBUTION OF DISTRIBUTION OF ROD-F, GAPS-F g GAPS-F) KTG FROM DENSIFICATION SPlKE DUE TO SINGLE MODEL(MAX.GAPSIZE GAPS-S, DRAW COMPUTER ATGIVENHEIGHT) CODE MODE I MODE 2 EXPECTED VALUES q SEEN BY INCORE DETECTOR O' PROBADILITIES OF EXCEEDING GIVEN SPIKE SIZE FOR EACH AXlAL I R0D CENSUS 1r CONVOLUTION 1r POWER SPlKE FACTOR-S(Z) i Figure 4.3-18 Flow Chart for Determining Spike h41 (Three and Four Loop) , E W ENT 1 4.3-101
6937-63 82.0 10.0 -
---- is x i s 17 X 17 i
g 8.0 - E G S g 6.0 - u ~~ ~~ 5 # t ,a# h 4.0 -
,h' O '
/
/
/
/
/
2.0 - j 0 I l 1 0 1.0 2.0 3.0 4.0 GAPSIZE(INCHES) Figure 4.3-19 Predicted Power Spike Due to Single Nonflattened Gap in the Adjacent Fuel (Three and Four Loop) AMENDMENT 1 4.3-102
O O O . b I.06
/
p s 1.05 - --- - - -- 15 x i s p' ! 17 X 17 # f p
/
,- 1 a i.on -
x ,- c ! 3u m* 1.03 -
? 2 W E IMITI AL GE0 METRIC DENSITY = 94.5%
m O g [ g 1.02 - i E i 1.01 9 E I I I I I I I 9 H 1.00 0 20 40, 60 80 100 120 140 160 , AXIALPOSITION(INCHES) O Figure 4.3-20 Power Spike Factor as a Function of Axial Position y (Three and Four Loop) 0
. .. . . . . - . . . . .. . - - - . - . . - . . ~ . .
6937-140 , i F
- POWER 9
2.6 - 2.4 - 1 O* & 2.2 - X WITH PL RODS X O NO PL RODS l X 2.0 - X Oo 1.8 - CO ~ O l 9 XMX X O i,g _ l.2 - i
- X i 1.0 -
J l 0.8 - 0.6 - 0.4 - 0.2 - I I l
-15 -10 -5 0 5 10 FLUX DIFFERENCE (AI)
- Figure 4.3-21 Peaking Factor Limits for Normal Operation at Beginning of
( First Cycle (Three and Four Loop) AMENDMENT 1 4.3-104 l 4 n. , n
6937-212 O FQ
- POWER 2.6 2.4 X 2.2 - xN x xx l.8h X x x 1.6 r--
O X l.49 - O p O *xX x xX *='2 x I- r-- .
.6 X WITH PL RODS
.6 O M0 PL RODS
.4
.2 I I I I I I
-15 -10 -5 0 5 10 15 FLUX DIFFERENCE (AI)
Figure 4.3-22 Peaking Factor umits For Normal Operation at End of Fint Cycle O' (Three and Four Loop) NDENT 1 4.3-105
,.. .. .= _ .. . . _ . . . _ . . -.. - . -. . . . . . . ..
l ^
- 6937-142 KW/FT 25 O = UNCONTROLLED ROD BANK MOTION
- X = OPERATOR ERROR 4
4 ~ 20 X XX dX X X ftX"qige/* O X X X Oo O X XX O 10 i l l l l l l l l
-40 -30 -20 -10 0 10 20 30 40 FLUX DIFFERENCE (AI)
Figure 4,3-23 Peak Linear Power versus Flux Differences for Determination of Protection Set Points O (4 Loop) AMENDMENT 1 4.3-106
_ _ . . = _ _____ m. _ _ _ _ _ _ _ . _ _ . . _._ _ . ~ _ _ _ ._ - _ _ _ . . . 6274-72 O ' O.759 0.792
- 4. 3 f-1.217 f.224 0.6f 0.774 1.255 0.800 1.249 3.4f -0.51 1.229 1.229 1.225 1.248
-0.3f -0.9?
1.109 1.077 1.107 1.092
-0.2% 1.4T l.202 1.223 1.170 1.256
-2.74 2.71 0.523 1.217 1.221 1.217 0.548 1.203 1.233 1.210 u I. it 1.0I -0.67
_ .6E l.229 1.229 C. l.189
-3.31 1.220
-0.71 1.217 -*-- CALCULATED 1.211 *--- MEASURED
-0 5E d---- DIFFERENCE PEAKING FACTORS F =
1.5 7 F[g* 1. E FN : 2.07 LOCATED Ar 0 M-8 SOUTH Figure 4. 3-24. Comparison Between Calculated and Measured Relative Fuel Assembly Power Distribution ( (Three and Four Lc0p) 4.3-107
O O O l.4 1 BOL NO XENON , POWER AT 70% l.2 - BANK D 25% INSERTED 1.0 - CORE AVERAGE
@ FROM INCORE MAP ,
g A0 = 2%
, 0.8 -
CALCULATED
,a w A0 = 3%
y , 0.6 O m Q
=
0.4 0.2 -
. I I I I I I I I I o
0 10 20 30 40 50 60 70 80 90 100 j BOTTON PERCENT OF CORE HEIGHT TOP l e ! Figure 4.3 25 Comparison of Calculated and Measured Axial Shape 3 ' (Three and Four Loop) g i f
._. . - . . ~ . _ . . . _ . . . _. - . . _ ._ __. . _ _ . _ _ _ _ _ _ . . . . _ - . _
6274-54 x REACTOR I O REACTOR 4 O' O
--F 9 V REACTOR 2 9 REACTOR 3 la REACTOR 5 i
0 -- l O -- 3.0 m -- ! O _ -
" 7 p - - 2. 5 V g 8
- l y O __
E O g
*Vx ~~
mE a x 6x x 'O
"(m x* -k O
. o0 -- 2.0 o
" O e
.=* pf- g =*
" + 89 o ,. .
So g %% O g -- O 9 *
# __ pi.s I I I I I I I I I I I I I I I I 45 35 25 15 -10 -5 0 5 10 15 20 25 30 INCOREAXIALOFFSET(%)
l Figure 4. 3-2E Measured Values of Fg or f Full Power Rod Configurations l (Three and Four Loop) r 4.3-109 ;
6694-5 O _ -l.2 u. R 5 6 -l.4 - I [ -1.6 - E0L J 5 -l.8 - G f o -2.0 - BOL r E R g -2.2 - E 5 g -2.4 d 5 I I I l l l
-2.6 500 600 700 800 900 1000 l100 1200 EFFECTIVE FUEL TEMPERATURE Teff (OF)
Figure 4.3 27 D ler Temperature Coeffielent at BOL and EOL C le I (three and Four Loop) 4.3-110
6694-6 O 1 l l I l
-8 d 1 - E0L 5-
-10 -
l b 8e f t 5E g g -12
- n. ~
/
L BOL j l >. 3 a a. i \
'
- i
( , a: I -14 - l ds I& 8 I !
-16 0 20 40 60 80 100 POWER LEVEL (PERCENT OF FULL POWER) l
- Figure 4.3 28 D ler - Only Power Coefficient - BOL, E , Cycle I (Three and Four Loop) 4.3-111
d i
\
6694-7 lO i . I i ' - i 0 -- a I i x , E l, w g -500 - l 8 tii i o l
= l l
b - 5 O g "i E0L
& -1000 -
8 l BOL
-1500 ! ! !
0 20 40 60 80 100 POWER LEVEL (PERCENT OF FULL POWER) Figure 4.3-29 ler Only Power Defect - BOL, EOL, C le I(Three and Four LW O 4.3-112
.. - - . . = . . - . . . _ - - . . . . . . - - . . _ _ . . . _ - - . _ . . - . . - - . . . . ... ..
6694-8 O l l 20 l C O y 1 2000 PPM
? 1 10 -
x ,3,3 pp, [ l 0 \ i : l 2 , u l C 1000 PPM 500 PPM u
-10 -
y 0 PPM
?
O < 5 I
-20 -
4 l l W E O -30 - 5 o , E I I I ! I
-40 O 100 200 300 400 500 600 MODERATOR TEMPERATURE (OF) l 1
1 : Figure 4.3-:30 Moderator Temperatwo Cos%n!- BOL, Cycle i, No Rods (Three and Four Loop) 4.3-113
6694-9 O 20 C h 10 -- E E T 500 PPy - 0 - d - UNRODDED 5 " 5 -10 - o ppy _ t iZ u. 8 a o PPw
-20 t
RODDED b -30 8 a 5 y -40 - i I
-50 O 100 200 300 400 500 600 MODERATOR TEMPERATURE (OF)
Figure 4,3 3'. Moderator Temperature Coefficient - EOL, Cycle I (Three and Four Loop) 4.3-114
. . ~ _ . . - . . . . - - . _ _ . . . _ . - - _ .
-.l 6694-11 O
5 e 680F /' Y E 4000F J -5 - L?. 2 I w' ~ 5 -10 - h 5570F I w 8 f _ -15 - ac 5 586 F E ' g-20 8 4
$-25 -
E
-30 0 500 1000 1500 SOLUBLE BORON CONCENTRATION (PPM)
Figure 4.3 32 Moderator Temperature Coefficient cm a Function of Baron Concentration - BOL Cycle I, No Rods (Four Loop) 4.3-115
l 6694-12
'O :
l i O I C -5 - EL i 5 : e. J. -10 l 42 s -15 -- l 5 5 , C ' . b -20 - 8 ! ,O = R -25 - 5 ; t l
- 5
* -30 -
i 8 4
- 5 o -35 -
E I I I
-40 O 5000 10000 15000 20000 BURNUP (WD/MTU)
Figure 4.3 33 Hot Full Power Temperature Coeffielent During Cycle l for the Critical Baron Concentration (Three and Four Loop) O - 4.3-116
6694-13
-12 BOL 14 ;-
d w2 2 5 -16 - tu
E $
~
E e&. -18 m . O i;
* * -20 -
EOL I I I I
-22 O 20 40 60 80 100 i
POWER LEVEL (PERCENT OF FULL POWER) Figure 4.3-34 Total Power Coefficient - BOL, EOL, Cycle 1 (Three and Four Loop) O 4.3-117
- . . . ~ -.. ... ~. ..... . ..~_. ... . . . . . . - . . - - . - .
l I 6694-14 1 0 lO 1 i l 500 - 1 1 l 2 l E l
- 1000 -
l BOL ! A j a bo W e , a 1500 - i ! O. < O g H j E0L l l l l l 2000 - l l l l l l l l l l l l l l l 2500 l 0 20 14 0 60 80 100 POWER LEVEL (PERCENT FULL POWER) Figure 4.3-35 Total Power Defect 4 BOL, EOL, Cycle I (Three and Four Loop) !O 4.3-118 I
6274-46 l lO n A e A >A
- 5) @ @ !s~c]
e e e I s A e-- i
@ @ r, s
A ! l u; et e v e ! A @ @ A ; e O V O V O e
@ G ;
- A e e v e @ ,S A ;
l L_Dj L _CJ lO a e S O 8 e S A3 3 n
<_s C
VB VsB <_; D A I FUNCTION R0D CLUSTERS SHUTDOWN BANK SA 0 SHUTDOWN BANK B Sg SHUTDOW:4 B ANK Sc&SD "A" l CONTF.0L BANK A 4 CONTROL BANK B E CONTROL BANK C 8 CONTROL BANK D 9 PART-LENGTH PL B l Figure 4.3-36 Rod Cluster Control Assembly Pattern CT (cour Loop) V N NT 1 4.3-119 l
. - _ . - .- - . . . . . . . - - ._ - . . . . . . . ~ .. . - . _ _ .._ -. _ _ - - .-
e 669'4-15 l l l l l l 50 E W I40 - R
- 5
! i { 30 - E i e E g 20 -
- c l
5 a: i d l $ 10 - I
- l 0
l 0 50 100 .I50 200 250 < l STEPS WITHDRAWN i 1 l
- Figure 4.3-37 Accidental Simultaneous Withdrawal of Two Control Banks <
, EOL HZP ks D and B Moving in the Some Plane, PL at I st.ps bree and Four Loop) 1 4.3-120 l
O U / 160 I I 140 - I I am - l l _sc ioo - I
- THIMBLE l DASHPOT -
=8 E
1 1 2 I . 60 - { l 5 i : I I l 20 - l , l o l I I I I I I I ! I I O 0.2 0.4 0.6 0.8 1. 0 1.2 1.4 1.6 1.8 2.0 2.2 2. 4 i TIMEOFTRAVEL(SECONDS) O Figure 4. 3-38 Design-Trip Curve y
- i l
6274-58 O l i l I i l 1.00 1 i l l 1 i l i l i I l ' i E l ~ 0. 5 -
-S s .
{ l l t I 1 l i l l 0 I l 100 50 0 l r RODPOSITION(PERCENTINSERTED) i l 1 l l l j, figure 4. 3- 39, Normalized Rod Worth versus Percent insertion l All Rods But One . ! l l r 4.3-122 I
O O O 5 , CURVE 2 CURVE 2 0 - C 5 -5 - y __ _ _ _ A , t
- -10 -
s E CURVE I g-15 q CURVE I d BURNUP PERIOD STABILITY INDEX g -20 ' CURVE (MWD /HTU) (HOURS) (HOURS ~I) I 1550 32.4 -0.041
-2s - 2 noo 27.2 -0.oi4 l I I I I I I I I I
$ 30 I I I I I I I
. M 0 4 8 12 16 20 24 28 32 36 40 44 48 52 56 60 64 68 72 ,
TIME (HOURS) t m Figure 4. 3-40. Axial Offset versus Time PWR Core with a 8 k 12-Ft Height and 121 Assemblies 5 g ~ ; M 1 I I
c !ll
'i '
% w',
O
&h*g >- ]
is "
**=s g 1 1 l 1 6 4 2 0 8 6 2 e 2 4 6 1
8 0 0 f g g g- ,- - - - - - - F 8 g i 4 I u r e 8 I
.O*..
4 3
- 1 I
4 2 1 QX 1 6 I % \
.s uY a
dX r e an H O U 2 S R 0 I g t- O n t o n A T F 2 I S iT T 4 T A E tl se R B I t D i T W 82 I L I I fh T T e f er fYI H r em D 3 N no R A 2 I D cc W E eou A X vp L 3 = a, O 6 - el r s e F
.0 u R R 4 0 s
e C C 0 I 7 Ts 6 ip M N mo E o w (.Hl _ 4 e ns I
- 4 4 I 3 o24 R e l 3 "
)
4 8 I 5 2 I N N 4 1 o O44 5 I 6 R C C 6 0 I E l . l 6 4 O {$ {) I l
oyg yO gMg u _ a8 a$W5 j . O 0 0 4 2 0 0 2 2 0 0 0 2 0 0 8 1 0 0 6 1 0 0 4 1 0 0 2 1 0 0 0 1 0 0 8 0 0 6 0 0 4 0 0 2 o 0
- - - - r- - - - - - - .
-- 10 I
0 9 t s n ere
" io iCc
- f
- +
f t 0 eo oo I 8 CF -
- d2 n1
%'" v W. I 0
7 f t a ce ei eb Dme rs l s, es l p A _ R p1 I 0 E o2 6 WO D1 P d t, L L en r U u N 0 F sPla a OI I 5 F ep T A S O Moo N T dL E N N n-O P O M O T C A K N I S S I E C 0 RE 4 P dw ao eT t N E A P T O N l aL B M N O E T I l cuO B L C O IMA L at R T M R U O E C I 0 3 Ca O N R P L O X A o C B E C 2 4 3 OA - I 0 2 4 e
- r u
s4 i g F N l 0 1 N s s
- - % - - 0 0 0 0 0 0 0 o 0 0 0 0 0 0 6 5 4 3 2 1
{ g7& h u.i8 i i gd*8 O
*.Ygm
4lll i ,
\11 la 5f m
n o 0 i 0 t 0 a 8 r 1 t n e 0 c 0 n I 0 o 6 Ce 1 r no oC r 0 ot 0 Bo I 0 o 4 dF 1 e r2 u1 0 s a , - 0 es I 0 D 2 Me E . 1 i R dl U nb S A U ame E 0 T H 0 /H 0 ds l 0 DW es 1 tA M a l1
, u2 D
P c 1, E 0 U l at T I 0 N Cn A 0 8 R U a L U B fl q C L A C I 0 0 oP np oo so 0 iL 6 r-a2 p mr oo 0 I 0 Cf 0 4 0 3. i 0 4 0 - w0 2 3 4
- e r
u g 0 0 0 0 0 0 g 0 0 0 0 0 0 i 4 1 2 1 0 1 8 6 O 2 F m e"
*'.m
!1l
O O ~O 1300 i i
-l200 -
l l100 - f 1000 - MEASURED l 900 - HEASURED 8* - i 4 l w 700 - L CALCULATED l 600 - r i 500 - i yoo l I I I I I I I 0 900 1600 2700 3600 4500 5400 6300 7200 8100 l 0 Figure 4.3-44. Comparison of Calculated and Measured CB 2-lUUP _ with 121 Assemblies,12 Foot Core a !
f'\ ') r d 1100 1000 - 900 - 800 MEASURED 700 - [600 - d500 - 400 - A C ALCUL ATEO / L, 300 - h m 200 - 100 - l I I I I I I I I I I I i o 0 1000 2000 3000 4000 5000 6000 7000 8000 9000 10000 l1000 12000 13000 14000 BURNUP, NWD/MTU Figure 4. 3-45 Comparison of Calculated and Measured CB in 3-Loop Plant, @ i 157 Assemblies, 12 Foot Core f C i
. _ . _ _ ___ .__. _____-m ...____._-_____..__.-________-________-_--__m_-.m
l R PN ML K J H G F E DC BA lO esa i i MMMMMMM MM VA % MM
- l MM VA YA % VAM l 5% VA % % % MM s l
W VA % % % VA M ' l
##VA VA % % VA VAMM WVA @ % % % VA VAM
##VA % % % VA @MM 9
% YA % % @ V4 & 'O NW % VA VA '4 WM
/ "
#% % VA VA VAM '
#W M VA WM bhM Mbbb I4 ggg 15 I o
l b FIRST CORE RELOAD CORE REGION I ONCE OR TWICE B' BURNED FUEL REGION 2 ONCE OR TWICE BURNED FUEL REGION 3 FRESH FUEL 1 Figure 4.3-1 Fuel Loading Arrangement , l (ThreeLoop) l O 4.3-129 j l
0 T3-0 0 ~ 6 rU 1 U E T s C M d N/ E D R o P R W UE M N F R F 0 N 0 0 0 n o U B D9 I 0 N 4 1 P i b s l e b a N 0 m u S O S N N 0 0 2 B t u o 1 N I O P ht i E B A L
\ d W
T N ) n U O R U 0 U T a S B 0 /M 0 ht D T 0 D i U l O O W W R H M
. T ( p R
W I P u E W U m O N u P R B U e L L U F N 0 B 0 0 E 8 GA l c y T R C O E t V s ~O ( H E N A E F i r T R s O O u s 0 C M 0 0 6 r e v S n
\ O S
l N i t o o m N P E t n) ep L 0 co N B A N ! 0 0 no oL R U 4 Ce N B ne o hr H T r I oT N W B( 0 N ! 0 0 2 3 3 4 N e r u l - '~ - 0 i F g 0 0 0 0 0 O 0 2 0 2 1 4 O 53 54E5M8 = O P&"
*.YC
, ,l
- O g aS S g 3
SOE O E E B B E i i 8 OE O E i E BB S E i 20 BP'S f i i S SO E S 1 3 l 8 OO O E i O E S O 1 E E
- 1O E B OB E IS BP'S t
- 0 9O S O
.) S OO O E O E E O
)
S OO O S 0 8O S O 12 BP'S Figure 4.3-4 Bumable Poisor. Rod Arrangement Within on Assembly (Three Loop) , l 4 .3-1 31
. . . . . _ . - - - , . ~ - - . . . - ~ . . . - - . - - . _ - - - . - - - . _ . . . . _ . . . . - . . . . ~ . . . - ~ ~ - . - - . ~ . - ~ - _
l i l-4 R P -N M L K J H 0 F E D C B A i l l I 16 16 2 3 3 12 16 12 16 12 I l 16 16 16 16 12 16 16 20 16 16 12 l 16 6 16 16 16 16' 46 l 16 16 16 20 16 16 16 l 6 12 20 20 20 20 12 i 16 16 16 20 16 16 46 l 16 16 16 16 16 16 10 e l 12 16 16 20 16 16 12 Il ! 16 16 16 16 32 3 33 12 16 12 16 12 l 16 16 34 i 15 NL'MBER INDICATES NUMBER OF BURN ABLE POISON RODS S INDICATES SOURCE RODS l Figure 4.3-5 Bumable Poison loading Pattem
- (Three Loop) 4.3-132 l'
l O ( H G F E D C B A k~ 1.14 --- --- --- ._ _ __.. ___ ._, _ 9 1.08 1.18 10 1.19 1.17 1.21 I il 1.10 1.20 1.14 1.ll 12 1.21 1.14 1.12 .97 .84 O 13 1.19 1.14 .98 .90 .62 I I4 1.06 1.01 .90 .59 N IS .82 .62 CALCULATED F st.35 KEY: VALUE REPRESENTS ASSEMBLY l RELATIVE POWER t Figure 4.3-6 Normalized Power Densit of Life, Unrodded Core, HotyFull Distribution Near Beginning Power, No Xenon (Three Loop) 4.3-133
1 l q G F lH E D C B A (- 1.16 --- --. -- ._ . __ . _,__ ,__ _q l 9 1.09 1.20 10 1.21 1.18 1.22 Il 3.11 1.21 1.14 1.12 12 1.21 1.14 1.13 .97 .85 , ! l 13 1.18 1.14 .98 .89 .62 I 14 1.05 .99 .88 .59 l \ i l N 15 .80 .61 A F al.33 AH KEY: VALUE REPRESENTS ASSEMBLY , l RELATIVE POWER t i l . Figure 4.3-7 Normalized Power Density Distribution
- Noor Beginning of Ufe, Unrodded Core,
- Hot Full Power, Equilibrium Xenon (ThreeLoop) 4.3-134
- e
. - - . . . . = . . . . . . . . . . _ . . -.. - . .. - -- -.-. - . __ _ . . - ........ _ . .._ _ _. = -
l l l I i lH G F E D C B A k- 1.19 --- -- - -- - .. . __ _ . . ,_,, _q l 1 l l l 9 1.11 1.20 i l l6 10 1.22 1.14 1.05 l ll 1.13 1.22 1.12 1.15 l l 12 1.24 1.17 1.16 1.01 .90 l ! I 13 1.15 1.14 1.01 .94 .66 l 14 .88 .94 .90 .62 ? l l N l 15 .72 .57 CALCULATED F g 1.325 i i KEY: VALUE REPRESENTS ASSEMBLY
! RELATIVE POWER i.
! Figure 4.3-8 Normlized Power Density Distribution Near Beginning of Life, Group D 30% Inserted, Hot Full Power, Equilibrium Xenon (ThreeLoop) 4.3-135
.. ~ . .- - . . . . .- -. . . . . . - - . ~ _ - . - . . - - . . - . . . . . . - _ . . . _ . - . _ . . , . - ~ _ - . .
i ( G lH F E D C B A 1.00 - i ,- -- --- ._- .__ __. ___ ._. __ q 9 1.05 1.17 t 10 1.19 1.13 1.06 11 1.07 1.19 1.14 1.18 12 1.05 1.13 1.18 1.05 .95 l 13 1.11 1.13 1.03 .98 .70 ! 14 .89 .96 .93 .64 1 l 15 .92 .59 CALCULATED F " 1.323 KEY: VALUE REPRESENTS ASSEMBLY RELATIVE POWER t l l i
- Figure 4.3-9 Normalized Power Density Distribution Near Beginning
- of Ufe, Group D 30% inserted, Plus PL Inserted, Hot Full Power, Equilibrium Xenon .
!. (Three Loop) i l 4.3-136 i l
- - - ~ - -
, - . ~ - . - - - . - . , , . . - - ~ . . . . . _ - . _ . - . . ~ - . - _ - . . . . . . . . ~ . - . -
. . . . . . . - . ~ . . .
i ( lH G F E D C B A
- 1. 18 ._ _q (8- __. ___ __ ___ .,__ ,_
9 1.21 1.18 l 10 1.18 1.23 1. 17 I i 11 1.20 1.16 1.20 1.11 12 1.14 1. 18 1. 10 1.07 .87 i 13 1. 15 1.06 1.04 .95 .63 i 14 .94 .95 .84 .59 M l 15 .69 .55 CALCULATED F 1.29 l AH KEY: VALUE REPRESLNTS ASSEMBLY RELATIVE POWER i Figure 4.3-10 Normalized Power Density Distribution Near Middle of Life, Unrodded Core, Hot Full Pbwer, Equilibrium Xenon (Three Loop) 4.3-137 l
_m__.._ _ . . _ _ _ . _ . . _ _ _ . _ -..-______..__.m.___ _ . . _ . _ . . _ _ . _ _ _ _ . _ - . _ . , . _ .) i J l
, I
( I.t lH G F E D C B A (-- 1.07 --- --. -- __ _ ._. ___ .__ _q 9 1.13 1.07 10 1.08 1.15 1.08 4 l 11 1.14 1.08 f.16 1.08 1 l
)
12 1.09 1.16 1.09 1.11 .92 ! 1 13 1.15 1.06 1.09 1.03 .69 i 14 .97 1.02 .88 . 6 4, l 15 .74 61 CALCULATED F " 1.19 AH KEY: l VALUE REPRESENTS ASSEMBLY C t RELATIVE POWER Figure 4.3-11 Normalized Power Density Distribution Near End of Life, Unrodded Core, Hot Full Power Equilibrium Xenon ( (Three Loop) 4.3-138
I l l i N l.08 I.08 f.10 l f 09 f. I f 1.f5 1.09 1.13 1.20 1.09 1.15 1.21 1.24 1.24 l 1.10 1.17 f.25 1.26 1.09 1.15 1.22 1.22 1.23 1.26 f.24 l 1.09 1.15 1.22 1.22 1.23 1.27 1.25 1.25 i l.10 1.18 8.24 1.26 1.27 1.28 1
- 1. 10 1. 16 1.22 1.22 1.24 1.27 1.25 1.26 1.28 1.26 l.10 1.16 1.22 1.23 1.24 1. 27 1.25 f.26 1.28 1.26 1.26 1.11 1. 18 1.26 1.27 f . 28 1.28 1.28 1.28 1.11 1.16 1.23 1.26 1.26 1.28 f.25 1.25 1.28 f.26 1.26 1.29 1.27 l
1.11 1.15 1.22 1.27 f.27 1.24 1.24 1.27 1.25 1.25 1.28 f.28 l 1.11 f.14 1.17 1.22 1.24 1.25 1.25 1.25 1.25 1.26 f.24 1. 19 l 1.11 1.12 1.14 1.16 1. 18 1.21 1.19 1.20 1.22 1.20 1.20 1.22 1.19 1. 18 1.16 1.15 l 1.11 1.11 1.12 1.12 1.13 1.14 1.14 f .15 1.16 1.15 1.15 1.15 1.14 1.14 1.13 1.13 1.14 l I l l l l r Figure 4.3-12 Rodwise Power Distribution in o Typical Assembly (Assembly G-9) Near Beginning of Life, Hot Full Power, Equilibrium Xenory Unn>dded Core (Three Loop) t 4.3-139 I l , l I
l O ! 3 99 1.00 1.00 1.02 l.05 1.01 1.04 1.09 1.02 1.05 1. I i 1.13 1.Ii ; i.02 1.08 1.13 1.13 1.02 1.06 1.10 1.09 1.10 1.12 1.10 l.02 1.06 1.10 1.09 1.09 1.12 1.10 1.10 a l 1.03 1.08 1.11 1.12 1.12 1.'2 1.02 1.06 1. 10 1.09 1.09 1.12 1. 40 1.10 1.12 1.10 j i 1.02 1.06 1. 10 1.09 1.10 1.12 1.10 1. 10 !.12 I.10 1.10 1.02 1.08 1.13 1.13 1.12 1.12 1.12 1.12 , 1.02 1.05 1.11 1.13 1.11 1.13 1. 10 1.10 1.12 1.10 1.10 1.13 1.42 1.01 1.04 1.09 1.13 1.13 1.09 1.09 1.12 1.09 1.09 1. 13 1.13
- 1.00 1.02 1.05 1.09 1. I I 1. I i 1. 10 1.10 1. 11 1. I i .09 1.05 1.00 1.00 1.02 1.04 1.05 1.08 l.06 1.06 1.08 1.06 1.06 1.08 1.06 1.04 1.02 1.01 1.00 1.00 1.00 1.01 1.02 1.03 1.03 1.03 1.03 1.03 l.03 1.03 1.02 1.01 1.00 1.00 1.00 l
~
Figure 4.3-13 Rodwise Power Distribution in a Typical Assembly (Assembi G-9) Near End of Life, Hot Full O Power, kullibrium Xenon, Unrodded Core V (Three Loop) 4.3-140 j
. . - - _ . . .- ~ ._. . . - . - _ =- _ -
6937-120 O KW/FT 25 O = UNCONTROLLED ROD BANK MOTION X = OPERATOR ERROR 20 X XX X X k 15 x[{ X O O XX ' O , 1 I l l l l l l l l
-% -30 -20 -10 0 10 20 30 %
FLUX DIFFERENCE ( AI) l Figure 4.3-23 Peak Linear Power versus Flux Differences for Determination of Protection Set Points (3 Loop) O , C/ AMENDMENT 1 4.3-141 x /
1 I 6694-10 O, 5 _ 0 - M u 2 3 400 r 6 x -5 - 2 g 547 F g -10 - G
. I b
8 -15 - 582 r O I E w g -20 - 8
,5 -25 -
E
-30 0 500 1000 1500 SOLUBLE BORON CONCENTRATION (PPM)
Figure 4.3-32 Moderator Temperature Coefficient as a Function of Baron Concentration - BOL Cycle 1, No Rods (Three Loops) O AMENDMENT 1 4.3-142
- _________ A
R P N M L K J H G F E D C B A O O @ @ A A
@ @ O @ @ 4 A A s
@ @ @ @ @ @ @ 6 A A A A 7
'o
@ G @ O @ O @ 27o 8 i A A & A '
& a "
O O O O G O 15 00 FUNCTION NUMBER OF CLUSTERS L ANK s"st"i"l1 '
!"M 7 l^"l '8 PART LENGTH P 5 Figure 4.3-36 Cluste control Assembly Pattern $ 4.3-u3 I
L. . ..
._ _A}}