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| number = ML14113A088 | | number = ML14113A088 | ||
| issue date = 04/11/2014 | | issue date = 04/11/2014 | ||
| title = | | title = Updated Final Safety Analysis Report, Revision 20, Chapter 4.0, Reactor | ||
| author name = | | author name = | ||
| author affiliation = Exelon Generation Co, LLC | | author affiliation = Exelon Generation Co, LLC | ||
| Line 17: | Line 17: | ||
=Text= | =Text= | ||
{{#Wiki_filter:}} | {{#Wiki_filter:LSCS-UFSAR 4.0-i REV. 20, APRIL 2014 CHAPTER 4.0 - REACTOR TABLE OF CONTENTS Page 4.0 REACTOR 4.1-1 4.1 | ||
==SUMMARY== | |||
DESCRIPTION 4.1-1 | |||
====4.1.1 Reactor==== | |||
Vessel 4.1-1 4.1.2 Reactor Internal Components 4.1-1 4.1.2.1 Reactor Core 4.1-1 4.1.2.1.1 General 4.1-1 4.1.2.1.2 Core Configuration 4.1-3 4.1.2.1.3 Fuel Assembly Description 4.1-4 4.1.2.1.3.1 Fuel Rod 4.1-4 4.1.2.1.3.2 Fuel Bundle 4.1-4 4.1.2.1.4 Assembly Support and Control Rod Location 4.1-5 4.1.2.2 Shroud 4.1-5 4.1.2.3 Shroud Head and Steam Separators 4.1-6 4.1.2.4 Steam Dryer Assembly 4.1-6 4.1.3 Reactivity Control Systems 4.1-7 4.1.3.1 Operation 4.1-7 4.1.3.2 Description of Rods 4.1-7 4.1.3.3 Supplementary Reactivity Control 4.1-8 4.1.4 Analysis Techniques 4.1-8 4.1.4.1 Reactor Internal Components 4.1-8 4.1.4.1.1 MASS (Mechanical Analysis of Space Structure) 4.1-9 4.1.4.1.1.1 Program Description 4.1-9 4.1.4.1.1.2 Program Version and Computer 4.1-9 4.1.4.1.1.3 History of Use 4.1-9 4.1.4.1.1.4 Extent of Application 4.1-9 4.1.4.1.2 SNAP (MU LTISHELL) 4.1-9 4.1.4.1.2.1 Program Description 4.1-9 4.1.4.1.2.2 Program Version and Computer 4.1-10 4.1.4.1.2.3 History of Use 4.1-10 4.1.4.1.2.4 Extent of Application 4.1-10 4.1.4.1.3 GASP 4.1-10 | |||
4.1.4.1.3.1 Program Description 4.1-10 4.1.4.1.3.2 Program Version and Computer 4.1-10 4.1.4.1.3.3 History of Use 4.1-11 4.1.4.1.3.4 Extent of Application 4.1-11 4.1.4.1.4 NO HEAT 4.1-11 4.1.4.1.4.1 Program Description 4.1-11 4.1.4.1.4.2 Program Version and Computer 4.1-11 LSCS-UFSAR Table of Contents (Cont'd) 4.0-ii REV. 20, APRIL 2014 4.1.4.1.4.3 History of Use 4.1-11 4.1.4.1.4.4 Extent of Application 4.1-12 4.1.4.1.5 FI NITE 4.1-12 4.1.4.1.5.1 Program Description 4.1-12 4.1.4.1.5.2 Program Version and Computer 4.1-12 4.1.4.1.5.3 History of Use 4.1-12 4.1.4.1.5.4 Extent of Application 4.1-12 4.1.4.1.6 SAMIS 4.1-12 4.1.4.1.6.1 Program Description 4.1-12 4.1.4.1.6.2 Program Version and Computer 4.1-13 4.1.4.1.6.3 History of Use 4.1-13 4.1.4.1.6.4 Extent of Application 4.1-13 4.1.4.1.7 General Matrix Manipulation Program(GEMOP) 4.1-14 4.1.4.1.7.1 Program Description 4.1-14 4.1.4.1.7.2 Program Version and Computer 4.1-14 4.1.4.1.7.3 History of Use 4.1-14 4.1.4.1.7.4 Extent of Application 4.1-14 4.1.4.1.8 SHELL 5 4.1-14 | |||
4.1.4.1.8.1 Program Description 4.1-14 4.1.4.1.8.2 Program Version and Computer 4.1-15 4.1.4.1.8.3 History of Use 4.1-15 4.1.4.1.8.4 Extent of Application 4.1-15 4.1.4.1.9 HEATER 4.1-15 4.1.4.1.9.1 Program Description 4.1-15 4.1.4.1.9.2 Program Version and Computer 4.1-15 4.1.4.1.9.3 History of Use 4.1-16 4.1.4.1.9.4 Extent of Application 4.1-16 4.1.4.1.10 FAP-71 (Fatigue Analysis Program) 4.1-16 4.1.4.1.10.1 Program Description 4.1-16 4.1.4.1.10.2 Program Version and Computer 4.1-16 4.1.4.1.10.3 History of Use 4.1-16 4.1.4.1.10.4 Extent of Application 4.1-16 4.1.4.1.11 CREEP/P LASTICITY 4.1-17 4.1.4.1.11.1 Program Description 4.1-17 4.1.4.1.11.2 Program Version and Computer 4.1-17 4.1.4.1.11.3 History of Use 4.1-17 4.1.4.1.11.4 Extent of Application 4.1-17 4.1.4.1.12 SAP4G07 and ANSYS 4.1-17 4.1.4.2 Fuel Rod Thermal Analysis 4.1-17 4.1.4.3 Reactor Systems Dynamics 4.1-18 4.1.4.4 Nuclear Engineering Analysis 4.1-18 4.1.4.5 Neutron Fluence Calculations 4.1-18 4.1.4.6 Thermal Hydraulic Calculations 4.1-19 LSCS-UFSAR Table of Contents (Cont'd) 4.0-iii REV. 15, APRIL 2004 4.1.5 References 4.1-19 4.2 FUEL SYSTEM 4.2-1 4.2.1 Design Bases 4.2-1 4.2.1.1 Safety Design Bases 4.2-1 4.2.1.2 Power Generation Design Basis 4.2-3 4.2.1.2.1 Material Selection 4.2-3 4.2.1.2.2 Effects of Irradiat ion and Fuel Swelling 4.2-3 4.2.1.2.3 Fuel De nsification 4.2-4 4.2.1.2.3.1 GE Fuel 4.2-4 4.2.1.2.3.2 FANP Fuel 4.2-5 4.2.1.2.4 Incipient UO 2 Center Melting 4.2-5 4.2.1.2.4.1 GE Fuel 4.2-5 4.2.1.2.4.2 FANP Fuel 4.2-5 4.2.1.2.5 Maximum Allowable Stresses 4.2-6 4.2.1.2.5.1 GE Fuel 4.2-6 4.2.1.2.5.2 FANP Fuel 4.2-7 4.2.1.2.6 Capacity for Fission Gas Inventory 4.2-7 4.2.1.2.7 Maximum Internal Gas Pressure 4.2-7 4.2.1.2.7.1 GE Fuel 4.2-7 4.2.1.2.7.2 FANP Fuel 4.2-8 4.2.1.2.8 Internal Pressure and Cladding Stresses During Normal Conditions 4.2-8 4.2.1.2.9 Cycling and Fatigue Limits 4.2-8 4.2.1.2.9.1 GE Analysis 4.2-8 4.2.1.2.9.2 FANP Analysis 4.2-9 4.2.1.2.10 Deflections 4.2-9 4.2.1.2.10.1 GE Evaluation 4.2-9 4.2.1.2.10.2 FANP Evaluation 4.2-9 4.2.1.2.11 Flow-Induced Fuel Rod Vibrations 4.2-10 4.2.1.2.12 Fretting Corrosion 4.2-10 4.2.1.2.13 Seismic Loadings 4.2-10 4.2.1.2.14 Chemical Proper ties of Cladding and Fuel Material 4.2-11 4.2.1.2.15 Design Ratios 4.2-11 4.2.1.2.15.1 Limiting Parameter Values 4.2-11 4.2.1.2.15.1.1 Normal and Upset Design Conditions 4.2-11 4.2.1.2.15.1.2 Emergency and Faulted Design Conditions 4.2-12 4.2.1.2.15.2 Actual Parameter Values 4.2-12 4.2.1.2.16 Fuel Assembly Limits 4.2-13 4.2.1.2.16.1 Fuel Rods 4.2-13 4.2.1.2.16.2 Fuel Spacer 4.2-13 4.2.1.2.16.3 Water Rods 4.2-14 LSCS-UFSAR Table of Contents (Cont'd) 4.0-iv REV. 15, APRIL 2004 4.2.1.2.16.4 Channel 4.2-14 4.2.1.2.16.5 Tie Plates 4.2-15 4.2.1.2.17 Reactivity Control Assembly and Burnable Poison Rods 4.2-15 4.2.1.2.17.1 Safety Design Bases for Reactivity Control 4.2-15 4.2.1.2.17.1.1 Specific Design Characteristics 4.2-16 4.2.1.2.18 Surveillance Program 4.2-17 | |||
====4.2.2 Description==== | |||
and Design Drawings 4.2-18 4.2.2.1 Core Cell 4.2-18 4.2.2.2 Fuel Assembly 4.2-18 4.2.2.3 Fuel Bundle 4.2-18 4.2.2.4 Fuel Rod 4.2-19 4.2.2.5 Fuel Pellets 4.2-21 4.2.2.6 Fuel Channel 4.2-22 4.2.2.7 Reactivity Control Assembly and Burnable Poison Rods 4.2-23 4.2.2.7.1 Control Rods 4.2-23 4.2.2.7.1.1 General Electric Control Rods 4.2-22 4.2.2.7.1.2 ASEA-ATOM Control Rods 4.2-24 4.2.2.7.2 Velocity Limiter 4.2-25 4.2.2.7.3 Burnable Poison Rods 4.2-25 4.2.3 Design Limits and Evaluation 4.2-26 4.2.3.1 Fuel Damage Analysis 4.2-26 4.2.3.2 Fuel Damage Experience 4.2-27 4.2.3.3 Potential for a Water-Logging Rupture 4.2-28 4.2.3.4 Potential for Hydriding 4.2-29 4.2.3.5 Dimensional Stability 4.2-29 4.2.3.6 Fuel Densification 4.2-30 4.2.3.7 Fuel Cladding Temperatures 4.2-30 4.2.3.8 Peaking Factors 4.2-31 4.2.3.8.1 Local Peak ing Factors 4.2-31 4.2.3.8.2 Axial and Gross Peaking Factors 4.2-31 4.2.3.9 Temperature Transients with Water- logged Fuel Element 4.2-31 4.2.3.10 Potential Damaging Temperature Effects During Transients 4.2-32 4.2.3.11 Energy Release During Fuel Element Burnout 4.2-32 4.2.3.12 Energy Release fo r Rupture of Water- logged Fuel Elements 4.2-34 4.2.3.13 Fuel Rod Behavior Effects from Coolant Flow Blockage 4.2-34 4.2.3.14 Channel Evaluation 4.2-35 LSCS-UFSAR Table of Contents (Cont'd) 4.0-v REV. 15, APRIL 2004 4.2.3.15 Fuel Reliability 4.2-36 4.2.3.16 Fuel Operating and Developmental Experience 4.2-37 4.2.3.17 Fuel Assembly 4.2-37 4.2.3.17.1 Loads Assessment of Fuel Assembly Components 4.2-38 4.2.3.18 Spacer Grid and Channel Boxes 4.2-38 4.2.3.19 Burnable Poison Rods 4.2-38 4.2.3.20 Control Rods 4.2-38 4.2.3.20.1 Materials Adequacy Throughout Design Lifetime 4.2-38 4.2.3.20.2 Dimensional and Tolerance Analysis 4.2-38 4.2.3.20.3 Thermal Analys is of the Tendency to Warp 4.2-39 4.2.3.20.4 Forces for Expulsion 4.2-39 4.2.3.20.5 Functional Failure of Critical Components 4.2-39 4.2.3.20.6 Precluding Excessive Rates of Reactivity Addition 4.2-39 4.2.3.20.7 Effect of Fuel Rod Failure on Control Rod Channel Clearances 4.2-39 4.2.3.20.8 Mechanical Damage 4.2-39 4.2.3.20.8.1 First Mode of Failure 4.2-40 4.2.3.20.8.2 Second Mode of Failure 4.2-40 4.2.3.20.9 Analysis of Guide Tube Design 4.2-40 4.2.3.20.10 Evaluation of Control Rod Velocity Limiter 4.2-41 4.2.3.21 Rod Bowing 4.2-41 4.2.3.21.1 GE Evaluation 4.2-41 4.2.3.21.2 FANP Evaluation 4.2-42 4.2.3.22 Fission Gas Release 4.2-42 4.2.3.23 Ballooning and Rupture 4.2-43 4.2.3.23.1 GE Evaluation 4.2-43 4.2.3.23.2 FANP Evaluation 4.2-44 4.2.4 Testing and Inspection Plan 4.2-44 4.2.4.1 Testing and Inspection (Enrichment and Burnable Poison Concentrations) 4.2-45 4.2.4.1.1 Enrichment Control Program 4.2-45 4.2.4.1.2 Gadolinia Inspections 4.2-46 4.2.4.1.3 Reactor Control Rods 4.2-47 4.2.4.2 Surveillance Inspection and Testing of Irradiated Fuel Rods 4.2-47 LSCS-UFSAR Table of Contents (Cont'd) 4.0-vi REV. 15, APRIL 2004 4.2.4.3 Operating Experience with Gadolinia-Containing Fuel 4.2-48 4.2.5 References 4.2-49 | |||
===4.3 NUCLEAR=== | |||
DESIGN 4.3-1 | |||
====4.3.1 Design==== | |||
Bases 4.3-1 4.3.1.1 Safety Design Bases 4.3-1 4.3.1.2 Power Generation Design Bases 4.3-2 4.3.2 Description 4.3-2 4.3.2.1 Nuclear Design Description 4.3-2 4.3.2.1.1 Fuel Nuclear Properties 4.3-3 4.3.2.2 Power Distributions 4.3-4 4.3.2.2.1 Local Power Distribution 4.3-5 4.3.2.2.2 Radial Power Distribution 4.3-5 4.3.2.2.3 Axial Power Distribution 4.3-5 4.3.2.2.4 Power Distributi on Calculations 4.3-6 4.3.2.2.5 Power Distribution Measurements 4.3-6 4.3.2.2.6 Power Distribution Accuracy 4.3-6 4.3.2.2.7 Power Distribution Anomalies 4.3-6 4.3.2.3 Reactivity Coefficients 4.3-7 4.3.2.3.1 Void Reactivity Coefficients 4.3-7 4.3.2.3.2 Moderator Temperature Coefficient 4.3-7 4.3.2.3.3 Doppler Reactivity Coefficient 4.3-8 4.3.2.3.4 Power Coefficient 4.3-9 4.3.2.4 Control Requirements 4.3-9 4.3.2.4.1 Shutdown Reactivity 4.3-9 4.3.2.4.2 Reactivity Variations 4.3-10 4.3.2.5 Control Rod Patterns and Reactivity Worths 4.3-11 4.3.2.5.1 Control Rod Withdrawal Sequences 4.3-11 4.3.2.5.1.1 Control Rod Withdrawal Sequences in the Startup Range 4.3-12 4.3.2.5.1.2 Control Rod Withdrawal Sequences in the RWM Power Range 4.3-13 4.3.2.5.1.3 Maximum Control Rod Worth Pattern with a Single Error in the RWM Power Range 4.3-14 4.3.2.5.2 Control Rod Worth Calculations 4.3-14 4.3.2.5.2.1 Control Rod Worth in the Startup Range and RWM Power Range 4.3-14 4.3.2.5.2.2 Control Rod Worth in the Reactor Power Range > 10% Rated Power 4.3-15 4.3.2.5.3 Scram Reactivity 4.3-15 LSCS-UFSAR Table of Contents (Cont'd) 4.0-vii REV. 15, APRIL 2004 4.3.2.6 Criticality of Reactor During Refueling 4.3-16 4.3.2.6.1 Criticality of Reactor 4.3-16 4.3.2.6.2 Criticality of Fuel Assemblies 4.3-16 4.3.2.7 Stability 4.3-16 4.3.2.7.1 Xenon Transients 4.3-16 4.3.2.7.2 Thermal Hydraulic Stability 4.3-17 4.3.2.8 Vessel Irradiation 4.3-17 4.3.3 Analytical Methods 4.3-17 | |||
====4.3.4 References==== | |||
4.3-18 | |||
===4.4 THERMAL=== | |||
AND HYDRAULIC DESIGN 4.4-1 | |||
====4.4.1 Design==== | |||
Bases 4.4-1 4.4.1.1 Safety Design Bases 4.4-1 4.4.1.2 Power Generation Design Bases 4.4-1 4.4.1.3 Requirements for Steady-State Conditions 4.4-1 4.4.1.4 Requirements for Transient Conditions 4.4-2 4.4.1.5 Summary of Design Bases 4.4-2 4.4.1.5.1 Fuel Cladding Integrity 4.4-3 4.4.1.5.2 Fuel Assembly Integrity 4.4-3 4.4.1.5.3 Fuel-Cladding Gap Characteristics 4.4-3 4.4.2 Description of Thermal Hydraulic Design of Reactor Core 4.4-3 4.4.2.1 Summary Comparison 4.4-3 4.4.2.2 Critical Power Ratio 4.4-3 4.4.2.2.1 Boiling Correlations 4.4-4 4.4.2.2.1.1 GE Fuel 4.4-4 4.4.2.2.1.2 FANP Fuel 4.4-4 4.4.2.3 Maximum Average Planar Linear Heat Generation Rate (MAPLHGR) 4.4-5 4.4.2.3.1 Design Power Distribution 4.4-6 4.4.2.4 Void Fraction Distribution 4.4-7 4.4.2.5 Core Coolant Flow Distribution 4.4-7 4.4.2.6 Core Pressure Drop and Hydraulic Loads 4.4-8 4.4.2.6.1 Friction Pressure Drop 4.4-8 4.4.2.6.2 Local Pressure Drop 4.4-9 4.4.2.6.3 Elevation Pressure Drop 4.4-9 4.4.2.6.4 Acceleration Pressure Drop 4.4-10 LSCS-UFSAR Table of Contents (Cont'd) 4.0-viii REV. 17, APRIL 2008 4.4.2.7 Correlation and Physical Data 4.4-11 4.4.2.7.1 Pressure Drop Correlations 4.4-11 4.4.2.7.2 Void Fraction Correlation 4.4-11 4.4.2.7.3 Heat Transfer Correlation 4.4-12 4.4.2.8 Thermal Effects of Operational Transients 4.4-12 4.4.2.9 Uncertainties in Estimates 4.4-12 4.4.2.9.1 Transition Boiling Uncertainties 4.4-12 4.4.2.9.2 Variation of Fuel Damage Limit 4.4-13 4.4.2.9.3 Effects of Misoriented Fuel Bundle 4.4-13 4.4.2.10 Flux Tilt Considerations 4.4-13 4.4.3 Description of the Thermal and Hydraulic Design of the Reactor Coolant System 4.4-13 4.4.3.1 Plant Configuration Data 4.4-13 4.4.3.2 Operating Restrictions on Pumps 4.4-14 4.4.3.3 Power-Flow Operating Map 4.4-14 | |||
4.4.3.4 Temperature-Power Operating Map (PWR) 4.4-14 4.4.3.5 Load-Following Characteristics 4.4-14 4.4.3.6 Thermal and Hydraulic Characteristics Summary Table 4.4-14 4.4.4 Evaluation 4.4-14 4.4.4.1 Critical Heat Flux 4.4-14 4.4.4.2 Core Hydraulics 4.4-14 4.4.4.3 Influence of Power Distribution 4.4-14 4.4.4.4 Core Thermal Response 4.4-15 4.4.4.5 Analytical Methods 4.4-15 4.4.4.5.1 Reactor Model 4.4-15 4.4.4.5.2 System Flow Balances 4.4-16 4.4.4.5.3 System Heat Balances 4.4-17 4.4.4.5.4 Uncertainties in Design Analyses 4.4-18 4.4.4.6 Reactor Stability Analysis 4.4-18 4.4.4.6.1 Introduction 4.4-18 4.4.4.6.2 Description 4.4-19 4.4.4.6.3 Solution Description for Thermal-Hydraulic Stability 4.4-19 4.4.4.6.4 Stability Criteria 4.4-20 4.4.4.6.5 Expected Oscillation Modes 4.4-21 4.4.4.6.6 Analysis Approach 4.4-22 4.4.4.6.7 Mathematical Model 4.4-23 4.4.4.6.8 Initial Core Analysis Results 4.4-24 4.4.5 Testing and Verification 4.4-25 4.4.6 Instrumentation Requirements 4.4-25 4.4.6.1 Loose Parts Monitoring System (Deleted) 4.4-25 | |||
====4.4.7 References==== | |||
4.4-27 LSCS-UFSAR Table of Contents (Cont'd) 4.0-ix REV. 15, APRIL 2004 4.5 REACTOR MATERIALS 4.5-1 4.5.1 Control Rod System Structural Materials 4.5-1 4.5.1.1 Material Specifications 4.5-1 4.5.1.2 Special Materials 4.5-2 4.5.1.3 Processes, Inspections and Tests 4.5-2 4.5.1.4 Control of Delta Ferrite Content 4.5-3 4.5.1.5 Protection of Materials During Fabrication, Shipping and Storage 4.5-3 4.5.2 Reactor Internals Materials 4.5-4 4.5.2.1 Material Spec ifications 4.5-4 4.5.2.2 Controls on Welding 4.5-6 4.5.2.3 Nondestructive Examination of Wrought Seamless Tubular Products 4.5-6 4.5.2.4 Fabrication and Processing of Austenitic Stainless Steel 4.5-6 4.5.2.5 Regulatory Guide Conformance Assessment 4.5-6 4.6 FUNCTIONAL DESIGN OF REACTIVITY CONTROL SYSTEMS 4.6-1 4.6.1 Information for Control Rod Drive Systems (CRDS) 4.6-1 4.6.1.1 Control Rod Drive System Design 4.6-1 4.6.1.1.1 Design Bases 4.6-1 4.6.1.1.1.1 General Design Bases 4.6-1 4.6.1.1.1.1.1 Safety Design Bases 4.6-1 4.6.1.1.1.1.2 Power Generation Design Basis 4.6-2 4.6.1.1.2 Description 4.6-2 4.6.1.1.2.1 Control Rod Drive Mechanisms 4.6-2 4.6.1.1.2.2 Drive Components 4.6-3 4.6.1.1.2.2.1 Drive Piston 4.6-3 4.6.1.1.2.2.2 Index Tube 4.6-4 4.6.1.1.2.2.3 Collet Assembly 4.6-4 4.6.1.1.2.2.4 Piston Tube 4.6-4 4.6.1.1.2.2.5 Stop Piston 4.6-5 4.6.1.1.2.2.6 Flange and Cylinder Assembly 4.6-5 4.6.1.1.2.2.7 Lock Plug 4.6-6 4.6.1.1.2.3 Materials of Construction 4.6-6 4.6.1.1.2.3.1 Index Tube 4.6-6 4.6.1.1.2.3.2 Coupling Spud 4.6-7 4.6.1.1.2.3.3 Collet Fingers 4.6-7 4.6.1.1.2.3.4 Seals and Bushings 4.6-7 4.6.1.1.2.3.5 Summary 4.6-7 LSCS-UFSAR Table of Contents (Cont'd) 4.0-x REV. 15, APRIL 2004 4.6.1.1.2.4 Control Rod Drive Hydraulic System 4.6-8 4.6.1.1.2.4.1 Hydraulic Requirements 4.6-8 4.6.1.1.2.4.2 System Description 4.6-9 4.6.1.1.2.4.2.1 Supply Pump 4.6-9 | |||
4.6.1.1.2.4.2.2 Accumulator Charging Pressure 4.6-10 4.6.1.1.2.4.2.3 Drive Water Pressure 4.6-10 | |||
4.6.1.1.2.4.2.4 Cooling Water Header 4.6-11 4.6.1.1.2.4.2.5 Return Line 4.6-11 4.6.1.1.2.4.2.6 Scram Di scharge Volume 4.6-11 4.6.1.1.2.4.3 Hydraulic Control Units 4.6-12 4.6.1.1.2.4.3.1 Insert Drive Valve 4.6-12 4.6.1.1.2.4.3.2 Insert Exhaust Valve 4.6-13 4.6.1.1.2.4.3.3 Withdr aw Drive Valve 4.6-13 4.6.1.1.2.4.3.4 Withdraw Exhaust Valve 4.6-13 4.6.1.1.2.4.3.5 Speed Control Valves 4.6-13 4.6.1.1.2.4.3.6 Scram Pilot Valves 4.6-13 4.6.1.1.2.4.3.7 Scram Inlet Valve 4.6-13 4.6.1.1.2.4.3.8 Scram Exhaust Valve 4.6-14 4.6.1.1.2.4.3.9 Scram Accumulator 4.6-14 4.6.1.1.2.4.3.10 Alternate Rod Insertion Scram Valves 4.6-14 4.6.1.1.2.5 Control Rod Drive System Operation 4.6-15 4.6.1.1.2.5.1 Rod Insertion 4.6-15 4.6.1.1.2.5.2 Rod Withdrawal 4.6-15 4.6.1.1.2.5.3 Scram 4.6-16 4.6.1.1.2.6 Instru mentation 4.6-17 4.6.1.2 Control Rod Driv e Housing Supports 4.6-17 4.6.1.2.1 Safety Objective 4.6-17 4.6.1.2.2 Safety Design Bases 4.6-17 4.6.1.2.3 Description 4.6-17 4.6.2 Evaluations of the CRDS 4.6-19 4.6.2.1 Failure Mode and Effects Analysis 4.6-19 4.6.2.2 Protection from Co mmon Mode Failures 4.6-19 4.6.2.3 Safety Evaluation 4.6-19 4.6.2.3.1 Control Rod Drives 4.6-19 4.6.2.3.1.1 Evaluation of Scram Time 4.6-19 4.6.2.3.1.2 Analysis of Malfunction Relating to Rod Withdrawal 4.6-20 4.6.2.3.1.2.1 Drive Housing Fails at Attachment Weld 4.6-20 4.6.2.3.1.2.2 Rupture of Hydraulic Line(s) to Drive Housing Flange 4.6-21 4.6.2.3.1.2.2.1 Pressure-U nder Line Break 4.6-21 4.6.2.3.1.2.2.2 Pressure-Over Line Break 4.6-22 4.6.2.3.1.2.2.3 Simultaneous Breakage of Pressure- Over and Pressure-Under Lines 4.6-22 LSCS-UFSAR Table of Contents (Cont'd) 4.0-xi REV. 15, APRIL 2004 4.6.2.3.1.2.3 All Drive Flange Bolts Fail in Tension 4.6-22 4.6.2.3.1.2.4 Weld Joinin g Flange to Housings Fails in Tension 4.6-23 4.6.2.3.1.2.5 Housin g Wall Ruptures 4.6-24 4.6.2.3.1.2.6 Flange Plug Blows Out 4.6-25 4.6.2.3.1.2.7 Drive Pressure Control Valve Closure (Reactor Pressure, 0 psig) 4.6-26 4.6.2.3.1.2.8 Ball Check Valve Fails to Close Passage to Vessel Ports 4.6-26 4.6.2.3.1.2.9 Hydraulic Co ntrol Unit (HCU) Valve Failures 4.6-26 4.6.2.3.1.2.10 Collet Fing ers Fail to Latch 4.6-27 4.6.2.3.1.2.11 Withdrawal Speed Control Valve Failure 4.6-27 4.6.2.3.2 Scram Reliability of CRDS 4.6-27 4.6.2.3.2.1 Reliabilit y Analysis 4.6-28 4.6.2.3.2.2 Control Rod Su pport and Operation 4.6-28 4.6.2.3.3 Control Rod Driv e Housing Supports 4.6-28 4.6.2.3.3.1 Safety Evaluation 4.6-28 4.6.3 Testing and Verifica tion of the CRDS 4.6-29 4.6.3.1 Control Rods 4.6-29 4.6.3.1.1 Testing an d Inspection 4.6-29 4.6.3.2 Control Rod Drives 4.6-29 4.6.3.2.1 Testing an d Inspection 4.6-29 4.6.3.2.1.1 Develo pment Tests 4.6-29 4.6.3.2.1.2 Factory Quality Control Tests 4.6-30 4.6.3.2.1.3 Operat ional Tests 4.6-31 4.6.3.2.1.4 Accepta nce Tests 4.6-31 4.6.3.2.1.5 Surveillance Tests 4.6-32 4.6.3.3 Control Rod Driv e Housing Supports 4.6-34 4.6.3.3.1 Testing an d Inspection 4.6-34 4.6.4 Information for Combined Performance of Reactivity Systems 4.6-34 4.6.4.1 Vulnerability to Common Mode Failures 4.6-34 4.6.4.2 Accidents Taking Credit for Two or More Reactivity Control Systems 4.6-34 4.6.5 Evaluation of Combined Performance 4.6-34 4.6.6 References 4.6-35 LSCS-UFSAR 4.0-xii REV. 20, APRIL 2014 CHAPTER 4.0 - REACTOR LIST OF TABLES NUMBER TITLE 4.2-1 Typical Limiting LHGR's for Gadolinia-Urania Fuel Rods (kW/ft) 4.2-2a GE Stress Intensity Limits 4.2-2b FANP Stress Intensity Limits 4.2-3 Conditions of Design Resulting from In-Reactor Process Conditions Combined with Earthquake Loading 4.2-4(a) Data for the 8X8R Fuel Design 4.2-4(b) Data for the GE 8x 8NB(GE 9B) Fuel Design 4.2-4(c) Data for the FANP ATRIUM-9B Fuel Design 4.2-4(d) Data for the FANP ATRIUM-10 Fuel Design 4.2-4(e) Data for the GE14 Fuel Design 4.2-5 Site Fuel Inspection Fu el Inspection Objectives 4.3-1 Maximum Incremental Rod Worths Using BPWS for Each of the Given Rod Groups 4.3-2 Neutron Fluxes Related to Vessel Irradiation 4.3-2a Bounding Neutron Flux and Fluences Related to Reactor Vessel Irradiation 4.3-3 24 Group Multigroup Neutron Flux at the Core Equivalent Radius 4.4-1 Thermal and Hydraulic Design Charac- teristics of the Reactor Core 4.4-2 Void Distribution 4.4-2a Axial Power Distribution Used to Generate Void and Quality Distributions (Typical) 4.4-3 Flow Quality Distribution (Typical) 4.4-4 Core Flow Distribution (Typical) 4.4-5 Typical Range of Test Data 4.4-6 Description of Uncertainties (Deleted) 4.4-7 Reactor Coolant System Geometrical Data 4.4-8 Lengths and Sizes of Safety Injection Lines 4.4-9 Bypass Flow Paths LSCS-UFSAR 4.0-xiii REV. 20, APRIL 2014 CHAPTER 4.0 - REACTOR LIST OF FIGURES AND DRAWINGS FIGURES NUMBER TITLE 4.1-1 Core Arrangement 4.1-2 Core Cell - GE 8X8R Fuel Type 4.1-2a Core Cell - GE 8X8NB Fuel Type 4.1-2b Core Cell - FANP ATRIUM-9B Fuel 4.1-2c Core Cell FANP ATRIUM-10 Fuel 4.1-2d GE14 Lattice Arrangement 4.1-2e GNF2 Lattice Arrangement 4.1-3 Fuel Assembly - (GE 8X8R Shown) 4.1-3a Fuel Assembly GE 8X8NB Fuel 4.1-3b Fuel Assembly FANP ATRIUM 9B Fuel 4.1-3c Fuel Assembly FANP ATRIUM-9B Fuel 4.1-3d Fuel Assembly FANP ATRIUM-10 Fuel 4.1-3e GE14 Fuel Bundle (Typical) 4.1-4 General Electric Control Rod Assembly 4.1-4a General Electric Original Equipment Control Rod Assembly 4.1-4b General Electric Typical Duralife 215 Control Rod Assembly 4.1-4c General Electric Typical Marathon Control Rod Assembly 4.1-5 Steam Separator 4.1-6 Steam Dryer 4.1-7 Steam Dryer Panel 4.2-1 Schematic of Four Bundle Cell Arrangement 4.2-2 Bypass Flow Paths 4.2-3 Fuel Bundle - 8X8R and BP8X8R Fuel Types 4.2-3a Fuel Bundle - GE 8X8EB Fuel Type 4.2-3b Fuel Bundle - GE 8X8NB Fuel Type 4.2-3c Fuel Bundle FANP ATRIUM 9B Type 4.2-3d Fuel Bundle FANP ATRIUM-10 Type 4.2-3e GE14 Fuel Bundle (Typical) 4.2-3f GNF2 Fuel Bundle (Typical) 4.2-4 [Deleted] | |||
4.2-5 Control Rod Velocity Limiter 4.2-5a Fabricast Velocity Limiter 4.2-6 Typical Cladding Temperature vs. Heat Flux - BOL - 8X8R Fuel Type 4.2-7 Typical Cladding Temperature vs. Heat Flux - Late Life- 8X8R Fuel Type 4.2-8 Typical Energy Release as a Function of Time 4.3-1 Initial Core Loading Map (Deleted) 4.3-1a Unit 1 Cycle 5 Core Loading Map (Deleted) | |||
LSCS-UFSAR 4.0-xiv REV. 20, APRIL 2014 FIGURES (Cont'd) | |||
NUMBER TITLE 4.3-1b Unit 2 Cycle 5 Core Loading Map (Deleted) 4.3-2 K as a Function of Exposure at Various Void Fractions, High Enrichment, Dominant Fuel Type (Typical) 4.3-3 Atom Fraction as a F unction of Exposure, High Enrichment, Dominant Fuel Type, 40% Voids (Typical) 4.3-4 Fission Fraction as a Function of Exposure, High Enrichment, Dominant Fuel Type, 40% Voids (Typical) 4.3-5 Neutron Generation Time vs. Exposure at 40% Voids (Typical) 4.3-6 Delayed Neutron Fraction vs. Exposure at 40% Voids (Typical) 4.3-7 Variation of Maximum Rod Power as a Function of Exposure for High Enrichment, 40% Voids (Deleted) 4.3-8 Variation of Maximum Rod Power as a Function of Exposure (Deleted) 4.3-9 Variation of Bundle Average Maximum R-Factor as a Function of Bundle Average Exposure for Uncontrolled | |||
High Enriched Bundle (Deleted) 4.3-10 Radial Power Factors (Deleted) 4.3-11 Typical Beginning of Cycle and End of Cycle Core Average Axial Power - 764 Core, BWR/4 and BWR/5 4.3-12 Moderator Void Reactivity C oefficient at EOC-1 Initial Cycle (GE) 4.3-13 Doppler Reactivity Coefficient as a Function of Fuel Exposure and Average Fuel Temperature at an Average Void Content of 40% High Enrichment Initial Cycle (GE) 4.3-14 Cold Shutdown Example of a Margin Curve 4.3-15 Control Rod Assignments for Groups 1 through 4 (Sequence A) 4.3-16 Control Rod Assignments for Groups 5 Through 10 (Sequence A) 4.3-17 Control Rod Assignments for Groups 1 Through 4 (Sequence B) 4.3-18 Control Rod Assignments for Groups 5 Through 10 (Sequence B) 4.3-19 Hot Operating EOC-1 Scram Reactivity 4.3-20 Xenon Reactivity Buildup and Burnout After Shutdown 4.3-21 Radial Power Distribution at 3323 MWt 4.3-21a Azimuthal Fast Flux Distribution 4.3-22 Axial Power Distribution at 3323 MWt 4.3-22a Axial Fast Flux Distribution 4.4-1 Damping Coefficient vs. Deca y Ratio (Second Order Systems) | |||
LSCS-UFSAR 4.0-xv REV. 20, APRIL 2014 FIGURES (Cont'd) | |||
NUMBER TITLE 4.4-2 Hydrodynamic and Core Stability Model 4.4-3 Model Block Diagram with Valve Flow Control 4.4-4 Comparison of Tests Results with Reactor Core Analysis (Deleted) 4.4-5 Core Reactivity Stability (End of Cycle) (Deleted) 4.4-6 10 psi Pressure Regulator Setpoint Step at 51.5% Rated Power (Natural Circulation) 4.4-7 10 Cent Rod Reactivity Step at 51.5% Rated Power (Natural Circulation) 4.4-8 6-inch Water Level Setpoint Step at 51.5% Rated Power (Natural Circulation) 4.4-9 10 psi Pressure Regulator Setpoint Step at 105% Rated Power and 100% Rated Flow 4.4-10 10 Cent Rod Reactivity Step at 105% Rated Power and 100% Rated Flow 4.4-11 10% Load Demand Step at 105% Rated Power and 100% | |||
Rated Flow 4.4-12 6-inch Water Setpoint Step at 105% Rated Power and 100% Rated Flow 4.4-13 10 psi Pressure Regulator Setpoint Step at 68% Power and 50% Rated Flow 4.4-14 10 Cent Rod Reactivity Step at 68% Rated Power and 50% Rated Flow 4.4-15 10% Load Demand Step at 68% Rated Power and 50% Rated Flow 4.4-16 6-inch Water Level Setpoint Step at 68% Rated Power and 50% Rated Flow 4.6-1 Control Rod to Control Rod Drive Coupling 4.6-2 Control Rod Drive Unit 4.6-3 Control Rod Drive Unit (Schematic) 4.6-4 Control Rod Drive Unit (Cutaway) 4.6-5 Control Rod Drive Hydraulic System Process Diagram 4.6-6 Process Data, Control Rod Drive Hydraulic System 4.6-7 Control Rod Drive Hydraulic Control Unit 4.6-8 Control Rod Drive Housing Support LSCS-UFSAR 4.0-xvi REV. 14, APRIL 2002 DRAWINGS CITED IN THIS CHAPTER* | |||
DRAWING* SUBJECT M-97 P&ID Reactor Water Cleanup System, Unit 1 M-100 P&ID Control Rod Drive Hydraulic Piping, Unit 1 M-143 P&ID Reactor Water Cleanup System, Unit 2 M-146 P&ID Control Rod Drive Hydraulic Piping, Unit 2 | |||
* The listed drawings are included as "Gen eral References" only; i.e., refer to the drawings to obtain additional detail or to obtain background information. These drawings are not part of the UFSAR. They are controlled by the Controlled Documents Program. | |||
LSCS-UFSAR 4.2-1 REV. 20, APRIL 2014 4.2 FUEL SYSTEM | |||
====4.2.1 Design==== | |||
Bases | |||
This section and its subsections were written to describe the design basis consideration used in the design of th e GE initial core and reload fuel. | |||
Detailed descriptions of the design basis co nsiderations used in the design of the AREVA reload fuel can be found in Reference 46 and 49. If a AREVA reference contains the equivalent information as what is being presented for GE, that reference is provided. | |||
4.2.1.1 Safety Design Bases The fuel assembly is designed to ensure, in conjunction with the core nuclear characteristics (Section 4.3), the core thermal and hydraulic characteristics (Section 4.4), the plant equipment characteristics and the instrumentation and protection system, that fuel damage does not result in the release of radioactive materials in excess of the guideline values of 10 CFR 20, 50, and 100. | |||
The mechanical design process emphasizes that: | |||
: a. the fuel assembly provides substantial fission product retention capability during all potential operational modes, and | |||
: b. the fuel assembly provides sufficient structural integrity to prevent operational impairment of any reactor safety equipment. | |||
Assurance of the design basis considerations is provid ed by the following fuel assembly capabilities: | |||
: a. Pressure and temperature capabilities The fuel assembly and its components are capable of withstanding the predicted thermal, pressure, and mechanical interaction loadings occurring during startup testing, normal operation, and abnormal operation without impairment of operational capability. | |||
LSCS-UFSAR 4.2-2 REV. | |||
13 b. Handling capability The fuel assembly and each component thereof is capable of withstanding loading predicted to occur during handling without impairment of operational capability. | |||
: c. Earthquake loading capability (OBE) | |||
The fuel assembly and each component thereof is capable of sustaining incore loading predicted to occur from an operating basis earthquake (OBE), when occurring during normal operating conditions without impairment of operational | |||
capability. | |||
: d. Earthquake loading capability (SSE) | |||
The fuel assembly and each component thereof is capable of sustaining incore loading predicted to occur from a safe | |||
shutdown earthquake (SSE) when occurring during normal operation without: | |||
: 1. exceeding deflection limits which allow control rod insertion, and | |||
: 2. fragmentation or severance of any bundle component. | |||
: e. Accident capability The capability of the fuel assembly to withstand the control rod drop accident, the pipe breaks inside and outside containment accidents, the fuel handling accident, and one recirculation pump seizure accident, is determined by analysis of the specific event. The ability of the fuel assemb ly to provide the preceding capabilities is evaluated by one or more of the following: | |||
: a. design ratios developed by utilizing continually evolving, state-of-the-art numerical analysis techniques (Subsection 4.2.1.2.15); | |||
: b. analytical procedures based on classical methods (Subsection 4.2.1.2.5); and | |||
: c. experience and testing (Subsection 4.2.3.2). | |||
LSCS-UFSAR 4.2-3 REV. 20, APRIL 2014 For the initial reloads of the AREVA ATRIUM-9B and ATRIUM-10 fuel, the control rod drop accident, the pipe breaks inside and outside containment accidents, and the fuel handling accident were all evaluated for the fuel assembly's capability to withstand their effects. The recirculation pump seizure event was not analyzed by AREVA for their ATRIUM-9B or ATRIUM-10 fuel; it was dispositioned as bounded by the LOCA accident. The control rod drop accident is evaluated each cycle for both AREVA and GE reloads. | |||
4.2.1.2 Power Generation Design Basis | |||
The fuel assembly is designed to ensure, in conjunction with the core nuclear characteristics, the core thermal and hydr aulic characteristics, the plant equipment characteristics and the instrumentation and protection system, that fuel damage limits will not be exceeded during either planned operation or abnormal operational transients caused by any single equipment malfunction or single operator error. | |||
4.2.1.2.1 Material Selection | |||
The basic materials used in fuel assemb lies are Zircaloy, natural zirconium, Type 304 stainless steel, Inconel-X, and ceramic uranium dioxide and gadolinia. | |||
These materials have been shown from earlier reactor experience to be compatible with BWR conditions and to retain their design function capability during reactor operation. Additional information on material properties is referenced in Reference 41. | |||
4.2.1.2.2 Effects of Irra diation and Fuel Swelling Irradiation affects both fuel and cladding material properties. The effects include increased cladding strength and reduced cladding ductility. In addition, irradiation in a thermal reactor environment results in the buildup of both gaseous and solid fission products within the UO 2 fuel pellet which tend to increase the pellet diameter, i.e., fuel irradiation swelling. Pellet internal porosity and pellet-to-cladding gap have been specifie d such that the thermal expansion and irradiation swelling are accommodated throughout life. The irradiation swelling model is based on the NRC approved methodology as described in Reference 41. | |||
Observations and calculations based on this refined model for relative UO 2 fuel/cladding expansion indicate that the as-fabricated UO 2 pellet porosity is adequate (without pellet dishing) to accommodate the fission-product-induced UO 2 swelling out to expected exposures. | |||
The primary purpose of the gap between the UO 2 fuel pellet and Zircaloy cladding is to accommodate differential diametral expansion of fuel pellet and cladding and, thus, preclude the occurrence of excessive gross diametral cladding strain. A short time after reactor startup, the fuel cracks radially and redistributes out to the LSCS-UFSAR 4.2-4 REV. 20, APRIL 2014 cladding. Experience has shown, however, the gap volume remains available in the form of radial cracks to accommodate gross diametral fuel expansion. | |||
The value of thermal conductance used in BWR fuel design is derived from postirradiation data on exposed fuel with an initial pellet-to-cladding gap which is significantly larger than that employed in the General Electric fuel design. | |||
Axial ratcheting of fuel cladding is no t considered in BWR fuel rod design. | |||
Prototypical fuel rods have been operated in the Halden test reactor with axial elongation transducers. No significan t axial ratcheting has been observed (Reference 5). | |||
Fission product buildup also tends to ca use a slight reduction in fuel melting temperature. The melting point of UO 2 is considered to decrease with irradiation based on data from Reference 6. | |||
In the temperature range of interest (500 C), the fuel thermal conductivity is not considered to be significantly affected by irradiation as reported in Reference 7. | |||
A small fraction of the gaseous fission products is released from the fuel pellets to produce an increase in fuel rod internal gas pressure as discussed further in Subsection 4.2.1.2.7. In general, such irradiation effects on fuel performance have been characterized by available data and are considered in determining the design features and performance. Thus, the irradiation effects on fuel performance are inherently considered when determining whether or not the stress intensity limits and temperature limits are satisfied. | |||
In Reference 49, AREVA states that the BWR evaluation models for densification and swelling are included in the NRC approved fuel performance codes, References 50 and 51. | |||
4.2.1.2.3 Fuel Densification | |||
4.2.1.2.3.1 GE Fuel Fuel performance calculations that account for some specific effects of fuel densification have been performed with an approved version of the General Electric analytical model as described in Reference 41. The approved analytical model incorporates time-dependent fuel densification, time-dependent gap closure and | |||
cladding creepdown for the calculation of gap conductance. Other fuel performance predictions, such as cladding response, ar e also calculated. Cladding collapse has not been observed in boiling water reactor fuel rods, but its theoretical occurrence is calculated with the NRC approved methodolo gy as described in Reference 41. All of the fuel cladding used at LSCS has been shown not to collapse during the life of the fuel. | |||
LSCS-UFSAR 4.2-5 REV. 20, APRIL 2014 4.2.1.2.3.2 AREVA Fuel | |||
Fuel densification and swelling are limited by design criteria specified for fuel temperature, cladding strain, cladding collapse, and internal pressure criteria (Reference 49). | |||
Creep collapse of the cladding and the subsequent potential for fuel failure is avoided in the AREVA fuel system design by eliminating the formation of axial gaps. The maximum cladding circumferential creep and ovalization consistent with the time of maximum densification is computed during a creep collapse evaluation to demonstrate that no axial gaps are present. The evaluation must show that the pellet column is compact at the burnup of maximum densification (approximately 6000 MWd/MTU). The internal plenum spring provides an axial load on the fuel stack that is sufficient to assist in the closure of any gaps caused by handling, shipping, and densification. Evaluation of cladding creep stability in the unsupported condition is performed cons idering the compressive load on the cladding due to the difference between primary system pressure and the fuel rod internal pressure. AREVA fuel is designed to minimize the potential for the formation of axial gaps in the fuel and to minimize clad creepdown which would | |||
prevent the closure of axial gaps or allow creep collapse (Reference 49). | |||
4.2.1.2.4 Incipient UO 2 Center Melting | |||
4.2.1.2.4.1 GE Fuel | |||
The fuel rod is evaluated to ensure that fuel rod failure due to fuel melting is not expected to occur during normal steady-state operation. Incipient center melting is not expected to occur in fresh GE UO 2 fuel rods at the linear heat generation rate (LHGR) described in Reference 41 and Reference 59. The LHGR values for incipient center melt decrease slightly with burnup. The effect of gadolinia concentration and fuel exposure on the LHGR at calculated incipient center melting is also described in Reference 41. | |||
4.2.1.2.4.2 AREVA Fuel | |||
Fuel failure from the overheating of the fuel pellets is not allowed. The centerline temperature of the fuel pellets must rema in below melting during normal operation and anticipated operational occurrences. | |||
The melting point of the fuel includes adjustments for burnup and gadolinia co ntent. AREVA establishes steady state and transient design LHGR limits for ea ch fuel type which protect against centerline melting. These LHGR limits ar e appropriate for normal operation and anticipated operational occurrences throug hout the design lifetime of the fuel (Reference 49). | |||
LSCS-UFSAR 4.2-6 REV. 18, APRIL 2010 4.2.1.2.5 Maximum Allowable Stresses The strength theory, terminology, and stress categories presented in the ASME Boiler and Pressure Vessel Code, Section III, are used as a guide in the mechanical | |||
design and stress analysis of the reactor fuel rods. The mechanical design is based on the maximum shear stress theory for combined stresses. The equivalent stress intensities used are defined as the difference between the most positive and least | |||
positive principal stresses in a triaxial field. Thus, stress intensities are directly comparable to strength values found from tensile tests. Table 4.2-2a and b present a summary of the basic stress intensity limits that are applied for Zircaloy-2 cladding for both GE fuel and AREVA ATRIUM-9B and ATRIUM-10 fuel. | |||
4.2.1.2.5.1 GE Fuel In this analysis of BWR Zircaloy-clad UO 2 pellet fuel, continuous functional variations of mechanical properties with exposure are not employed since the irradiation effects become saturated at very low exposure. At beginning of life, the cladding mechanical properties employ ed are the unirradiated values. At subsequent times in life, the cladding mechanical properties employed are the | |||
saturated irradiated values. The only exception to this is that unirradiated mechanical properties are employed above the temperatures for which irradiation effects on cladding mechanical properties are assumed to be annealed out. It is significant that the values of clad yield strength and ultimate tensile strength employed represent the approximate lower bound of data on cladding fabricated by General Electric, i.e., approximately two standard deviations below the mean value. | |||
In this analysis the calculated stress and the yield strength or ultimate strength are combined into a dimensionless quantity called the design ratio. This quantity is the ratio of calculated stress intensity to the design stress limit for a particular stress category. The design stress limit for a particular stress category is defined as a fraction of either the yield strength or ultimate strength, whichever is lower. Thus, the design ratio is a measure of the fracti on of the allowable stress represented by the calculated stress. | |||
Analyses are performed to show that the stress intensity limits given in Table 4.2-2a and b are not exceeded during continuous operation with linear heat generation rates up to the design operating limit, or during transient operation above the design operating limit. Stresses due to external coolant pressure, internal gas pressure, thermal effects, spac er contact, flow-induced vibration, and manufacturing tolerances are considered. Cladding mechanical properties used in stress analyses are based on test data of fuel rod cladding for the applicable temperature. | |||
Fuel rods are evaluated to assure that the fuel will not fail due to stresses or strains exceeding the fuel rod mechanical capability. The analysis performed is described in Reference 41. | |||
LSCS-UFSAR 4.2-7 REV. 20, APRIL 2014 4.2.1.2.5.2 AREVA Fuel AREVA requires compliance with both Standard Review Plan criteria for pellet/cladding interaction for steady state and transient conditions over the lifetime of the fuel. The first one is that transient - induced deformations must be less than 1% uniform cladding strain. The second is that fuel melting cannot occur. Compliance with the fuel melting criteria is discussed in Section 4.2.1.2.4.2. | |||
The design basis for the fuel cladding stress limits is that the fuel system will not be damaged due to fuel cladding stresses. Conservative limits are derived from the ASME Boiler Code, Section III, Article-2000; and the specified 0.2% offset yield strength and ultimate strength for Zircaloy (Reference 49). | |||
4.2.1.2.6 Capacity for Fission Gas Inventory | |||
The available fission gas retention volume is determined based upon the following assumptions: | |||
: a. Nominal as-built plenum length and cladding inside diameter. | |||
: b. Maximum expected fuel-cladding differential expansion. | |||
: c. No credit for fuel-cladding annulus (gap). | |||
: d. The "net" volume is corrected for the volume of the components contained within the fuel rod plenum. | |||
4.2.1.2.7 Maximum Internal Gas Pressure | |||
Fuel rod internal pressure is due to the helium which is backfilled during rod fabrication, the volatile content of the UO 2 , and the fraction of gaseous fission products which are released from the UO | |||
: 2. Nominal tolerances are assumed in defining the hot plenum volume used to compute fuel rod internal gas pressure. | |||
4.2.1.2.7.1 GE Fuel The fuel rod internal pressure is calculat ed using the perfect ga s law (P = NRT/V). | |||
A quantity of 1.35 milligram-moles of fission gas is pr oduced per MWd of power production. In fuel rod pressure and stress calculations, fission gas release is calculated as per the NRC approved meth odology as described in Reference 41. This fission gas release model has been demonstrated by experiment to be conservative over the complete range of design temperature and exposure conditions (References 4 and 41). The calculated maximum fission gas release fraction in the highest design power density rod is less than 25%. This calculation is conservative because it assumes the worst peaking LSCS-UFSAR 4.2-8 REV. 18, APRIL 2010 factors applied constantly to this rod. The percentage of total fuel rod radioactivity released to the rod plenum is much less than 25% because of radioactive decay during diffusion from the UO | |||
: 2. | |||
Creepdown and creep collapse of the plenum are not considered because significant creep in the plenum region is not expected. The fuel rod is designed to be free-standing throughout its lifetime. The temperature and neutron flux in the plenum region are considerably lower than in the fueled region, thus the margin to creep collapse is substantially greater in the plenum. Direct measurements of irradiated fuel rods have given no indication of significant creepdown of the plenum. | |||
The fuel rod is evaluated to assure that the effects of rod internal pressure during normal steady state operation will not result in fuel failure. The analysis is further described in Reference 41. | |||
4.2.1.2.7.2 AREVA Fuel | |||
To prevent unstable thermal behavior and to maintain the integrity of the cladding, AREVA limits the maximum internal rod pressure relative to system pressure to avoid significant hydride reorientation during cooldown conditions or depressurization conditions. When the fu el rod internal pressure exceeds system pressure, the pellet-cladding gap has to remain closed if it is already closed or it should not tend to open for steady or increasing power conditions. Outward circumferential creep which may cause an increase in pellet-to-cladding gap must be prevented since it would lead to higher fuel temperature and higher fission gas release. The maximum internal pressure is also limited to protect embrittlement of | |||
the cladding caused by hydride reorientation during cooldown and depressurization conditions (Reference 49). | |||
4.2.1.2.8 Internal Pressure and Cladding Stresses During Normal Conditions | |||
The internal pressure is applied coincident with the applicable coolant pressure to compute the resulting cladding stresses, which, combined with cladding stresses from other sources, must satisfy the stress limits described in Subsection 4.2.1.2.5. | |||
4.2.1.2.9 Cycling and Fatigue Limits | |||
4.2.1.2.9.1 GE Analysis | |||
The fatigue analysis utilizes the linear cumulative damage rule (Miner's hypothesis) as documented in "Fatigue Design Basis for Zircaloy Components" (Reference 12). The fatigue analysis is based on the estimated number of temperature, pressure, and power cycles. Th e fuel assembly and fuel rod cladding are evaluated to ensure that strains due to cyclic loadings will not exceed the fatigue capability. | |||
LSCS-UFSAR 4.2-9 REV. 20, APRIL 2014 4.2.1.2.9.2 AREVA Analysis Cycle loading associated with relatively large changes in power can cause cumulative damage which may eventually lead to fatigue failure. Therefore, AREVA requires that the cladding not exceed a cumulative fatigue usage factor of 0.67. The O'Donnell and Langer fatigue curves are used in the analysis. These fatigue curves have been adjusted to incorporate the recommended '2 or 20' safety factor. This safety factor reduces the stress amplitude by factor of 2 or reduces the number of cycles by a factor of 20, whichever is more conservative. The fatigue curves provide the maximum allowed number of cyclic loading for each stress amplitude. The fatigue usage factor is the number of expected cycles divided by the number of allowed cycles. The total cladding usage factor is the sum of the individual usage factors for each duty cycle (Reference 49). | |||
4.2.1.2.10 Deflections | |||
4.2.1.2.10.1 GE Evaluation | |||
The operational fuel rod deflections considered are the deflections due to: | |||
: a. manufacturing tolerances, b. flow-induced vibration, c. thermal effects, and | |||
: d. axial load. | |||
There are two criteria that limit the magnitud e of these deflections. | |||
One criterion is that the cladding stress limits must be satisfied; the other is that the fuel rod-to-rod and rod-to-channel clearances must be suffi cient to allow free passage of coolant water to all heat transfer surfaces. The fuel rod is evaluated to ensure that fuel rod bowing does not result in fuel failure due to boiling transition. | |||
4.2.1.2.10.2 AREVA Evaluation | |||
Differential expansion between the fuel rods, and lateral thermal and flux gradients can lead to lateral creep bow of the rods in the spans between sp acer grids. This lateral creep bow alters the pitch between the rods and may affect the peaking and | |||
local heat transfer. The AREVA design basis for fuel rod bowing is that lateral displacement of the fuel rods shall not be of sufficient magnitude to impact thermal margins. Extensive post-irradiation exam inations have confirmed that such rod LSCS-UFSAR 4.2-10 REV. 18, APRIL 2010 bow has not reduced spacing between adjacent rods by more than 50%. The potential effect of this bow on thermal margins is negligible. Rod bow at extended burnup does not affect thermal margins du e to the lower power achieved at high exposure (Reference 49). | |||
4.2.1.2.11 Flow-Induced Fuel Rod Vibrations Flow-induced fuel rod vibrations depend primarily on flow velocity and fuel rod geometry. The stress levels resulting from the vibrations are negligibly low and well below the endurance limit of all affected components. This phenomenon is further described in GE References 13 and 41. | |||
Reference 47 discusses the AREVA calculatio ns for flow induced vibrations. Vibrational stresses due to flow induce d vibrations are calculated with the Paidoussis analysis which assumes: | |||
: 1) The structural stiffness of the rod is due to cladding only. | |||
: 2) The sections of the fuel rod be tween spacers and/or tie plate supports are modelled structurally as a simple beam with pinned ends. | |||
: 3) Flow velocity, viscosity, and virtual mass for the amplitude calculations are evaluated as suggested by Paidoussis. | |||
4.2.1.2.12 Fretting Corrosion | |||
Fretting wear has been considered in establishing the fuel mechanical design basis. | |||
Specific GE fuel designs described in Reference 41 have been incorporated to eliminate fretting wear. Tests of these designs have been conducted both | |||
out-of-reactor as well as in-reactor prior to application in a complete reactor core basis. All tests and post-irradiation exam inations have indicated that fretting corrosion does not occur. Post-irradiation examination of many fuel rods indicates only minor fretting wear. Excessive wear at spacer contact points has never been observed with the current spacer configuration. Additional information on testing relative to fretting wear is contained in Reference 41. AREVA discusses fretting wear in Reference 49. | |||
4.2.1.2.13 Seismic Loadings | |||
The fuel is analyzed for loading in the reactor resulting from seismic accelerations. The fuel seismic design basis is the design basis presented in References 15, 17 and 41 for GE fuel. The fuel seismic design basis for AREVA fuel is presented in Reference 49. Reference 48 verifies that the AREVA seismic criterial were met for a particular reload. | |||
LSCS-UFSAR 4.2-11 REV. 18, APRIL 2010 4.2.1.2.14 Chemical Properties of Cladding and Fuel Material | |||
The fuel material, fuel rod, pellets, and cladding are discussed generally in Subsections 4.2.2.2 through 4. | |||
====2.2.5. Testing==== | |||
and inspection of fuel is covered in Subsection 4.2.4. Reference 41 reports the specific fuel parameters of the fuel used for LSCS. Reference 19 presents the BWR fuel experience through September 1974. | |||
Reference 42 represents later BWR fuel ex perience. References 46, 48, and 49 report specific fuel parameters for AREVA fuel. | |||
4.2.1.2.15 Design Ratios | |||
Design ratios are defined by the following relationship: D.R. = A/L where D.R. is the design ratio, L is the limiting parameter value, and A is the actual parameter value. Design ratios of less than one are demonstrated for component parameters influenced by loading conditions which may affect the structural or dimensional integrity of the fuel assemb ly or any component thereof. | |||
4.2.1.2.15.1 Limiting Parameter Values The following information is based on GE methodology. For a discussion on AREVA methodologies see Reference 46, 47, 48, 49 and 55. | |||
4.2.1.2.15.1.1 Normal and Upset Design Conditions | |||
Limiting parameter values for each component are determined in the following manner as defined by Table 4.2-3: | |||
: a. For stress resulting from mean value or steady-state loading, the limiting value is determined by consideration of the material | |||
0.2% offset yield strength or the equivalent strain, as established at operating temperature. | |||
: b. For stress resulting from load cycling, limiting parameter values are determined from fatigue limits. | |||
: c. For stress resulting from loading of significant duration, the limiting parameter is determined from consideration of stress rupture as defined by the Larson-Miller parameter. If metal temperatures are below the level of applicability of stress rupture for the material or if the yield strength is more limiting then the limiting value of stress is determined from consideration of the material 0.2% offset yield strength or the equivalent strain, as establis hed at operating temperatures. | |||
LSCS-UFSAR 4.2-12 REV. 15, APRIL 2004 | |||
: d. Where stress rupture and fatigue cycling are both significant, the following limiting condition is applied: I = 1 to n I = 1 to m | |||
: e. Critical instability loads shall be derived from test data when available or from analytical methods when applicable test data are not available. | |||
: f. Deflection limits are those values of component deformation which could cause an undesirable event such as impairment of control rod movement or an excessive leakage flow rate. | |||
4.2.1.2.15.1.2 Emergency and Faulted Design Conditions Limiting parameter values are determined in the following manner as defined by Table 4.2-3: | |||
: a. Stress limits are determined from consideration of the ultimate tensile strength or equivalent strain of the material, as | |||
established at operating temperatures. | |||
: b. Critical instability loads are determined from test data when available or from analytical methods when applicable test data is not available. | |||
: c. Deflection limits are those values of deformation that if occurring could lead to a more serious consequence such as | |||
prevention of control rod insertion. | |||
4.2.1.2.15.2 Actual Parameter Values | |||
The following information is based on GE methodology. For a discussion on FANP methodologies see References 46, 47, 48, 49 and 55. | |||
Actual parameter values are determined from the following considerations: | |||
: a. Effective stresses are determined at each point of interest using the theory of constant elastic strain energy of distortion: | |||
1stressat cycles allowable cycles ofnumber actual stressat time allowable stressat timeactual LSCS-UFSAR 4.2-13 REV. 14, APRIL 2002 Stress concentration may be applied only to the alternating stress component. | |||
: b. Design values of instability loads are scaled up to allow for uncertainty in manner of load application, variation in modulus of elasticity, and difference between the actual case and the | |||
theoretical one. | |||
: c. Calculated values of deflection for comparison with deflection limits may be based on the resulting permanent set after load removal if load removal occurs before damage may result. | |||
4.2.1.2.16 Fuel Assembly Limits The design limits applicable to each component are discussed in the following paragraphs. In order to provide a fulle r understanding of how the limits will be applied, a functional description of each component and a discussion of the loadings on each component are provided. | |||
The general configuration of the fuel asse mbly and the detailed configurations of the assembly components are the result of the evolutionary change in customer, performance, manufacturing and serviceability requirements and the experience obtained since the initial design conception. In general, the experience obtained in prior fuel designs is relied upon very heavily to qualify particular component configurations for production fuel applic ation. More sophisticated analytical techniques are continually being developed and applied to fuel design. | |||
4.2.1.2.16.1 Fuel Rods | |||
A discussion of the mechanical analysis of the fuel rod and the appropriate stress intensity limits was provided in Subsection 4.2.1.2.5. In additi on, a fuel rod fatigue analysis is performed as descri bed in Subsection 4.2.1.2.9. | |||
As explained in Subsection 4.2.3.21, significant fuel rod bowing due to binding at the spacers is not expected to occur. Other contributors to rod bowing during normal operation and transients are manufacturing tolerances and thermal gradients. These factors are considered in the design. | |||
4.2.1.2.16.2 Fuel Spacer | |||
The primary function of the fuel spacer is to provide lateral support and spacing of the fuel rods, with consideration of thermal-hydraulic performance, fretting wear, strength, neutron economy, and producibility. | |||
The mechanical loadings on the spacer structure during normal operation and transients result from the rod positioning spacer spring forces and from local LSCS-UFSAR 4.2-14 REV. 18, APRIL 2010 loadings at the water rod-spacer positioning device. During a seismic event, the spacer transmits the lateral acceleration loadings from the fuel rods into the channel, while maintaining the spatial relationship between the rods. | |||
As noted above, the spacer represents an optimization of a number of considerations. Thermal-hydraulic development effort has gone into designing the particular configuration of the spacer parts. The resu ltant configurations give enhanced hydraulic performance. Extensive flow testing has been performed employing prototypical spacers to define single-phase and two-phase flow characteristics. Details of the mechanical design of the spacers used at LSCS can be found in Reference 41 for GE fuel and References 46, 47, 48 and 55 for AREVA fuel. | |||
4.2.1.2.16.3 Water Rods or Water Channel The main mechanical function of the water ro d(s) (or water channel) is to maintain the axial position of the fuel spacers. For the ATRIUM-10 Fuel, the water channel also provides the structural connection between the upper and lower tie plates. | |||
Differential thermal expansion between fuel ro ds and the water rods (or water channel) can introduce axial loadings into the water rod (or water channel) through the frictional forces between the fuel rods and the spacers. This differential growth is considered in the design process as discussed in Referenc e 41 for GE fuel and References 46 and 55 | |||
for AREVA fuel. | |||
The water rods or water channel provide flow through the center portion of the fuel assembly, thereby providing additional modera tion within the bundle interior. This improves uranium utilization and operational flexibility. | |||
4.2.1.2.16.4 Channel Assurance that the channels maintain their dimensional integrity, strength, and spatial position throughout their lifetime is provided in the following ways: | |||
: a. Dimensional integrity, as related to relaxation of residual forming stresses, is provided through the channel specifications and by qualification of the manufacturing process to these specifications. | |||
The operational experience with channels produced using the current process has demonstrated satisfactory relaxation characteristics (Reference 17). | |||
: b. The performance of the channels currently in operation has shown no tendency for gross inservice deformations, although long-term creep deformation and channel bulg e have been identified as a potential life-limiting phenomenon (References 17 and 45). | |||
: c. Channel material strength is assured through the material specification of yield and ultimate strength. Quality LSCS-UFSAR 4.2-15 REV. 18, APRIL 2010 measurements are made to show compliance with this specification. Irradiation subs tantially increases the material strength. | |||
: d. Mechanical integrity of the channel (that is, assurance that the channel will maintain its spatial position and integrity) is provided by designing the channel to the limits stated in Subsection 4.2.2.6 and item e following. The design limits used are based on the unirradiated strength of the material, thereby providing substantial material strength margin throughout most | |||
of the life of the channel. | |||
: e. During normal and transient operation, the channel is subjected to differential pressure loadings. The pressure loadings are evaluated to ensure the channel will not experience excessive deflection and subsequent channel wear. | |||
4.2.1.2.16.5 Tie Plates | |||
The upper and lower tie plates serve the functions of supporting the weight of the fuel and positioning the rod ends during all phases of operation and handling. The loading on the lower tie plate during operation and transients comprise the fuel weight, the weight of the channel, and the forces from the expansion springs at the top of the fuel rods. The loading of the upper tie plate is the expansion springs' force. The expansion springs permit differential expansion between the fuel rods without introducing high axial forces into the rods. | |||
Most of this loading arises from the weight of the fuel rods and the channel, which are not cyclic loadings. During accidents the tie plates are subjected to the normal operational loads plus the blowdown and seismic loadings. During handling, the tie plates are subjected to acceleration and impact loading. The stress design limit for the tie plates for all phases of operat ion and normal handling is discussed in Reference 41 for GE fuel. Reference 48 contains information regarding the upper and lower tie plate loads for AREVA ATRIUM-9B and ATRIUM-10 fuel. | |||
The ATRIUM-10 fuel design includes AREVA's FUELGUARD debris resistant lower tie plate. This design was chosen for two reasons: 1) to address the main cause of BWR fuel failures over the past few years - de bris induced fuel rod fretting; and 2) to reduce overall assembly pressure drop to be tter assure adequate core flow is available for reactor power maneuvering. | |||
The FUELGUARD lower tie plate design consis ts of a parallel array of blades with curved portions in the middle. The blades are arranged so that there is no line of sight through the grid thus preventing the passage of long narrow objects and objects larger than the pitch of the blades. The blades for the FUELGUARD on the | |||
ATRIUM-10 are brazed in position. | |||
LSCS-UFSAR 4.2-15a REV. 20, APRIL 2014The GE14 fuel design is assembled with a debris filter Lower Tie Plate (LTP) as standard equipment. The debris filter LTP increases the single phase pressure drop by approximately 0.3 psi over the non-debris filter LTP. The debris filter LTP has an underlying grid that screens out the debris and mitigates the debris related fuel rod failures by reducing the size of debris that can enter the fuel assembly. More detailed description of the debris filt er LTP can be found in Reference 58. | |||
The GNF2 fuel design is assembled with the Defender LTP as standard equipment. | |||
This LTP provides essentially the same si ngle-phase pressure drop as the GE14 debris filter LTP, while preventing smaller debris from entering the fuel assembly. | |||
Additional description of the Defender LTP can be found in Reference 59. | |||
4.2.1.2.17 Reactivity Control Assembly and Burnable Poison Rods 4.2.1.2.17.1 Safety Design Bases for Reactivity Control | |||
The limiting criteria for shutdown reactivity margins are given in Subsection 4.3.1.1 as items a and f. The cold-clean shutdown margin is shown in Figure 4.3-14 for the initial cycle of Units 1 and 2. Th e presence of the burnable poison Gd 2 O 3 is apparent in the curve shape as keff rises concurrent with poison depletion. The negative reactivity worth of the gadolinia-containing fuel rods decreases in a nearly linear manner so that it closely matches the depletion of fissile material. The curve LSCS-UFSAR 4.2-16 REV. 14, APRIL 2002 shown in Figure 4.3-14 is typical for most cycles, although differences will exist from cycle to cycle. | |||
The reactivity control mechanical design includes control rods and gadolinia burnable poison in selected fuel rods within fuel asse mblies and meets the following safety design bases. | |||
: a. The control rods have sufficie nt mechanical strength to prevent displacement of their reactivity control material. | |||
: b. The control rods have sufficient strength and are so designed as to prevent deformation that could inhibit their motion. | |||
: c. Each control rod has a device to limit its free-fall velocity sufficiently to avoid damage to the nuclear system process barrier by the rapid reactivity increase resulting from a free-fall of one control rod from its fully inserted position to the position where the drive was withdrawn. | |||
4.2.1.2.17.1.1 Specific Design Characteristics The acceptability of the control rod and control rod drive under scram loading condition is demonstrated by functional testing instead of analysis or adherence to formally defined stress limits. The results of such testing are given in Reference 10. | |||
The basis of the mechanical design of the control rod blade clearances is that there is no interference which will restrict the passage of the control rod blade. | |||
Mechanical insertion requirements during normal operation are selected to provide adequate operability and load following capability, and are able to control the reactivity addition resulting from burnou t of peak shutdown xenon at 100% power. | |||
Scram insertion requirements are chosen to provide sufficient shutdown margin to meet all safety criteria for plant operational transients (Chapter 15.0). | |||
The selection of materials for use in the control rod design is based upon their in-reactor properties. The irradiated properties of Type 304 austenitic stainless | |||
steel, 316 stainless steel and CF3 which compr ise the major portion of the assembly, B 4 C powder, hafnium Inconel-X, and stellite are well known and are taken into account in establishing the mechanical design of the control rod components. The basic cruciform control rod design and materials have been operating successfully in all GE reactors. No problems associat ed with component materials have been observed. | |||
The radiation effects on B 4C powder include the release of gaseous products, and the B 4C cladding is designed to sustain the resulting internal pressure buildup. The corrosion rate and the physical properties, e.g., density, modulus of elasticity, LSCS-UFSAR 4.2-17 REV. 14, APRIL 2002dimensional aspects, etc., of austenitic stainless steel, 316 stainless steel, CF3 and Inconel-X are essentially unaffected by the irradiation experienced in the BWR reactor core. The effects upon the mech anical properties, i.e., yield strength, ultimate tensile strength, percent elon gation, and ductility on the Type 304 stainless steel cladding also are well known and are considered in mechanical design. Visual examinations of control rods whic h have been subjected to high exposure rates have disclosed no significant material degradation (Reference 11). | |||
Rod positioning increments (notch lengths) are selected to provide adequate power shaping capability. The combination of rod speed and notch length must also meet the limiting reactivity addition rate criteria. | |||
For all LaSalle cores, supplementary reactivity control must be provided in such a way that the high initial keff can be compensated throughout the active core. Gadolinia containing fuel rods are used in normal fuel assemblies to attain this objective. Some assemblies contain mo re gadolinia than others to improve flattening both in the radial and axial directions. | |||
The gadolinia is uniformly distributed in the UO 2 pellet and forms a solid solution. | |||
The presence of the high cross section gadolinium isotopes results in a relatively low heat generation rate in those rods (this heat generation rate is also adjusted by the position of the gadolinia rods within the fu el assembly). During a fuel cycle, the gadolinia essentially burns out thus enab ling a progressive increase in rod power and a concurrent increase in net assembly power. At later stages of fuel exposure the power of the gadolinia-urania fuel rods decreases. | |||
Precise quality control measures are utilized during the manufacture of gadolinia bearing UO 2 pellets and also during the assembly of these fuel pellets into fuel rods. Special procedures assure accurate placem ent and quantity control for placement of gadolinia rods. | |||
4.2.1.2.18 Surveillance Program See Subsection 4.6.3.2 for information regarding the control rod surveillance program. The surveillance tests for the control rod drive system include an acceptance test, preinstallation test, operational test prior to startup, and tests during startup. Specific surveillance tests are performed following a refueling outage when core alterations are made, to demonstrate that the core can be made subcritical with a margin of 0.0038 k at any time in the subsequent fuel cycle with the strongest operable control rod fully withdrawn and all other operable control rods fully inserted. | |||
LSCS-UFSAR 4.2-18 REV. 20, APRIL 20144.2.2 Description and Design Drawings 4.2.2.1 Core Cell | |||
A core cell consists of a control rod and the four fuel assemblies which immediately surround it (Figure 4.2-1). Each core cell is associated with a four-lobed fuel support piece. Around the outer edge of the core, certain fuel assemblies are not immediately adjacent to a control rod and are supported by individual peripheral fuel support pieces. | |||
The top guide is an "egg-crate" structur e of stainless steel bars which form a four-bundle cell. The four fuel assemblies are lowered into this cell and, when seated, springs mounted at the tops of th e channels force the channels into the corners of the cell such that the sides of the channels contact the grid beams (Figure 4.2-1). | |||
4.2.2.2 Fuel Assembly | |||
A fuel assembly consists of fuel bundle and the channel whic h surrounds it (Figure 4.1-3). The fuel assemblies are arranged in the reactor core to approximate a right circular cylinder inside the core shroud. Each fuel assembly is supported by a fuel support piece and the top guide. A summary of nuclear fuel data for the GNF (formerly GE) 8x8R, 8x8NB, GE14, and GNF2 fuel designs are presented in Tables 4.2-4a, 4.2-4b, 4.2-4e, and 4.2-4f, respectively. A summary of nuclear fuel data for the AREVA ATRIUM-9B and ATRIUM-10 fuel designs are presented in Tables 4.2-4c and 4.2-4d, respectively. Other pertinent data are presented in References 41, 44, 46, 48, 49, 58, and 59. | |||
Beginning with LaSalle Unit 2 Cycle 13 and co ntinuing in subsequent cycles of Unit 2, AREVA Lead Fuel Assemblies (LFAs) ar e inserted into non-limiting core locations for demonstration purposes. The LFA program consists of eight Lead Fuel Assemblies of ATRIUM 10XM. | |||
The ATRIUM 10XM fuel bundle shares a ge neral geometry with the ATRIUM 10, consisting of a 10x10 fuel rod lattice with a square internal water channel. | |||
Variations from ATRIUM 10 include water channel crowns, a change in spacer material, addition of one spacer for a total of nine, larger fuel rod diameter, longer active fuel length, additional uranium mass, and in selected fuel rods doped fuel pellets and re-crystallized cladding. ATRIUM 10XM is described in Mechanical Design Report for LaSalle Unit 2 Cycle 13 Atrium 10XM Lead Test Assemblies , ANP-2756P, Revision 0, October 2008. | |||
LSCS-UFSAR 4.2-18a REV. 20, APRIL 20144.2.2.3 Fuel Bundle The 8x8R and BP8x8R (Figure 4.2-3) fuel bundles contain 62 fuel rods and two water rods which are spaced and supported in a square (8 x 8) array by the lower and upper tie plates. The GE8x8EB fuel de sign (Figure 4.2-3a) provides for the use of up to four water rods. However, the GE 8X8EB fuel bundles at LaSalle have two water rods. The GE8X8NB fuel design (Fig ure 4.2-3b) contains 60 fuel rods and one large centrally located water rod. | |||
The GE14 fuel design (Figure 4.2-3e) is bas ed on a 10x10 array that contains 78 full length rods, 14 part length rods and 2 large water rods that effectively replaced 8 fuel rods. The 14 part length rods terminate just past the top of the fifth spacer. Eight full length rods are used as tie rods. The rods are spaced and supported by the upper and debris filter lower tie plates and eight spacers over the length of the fuel rods. This assembly is encased in an interactive thick corner/thin wall fuel channel. Finger springs control the coolant leakage flow between the debris filter lower tie plate and the channel. | |||
Additional assembly and component description for the GE14 fuel design are provided in Reference 58. | |||
The GNF2 design consists of 92 fuel rods and two large central water rods contained in a 10x10 array. The two water rods encompass eight fuel rod positions. Eight of the fuel rods terminate at approximately two-thirds of the bundle length and are designated as long part length fuel rods. | |||
Six fuels rods terminate at approximately one-third of the bundle length and are desi gnated as short part length fuel rods. Eight fuel rods are used as tie rods. The rods are spaced and supported by the upper and lower tie plates and eight spacers over the length of the fuel rods. For GNF2, the channel interacts with the Lower Tie Plate (LTP). | |||
The fuel rods consist of high-density ceramic UO2 or (U, Gd)O2 fuel pellets stacked within Zircaloy-2 cladding. The cladding will generally have an inner zirconium liner. The fuel rod is evacuated and backfilled with helium. | |||
Additional assembly and component description for the GE14 fuel design are provided in Reference 59. | |||
The ATRIUM-9B reload fuel assembly design (Figure 4.2-3c) is a 9 x 9 array with 72 enriched uranium fuel rods. The interior is an inert water channel. The ATRIUM-10 reload fuel assembly design (Figure 4.2-3d) is a 10X10 array with 83 full-length fuel rods, 8 part-length fuel rods, and one centrally located water channel. The lower tie plate has a nosepiece which has the function of supporting the fuel assembly in the reactor. The upper tie plate has a handle for transferring the fuel bundle from one location to another. The identifying assembly number is engraved on the top of the handle and a boss projects from one side of the handle to LSCS-UFSAR 4.2-19 REV. 20, APRIL 2014 aid in assuring proper fuel assembly orientation. Both upper and lower tie plates position the rod ends for operation and handling. The tie plates also support the weight of the fuel during operation and handling in the 8x8R, 8x8NB, and ATRIUM-9B fuel designs. For the ATRIUM-10 fuel design, the weight of the fuel is supported by the water channel. Finger springs are also employed with the LSCS design. The finger springs are located between the lower tie plate and the channel for the purpose of controlling the bypass flow through that flowpath (Figure 4.2-2, Flow Path 8). | |||
Additional details of the finger springs ar e provided in Section 9 of References 14 and 49. Zircaloy fuel rod spacers equipped with Inconel springs maintain rod-to-rod | |||
spacing. | |||
AREVA Fuel | |||
For the AREVA ATRIUM-9B fuel, eight of the fu eled rods are tie rods. Some of the rods contain gadolinia as a burnable absorber. Fuel rod pitch is maintained by seven spacers. The spacers are a weld ed zircaloy-4 structure with Inconel 718 springs. The centrally located water channel captures the spacers to maintain the proper axial spacing. | |||
The assembly contains one water cha nnel to improve uranium utilization and operational flexibility. It provides unv oided water to the inner portion of the assembly, thereby, providing additional mo deration. The relatively large amounts unvoided water in the interior of the assembly increases the hot-cold reactivity swing. This feature allows greater operat ional flexibility by allowing longer cycles while maintaining appropriate shutdown margin. | |||
For fuel rod removal, the upper tie plate must be depressed against the compression springs a short distance in order to allow the locking sleeves to be rotated 90. After rotating the locking sleeves, the upper tie pl ate is then free to be removed for fuel rod extraction or replacement. | |||
The lower tie plate consists of a machined stainless steel casting with a grid plate for lower end cap engagement and a lower nozzle to distribute coolant to the | |||
assembly. | |||
The upper tie plate is a cast and machined grid plate with attached bail handle to provide for fuel assembly handling and orie ntation. A unique serial identification number is engraved on the bail handle of each tie plate. This number can be read under water to allow identification of the assemblies in the core. | |||
The identification of fuel type and enrichment may be marked on the end of each fuel rod upper end cap. | |||
Additional assembly and component descriptions for the ATRIUM-9B fuel are provided in References 46 and 48. | |||
LSCS-UFSAR 4.2-19a REV. 20, APRIL 2014The ATRIUM-10 fuel assembly consists of many of the same design features as the ATRIUM-9B presented above. Specifically, the ATRIUM-10 consists of a lower tie plate with a debris filter (F UELGUARD), an upper tie plate, 91 fuel rods, 8 spacer grids, a central water channel (or box) and miscellaneous assembly hardware. Of the 91 fuel rods, 8 are part-length fuel ro ds. The structural members of the fuel assembly include the tie plates, spacer grids, water channel, and connecting hardware. The structural connection between the lower tie plate and upper tie plate is provided by the water channel. | |||
Seven spacers occupy the normal axial locations, while an eighth spacer is located a few inches above the lower tie plate. In a manner similar to an AREVA PWR desi gn, the lowermost spacer restrains the fuel rods just above the lower tie plate. | |||
Additional assembly and component descriptions for the ATRIUM-10 fuel are provided in References 55 and 56. | |||
4.2.2.4 Fuel Rod | |||
Each fuel rod consists of high density (> | |||
95% of theoretical density) UO 2 fuel pellets stacked in a Zircaloy cladding tube which is evacuated, backfilled with helium, and sealed by Zircaloy end plugs welded in each end. Beginning with the fresh fuel in LaSalle Unit 1 Cycle 2, all fuel rods are zirconium-barrier fuel with the exception of the fuel rods in 48 ATRIUM-10 fuel bundles first loaded in LaSalle Unit 2 Cycle 10. | |||
The zirconium-barrier fuel has a zircaloy fuel cladding with a metallurgically bonded layer of zirconium on the inner surface. Adequate free volume is provided within each fuel rod in the form of pellet-to-cladding gap and a plenum region at the top of the fuel rod to accommodate thermal and irradiation expansion of the UO 2 and the internal pressures resulting from the helium fill gas, impurities, and gaseous fission products liberated over the design life of the fuel. A plenum spring, or retainer, is provided in the plenum space to prevent movement of the fuel column inside the fuel rod during fuel shipping and handling (Figure 4.1-3). For GE fuel bundles, a hydrogen getter is also prov ided in the plenum space as assurance against the inadvertent admission of moisture or hydrogenous impurities into a fuel rod. Additional information concerning the getter is provided in Section 8 of Reference 14 and in Reference 41. | |||
Prior to the introduction of ATRIUM-10 fuel design at LaSalle, three types of rods were used in GE and ATRIUM-9B fuel bundles: standard rods, tie rods, and nonfueled water rods (Figures 4.2-3 throug h 4.2-3f). The eight tie rods in each bundle have upper end plugs which extend through the upper tie plate casting. The eight tie rods are structural members of th e fuel assembly. They serve to connect the upper and lower tie plates. The tie rods contain fuel and have upper and lower end caps designed for connection to the tie plates. These rods are threaded into the lower tie plate and latch into the upper tie plate to hold the assembly together. The tie rods carry the assembly weight duri ng handling and provide the coil spring reaction support. These tie rods support the weight of the assembly only during fuel handling operations when the assembly hangs by the handle; during operation, LSCS-UFSAR 4.2-20 REV. 20, APRIL 2014the fuel rods are supported by the lower tie plate. Fifty-four rods in the 8x8R and BP8x8R bundles are standard fuel rods. The GE8x8EB bundle has fifty-four standard fuel rods, eight tie rods, and two water rods. The GE8X8NB fuel design contains fifty-two standard fuel rods, eight tie rods, and one centrally located water rod. | |||
The GE14 fuel design, inserted in LaSalle after the ATRIUM-10 design, contains 70 full length standard rods, 14 part length fu el rods, 8 tie rods and 2 large non fueled water rods. | |||
The GNF2 fuel design (Figure 4.2-3f), in serted starting in LaSalle-1 Cycle 15, contains 70 full length standard rods, 8 long part length fuel rods, 6 short part length rods, 8 tie rods and 2 large non fueled water rods. | |||
The ATRIUM-9B has 64 standard fuel rods, 8 tie rods, and one centrally located square water channel. The end plugs of the standard rods have shanks which fit into bosses in the tie plates. An expansion spring is located over the upper end plug shank of each rod in the assembly to keep the rods seated in the lower tie plate while allowing independent axial expansion by sliding within the holes of the upper tie plate. For AREVA 9X9 fuel, all fuel rods except for the tie rods have coil compression springs located between the top of the fuel rods and the bottom surface of the upper tie plate. These compression springs provide a force to aid in seating the fuel rods in the lower tie plate and react against the upper tie plate. The springs accommodate variations in rod lengths arising from manufacturing tolerances and permit axially non-uniform thermal and irradiation induced growth of the fuel rods (Reference 49). | |||
Two rods in each 8x8R and BP8x8R fuel bundle are hollow water tubes, one of which (the spacer-positioning water rod) positions seven Zircaloy fuel rod spacers axially in the bundle. The GE8x8EB fuel bundle may have more water rods, and the GE8X8NB has one large centrally located water rod. The spacer rods are hollow Zircaloy tubes. The spacer-positioning water rod is equipped with the square bottom end plug. The spacer-positioning water rod is assembled to the spacers by sliding the rod through the spacer cells with the welded tabs oriented in the direction of the corner of the spacer cell. The rod is then rotated so that the tabs fit above and below the elements of the spac er structure, thereby positioning the spacer in the required axial position. The rod is prevented from rotating and unlocking the spacers by the engagement of its (square) lower end plug with the tie plate hole. Several holes are punched around the circumference of each of the water rods near each end to allow coolan t water to flow through the rod. | |||
LSCS-UFSAR 4.2-21 REV. 20, APRIL 2014In the GE14 and GNF2 fuel design, two rods in the bundle are hollow water tubes, one of which positions eight high performance Zr-2 fuel rod spacers axially in the bundle. These two water rods are hollow Zircaloy tubes that encompass eight fuel rod positions. The spacer positioning wa ter rod has tabs welded on it above and below each spacer position. This water rod acts as the spacer capture rod for the fuel assembly. The tabs pr event excessive movement of the fuel spacers in either the upward or downward directions. | |||
Several holes are punched around the circumference of each of the water rods near each end to allow coolant water to flow through the rod. | |||
For the ATRIUM-9B fuel, one essentially square water channel is located in the central region of the fuel assembly replacing nine fuel rods in a 3 x 3 array. The water-filled channel has inlet and outlet holes located at the lower and upper end caps. These holes are dimensioned to maintain unvoided water during steady-state operation inside the water channel. The end fittings are made of zircaloy-4. The channel is made from two "U" shaped strips cut and formed from the same sheet of zircaloy-4. | |||
The wall thickness is 0.0285 inches and prov ides adequate strength. The lower end cap of the water channel is threaded and it connects to the lower tie plate. The upper end cap penetrates the upper tie plate and provides a sliding joint to allow for differential growth. | |||
The water channel has zircaloy stops welded on the outside of the channel at axial locations corresponding to the spacer locations. There is a small gap between the stops and each spacer to allow differential thermal expansion between the channel and the fuel rods (Reference 49). | |||
The ATRIUM-10 fuel bundle design is similar in design to the ATRIUM-9B. The most significant difference is in the load-b earing member of the fuel bundle. The ATRIUM-10 does not utilize tie-rods. Instead the central water channel bears the load of the assembly. The attachment of the upper tie plate is accomplished using a simple locking mechanism. All moveable parts in the mechanism are captured such that no parts can come loose during tie plate removal or reactor operation. The reduced number of components results in pa rt from having a single upper tie plate locking mechanism. To keep the upper ti e plate in place, there is one, large compression spring on the water channel rather than the multitude of compression springs on individually fuel rods commonly associated with other designs. Also, no tie rod nuts or locking tabs are required as the water channel carries the weight of the fuel assembly during movement rather than tie rods as in most other BWR fuel designs. Additional component information for the ATRIUM-10 fuel design is provided in References 55 and 56. | |||
LSCS-UFSAR 4.2-21a REV. 20, APRIL 20144.2.2.5 Fuel Pellets The fuel pellets consist of high density ceramic uranium dioxide manufactured by compacting and sintering uranium dioxide powder into right cylindrical pellets. The GE pellets have flat ends and chamfered edges while ATRIUM-9B and ATRIUM-10 pellets are dished and have an outward land taper. Ceramic uranium dioxide is chemically inert to the cladding at operating temperatures and is resistant to attack by water. | |||
Several U-235 enrichments are used in the fuel assemblies. Fuel element design and manufacturing procedures have been developed to prevent errors in enrichment location within a fuel assembly. The LSCS fuel bundle incorporates the use of small amounts of gadolinium as a burnable poison in selected fuel rods. | |||
The GE 8x8R, GE 8x8NB, GE14, GNF2, ATRIUM-9B and ATRIUM-10 fuel design features are summarized in Tables 4.2-4 through 4.2-4(f). Characteristics of other fuel types used at LSCS are given in References 41, 44, 46, 49, 58, and 59. | |||
4.2.2.6 Fuel Channel Separate licensing topical reports (Ref erences 17, 41, 45, 58, and 59) provide complete descriptions and analytical results for channels supplied by General | |||
Electric Company and used in conjunction with the fuel described herein. The use of the GE14 fuel design at LaSalle introd uces the first use of a non-uniform wall thickness channel. The GE14 channel is an interactive channel with a thick corner-thin wall design (120 mil corners and 75 mil wall thickness). | |||
This channel is described in more detail in Reference 58. | |||
Reference 57 contains the specific design details for the fuel channels supplied by AREVA. The GNF2 channel is essentially the same as the GE14 channel with the ex ception that the GNF2 channel interacts directly with the lower tie plate and does not require finger springs (Reference 59). | |||
Beginning with LaSalle Unit 2 Cycle 13 and co ntinuing in subsequent cycles of Unit 2, AREVA Lead Fuel Assemblies (LFAs) ar e inserted into non-limiting core locations for demonstration purposes. The LFA program consists of eight Lead Fuel Assemblies of ATRIUM 10 fuel with advanced alloy channels. | |||
The advanced alloy channel is dimensionally similar to channels described in Reference 57. The variation from previous ly approved channels is the zircaloy comprising the channel is Zircaloy-BWR vers es Zircaloy-2 or Zircaloy-4 described in Reference 57. Advanced alloy channels are included in Mechanical Design Report for LaSalle Unit 1 and 2 ATRIUM 10 Fuel Assemblies , ANP-2741P, Revision 0, August 2008. However, the following functional description is included in this report for completeness. | |||
LSCS-UFSAR 4.2-22 REV. 13 The BWR Zircaloy fuel channel performs the following functions: | |||
(1) Forms the fuel bundle flow path outer periphery for bundle coolant flow. | |||
(2) Provides surfaces for control rod guidance in the reactor core. | |||
(3) Provides structural stiffness to the fuel bundle during lateral loadings applied from fuel rods through the fuel | |||
spacers. | |||
(4) Minimizes, in conjunction with the finger springs and bundle lower tieplate, coolant bypass flow at the channel/lower tieplate interface. | |||
(5) Transmits fuel assembly seismic loadings to the top guide and fuel support of the core internal structures. | |||
(6) Provides a heat sink during loss-of-coolant accident (LOCA). | |||
(7) Provides a stagnation envelope for in-core fuel sipping. | |||
The channel is open at the bottom and makes a sliding seal fit on the lower tieplate surface. The upper end of the fuel assemblies in a four-bundle cell are positioned in the corners of the cell against the top guide beams by the channel fastener springs. At the top of the channel, two diagonally opposite corners have welded tabs, one of which supports the weight of the channel from a threaded raised post and the upper tieplate. One of these raised posts has a threaded hole. The channel is attached using the threaded channel fastener a ssembly, which also includes the fuel assembly positioning spring. Channel-to-c hannel spacing is provided for by means of spacer buttons located on the upper port ion of the channel adjacent to the control rod passage area. | |||
In the mid 1970s, channel box wear and cracking was observed, first in a foreign plant and later in a few domestic boiling water reactors. The wear was located adjacent to incore neutron monitor and startup source locations. It was postulated and later confirmed by out-of-reactor testing, that the wear was caused by vibration of the incore tubes due primarily to a high-velocity jet of water flowing through the bypass flow holes in the lower core plate. To eliminate significant vibration of instrument and source tubes and the resultant wear on channel loop corners, LaSalle incorporated modifications similar to those described in Reference 36. | |||
These modifications involve the elimination of the bypass holes in the lower core plate and addition of two holes in the lower tie plate of each assembly to provide an alternate flow path. This design modification has been determined to have LSCS-UFSAR 4.2-23 REV. 14, APRIL 2002 negligible adverse effects on the mechanical, thermal, and nuclear performance of the channel boxes. Channel box wear has been observed to have been significantly reduced in operating boiling water reactors following the design modification. | |||
Proper orientation of fuel assemblies in the reactor core is readily verified by visual observation and is assured by verification procedures during core loading. Five separate visual indications of proper fuel assembly orientation exist: | |||
: a. The channel fastener assemblies, including the spring and guard used to maintain clearances between channels, are located at one corner of each fuel assembly adjacent to the center of the control rod. | |||
: b. The identification boss on the fuel assembly handle points toward the adjacent control rod. | |||
: c. The channel spacing buttons are adjacent to the control rod passage area. | |||
: d. The assembly identification numbers which are located on the fuel assembly handles are all readable from the direction of the center of the cell. | |||
: e. There is cell-to-cell symmetry. | |||
Experience has demonstrated that these desi gn features are clearly visible so that any misoriented fuel assembly would be re adily distinguished during core loading verification. | |||
Appropriate description and design drawings of reactivity control assemblies are included in Subsection 4.6.1.1.2. | |||
4.2.2.7 Reactivity Control Assembly and Burnable Poison Rods 4.2.2.7.1 Control Rods The control rods perform the dual function of power shaping and reactivity control. Four types of control rods are used at LSCS. Three designs are supplied by General Electric, and the fourth type supplied by ASEA-ATOM (ABB). Power distribution in the core is controlled during operation of the reactor by manipulating selected patterns of control rods. Control rod displacement tends to counterbalance steam void effects at the top of the core and re sults in significant axial power flattening. | |||
4.2.2.7.1.1 General Electric Control Rods Figures 4.1-4(a,b,c) show drawings of the General Electric Control Rods. | |||
LSCS-UFSAR 4.2-24 REV. 18, APRIL 2010 The General Electric original equipme nt and Duralife 215 control rod designs consist of a sheathed cruciform array of stainless steel tubes filled with boron-carbide powder. The control rods are 9.74 inches in total span and are separated uniformly throughout the core on a 12-inch pitch. Each control rod is surrounded by four fuel assemblies. | |||
The main structural member of Original Equipment and Duralife 215 control rod designs is made of Type 304 stainless steel and consists of a top handle, a bottom casting with a velocity limiter and contro l rod drive coupling, a vertical cruciform center post, and four U-shaped absorber tube sheaths. The top handle, bottom casting, and center post are welded into a single skeletal structure. The U-shaped sheaths are resistance-welded to the center post, handle, and castings to form a | |||
rigid housing to contain the boro n-carbide-filled absorber rods. | |||
Rollers at the top and bottom of the control rod guide the control rod as it is inserted and withdrawn from the core. The control ro ds are cooled by the core bypass flow. The U-shaped sheaths are perforated to al low the coolant to circulate freely about the absorber tubes. Operating experience has shown that control rods constructed as described above are not susceptible to dimensional distortions. | |||
The boron-carbide (B 4 C) powder in the absorber tubes is compacted to about 70% of its theoretical density. The boron-carbid e contains a minimum of 76.5% by weight natural boron. The boron-10 minimum content of the boron is 18% by weight. | |||
Absorber tubes are made of Type 304 (or 304 rad resist) stai nless steel. Each absorber tube is 0.188 inch in outside diam eter and has a 0.025-inch wall thickness. | |||
Absorber tubes are sealed by a plug welded into each end. The boron-carbide is longitudinally separated into individual compartments by stainless steel balls at approximately 16-inch intervals. The steel balls are held in place by a slight crimp of the tube. Should boron-carbide tend to compact in service, the steel balls distribute the resulting voids over the length of the adsorber tube. | |||
The Marathon design consists of square outer tubes with round inner diameters welded together and filled with B 4C capsules and hafnium rods. The Marathon design utilizes a 316 stainless steel handle, tie rod, transition piece, fins and locking plug. As of 1999, velocity limiter utilized on General Electric designs (fabricast) is made of CF3 casting. The absorber tubes are made of Rad Resist 304S stainless steel and welded together for rigidity. Some Marathon control blade handles have rollers or buttons to provide guidance for control rod insertion and withdrawal. Some Marathon control blade handles have no rollers or pads. | |||
4.2.2.7.1.2 ASEA-ATOM (ABB) Control Rods The second type of Control Rod utilized at LSCS is the ASEA-ATOM (ABB) CR82B. The ASEA-ATOM control rod functions the same as the General Electric control LSCS-UFSAR 4.2-24a REV. 18, APRIL 2010 rod, however the design of the ASEA-ATOM control rod is slightly different. Each of the four ASEA-ATOM control blade wi ngs has 520 horizontal holes (0.20 inch diameter) drilled directly into the blade wing (thus eliminating the perforated U-shaped absorber tube sheaths used in th e General Electric Control Rod design). | |||
The first 6 inches of the blade (beneath the top handle) consist of 22 holes containing hafnium rodlets. The remaining 498 holes contain boron-carbide powder compacted to above 70% of its theoretical density. The boron-carbide contains between 76.5-81% by weight natural boro | |||
: n. The boron-10 content in the ASEA-ATOM control rods is 19.9 +/- 0.3 atom %. The horizontal holes are covered with a stainless steel bar at the outer edge of the blade wing and are connected through a narrow slit. This allows gas pressure equalization between holes and prevents significant displacement of LSCS-UFSAR 4.2-25 REV. 17, APRIL 2008 the B 4 C powder. | |||
4.2.2.7.2 Velocity Limiter | |||
The control rod velocity limiter (Figures 4.2-5 and 4.2-5a) is an integral part of the bottom assembly of each control rod. This engineered safeguard protects against a high reactivity insertion rate by limiting the control rod velocity in the event of a control-rod-drop. It is a one-way device in that the control rod scram velocity is not significantly affected but the control rod dropout velocity is reduced to a permissible limit. | |||
The velocity limiter is in the form of two nearly mated conical elements that act as a large clearance piston inside the control rod guide tube. The lower conical element is separated from the upper conical element by four radial spacers 90 degrees apart and is at a 15-degree angle relative to the upper conical element, with the peripheral separation less than the central separation. | |||
The hydraulic drag forces on a control rod are proportional to approximately the square of the rod velocity and are negligible at normal rod withdrawal or rod insertion speeds. However, during the scram stroke, the rod reaches high velocity and the drag forces must be ove rcome by the drive mechanism. | |||
To limit control rod velocity during dropout but not during scram, the velocity limiter is provided with a streamlined profile in the scram (upward) direction. Thus, when the control rod is scrammed, water flows over the smooth surface of the upper conical element into the annulus between the guide tube and the limiter. In the dropout direction, however, water is trapped by the lower conical element and discharged through the annulus between th e two conical sections. Because this water is jetted in a partially reversed direction into water flowing upward in the annulus, a severe turbulence is created, th ereby slowing the descent of the control rod assembly to less than 3.11 ft/sec for current control blade designs. | |||
4.2.2.7.3 Burnable Poison Rods To meet the reactivity control requirements of any core load with excess reactivity, gadolinia-urania fuel rods are placed in each fuel assembly except for the natural uranium assemblies used in the initial cycle for both units and 48 low enriched ATRIUM-10 bundles first loaded in LaSalle Unit 2 Cycle 10. Some assemblies contain more gadolinia than others to improve transverse power flattening. Also, some assemblies contain axially distributed gadolinium to improve axial power flattening. GD 2 O 3 is uniformly distributed in the UO 2 pellet and forms a solid solution. | |||
LSCS-UFSAR 4.2-26 REV. 20, APRIL 2014 4.2.3 Design Limits and Evaluation A discussion of the fuel thermal-mechanical design limits and evaluation results for the BP8x8R, GE8x8EB, GE8X8NB, GE14, and GNF2 fuel designs is given in Section 2 of Reference 41. A similar discussion of the limits and results for the 8x8R fuel design is given in Appendix C of this reference. A similar discussion of the thermal mechanical design limits and evaluation results for the AREVA fuel can be found in Reference 46 through 49, 55, and 56. The information contained in the following Subsections is provid ed as a historical reference. | |||
4.2.3.1 Fuel Damage Analysis | |||
Fuel damage is defined as a perforation of the fuel rod cladding which would permit the release of fission produc ts to the reactor coolant. | |||
The mechanisms which could cause fuel da mage in reactor operational transients are: (a) severe overheating of the fuel ro d cladding caused by inadequate cooling, and (b) rupture of the fuel rod cladding due to strain caused by relative expansion of the UO 2 pellet. Cladding failure due to ov erpressure from vaporization of UO 2 following a rapid reactivity transient is not considered to be an operational transient. | |||
A value of 1% plastic strain of the Zircaloy cladding has traditionally been defined as the limit below which fuel damage due to overstraining of the fuel cladding is not expected to occur. The 1% plastic strain value is based on General Electric data on the strain capability of irradiated Zircaloy cladding segments from fuel rods operated in several BWR's (Reference 4). None of the data obtained falls below the 1% plastic strain value. However, a statistical distribution fit to the available data indicates the 1% plastic strain value to be approximately the 95% point in the total population. This distribution implies, th erefore, a small (< 5%) probability that some cladding segments may have plastic elongation less than 1% at failure. | |||
For fresh UO 2 fuel the calculated linear heat generation rate (LHGR) corresponding to 1% diametral plastic strain of the cla dding is approximately 25 kW/ft. Later in life, the calculated LHGR corresponding to 1% diametral plastic strain decreases to | |||
approximately 24 kW/ft at 20,000 MWd/tU and approximately 22 kW/ft at 40,000 MWd/tU. However, due to a depletion of fissionable material, the high-exposure fuel has less nuclear capability and will operate at correspondingly lower powers, so that a wide margin is maintained throughout life between the operating LHGR and | |||
the LHGR calculated to cause 1% cladding diametral strain. | |||
The addition of small amounts of gadolinia to UO 2 results in a reduction in the fuel thermal conductivity and melting temperature. The result is a reduction in the LHGR's calculated to cause 1% plastic diametral strain for gadolinia-urania fuel rods. However, to compensate for this th e gadolinia-urania fuel rods are designed | |||
to provide margins si milar to standard UO 2 rods. | |||
LSCS-UFSAR 4.2-27 REV. 20, APRIL 2014 4.2.3.2 Fuel Damage Experience The early GE BWR fuel experience has been extensively described in previous reports. In general, the Zircaloy cladding performance in the very early plants was good; however, some fuel failure mechanis ms were encountered and corrected. They are not significantly affecting current fuel performance. Details of this experience are provided in References 4, 19, 20 and 40. Later BWR fuel experience is given in Reference 41. | |||
One of the early causes of fuel failures was internal hydriding of the Zircaloy cladding due to internal attack by hydrogen. The source of hydrogen was primarily small amounts of moisture introduced into the fuel rod. A detailed analysis of the potential sources of hydrogen or moisture shows that the only source large enough to explain primary hydride failure was the UO 2 pellet itself. Major process steps such as increased fuel rod drying temperatures and dry grinding of pellets were incorporated in the manufacture of UO 2 pellets to ensure that no significant moisture could be present in the as-fabricat ed fuel rod. In addition, the fuel rod design was changed to incorporate a hydrogen gettering system to further assure that neither moisture nor any sporadic hydrogen is ever available to cause hydride failure of the cladding. Newer fuel designs, such as GNF2 fuel rods, are manufactured to a tighter hydrogen control limit as described in Reference 41; therefore, these newer fuel designs do not include the hydrogen getter. | |||
Another fuel failure mechanism encountered in operating BWR fuel is crud induced localized corrosion (CILC). CILC, however, has not been experienced at LaSalle. | |||
The one class of fuel failure mechanisms which has restricted operation on LaSalle Units 1 and 2 is known as "pellet-cladding interaction" (PCI). The failures are caused by the direct interaction between the irradiated urania fuel, including its inventory of fission products, and the zi rcaloy fuel sheath, or cladding. The incidence of such failures is closely linked to the power history of the fuel rod and to the severity and duration of power changes. Consequently, in order to reduce the probability of fuel failures due to the PCI phenomenon, operational constraints were placed on the reactors. | |||
These constraints were placed on local nodal power increases (ramp rates). Although these constraints have been very successful in reducing the incidence of fuel failures, they were costly in terms of operational flexibility. Consequently, there was strong incentive to provide a type of fuel resi stant to PCI. There have been a number of fuel design improvements that were made to minimize PCI failures. These improvements include: | |||
(a) the pellet geometry has been modified to include chamfered pellet ends and a shorter length in order to reduce the magnitude of inservice pellet distortions contributing to local cladding strains. | |||
For AREVA Fuel, the pellet geometry includes a land taper, dish and short length for enriched and gadolinia pellets. These features have been shown to reduce PCI. | |||
LSCS-UFSAR 4.2-28 REV. 18, APRIL 2010 (b) the cladding heat treatment temperature has been increased in order to reduce the statistical variability in cladding mechanical properties; (c) change from 7 x 7 to 8 x 8 to 9 x 9 to 10x10 lattice design to reduce fuel thermal duty; and (d) introduction of zirconium-barrier fuel. | |||
Improvements (a), (b) and (c) were made prio r to 1975. These, however, did not totally eliminate the PCI problem and it was necessary for plants to continue operation within the ramp rate guidelines. Extensive testing at Quad Cities Unit 2 showed that the introduction of zirconium-barrier fuel eliminated the need for use of the ramp rate guidelines on those fuel assemblies. | |||
The initial cycle fuel for LaSalle Units 1 and 2 did not incorporate the zirconium-barrier fuel. Consequently, operation was maintained within the PCIOMR guidelines for all fuel assemblies. However, reload fuel for subsequent cycles will be zirconium-barrier fuel with the exception of 48 low enriched ATRIUM-10 bundles first loaded in LaSalle Unit 2 Cycle 10. Operation of the zirconium-barrier fuel will be restrained only by the Technical Specifications. However, industry experience will continue to be utilized in order to implement appropriate administrative operating policies that may be more conservative than Technical Specifications. Operation of the non-barrier ATRIUM-10 fuel will be restrained by the guidelines provided by the fuel manufacturer (AREVA). | |||
Operation with failed fuel rods has demonstr ated that the fission product release rate from defective fuel rods can be controlled by regulating power level. The rate of increase in released activity apparently associated with progressive deterioration of failed rods has been deduced from chrono logical plots of the off-gas activity measurements in operating plants. These data indicate that the activity release level can be lowered by lowering the local power dens ity in the vicinity of the fuel rod failure. This measured data also indicates that cata strophic failure of the fuel assembly does not occur upon continued operation and that the presence of a failed rod in a fuel assembly does not result in propagation of failure to neighboring rods. Shutdown can be scheduled, as required, to repair or repl ace fuel assemblies that have large defects. | |||
Evaluation of the fission product release rate for failed fuel rods shows a wide variation in the activity release levels. Correlation of the release rates to defect type, size and specific power level indicates that fission product release rates are functions of power density and that progressive deterioration is a function of time. Available failure data are insufficient to quantify the detaile d correlation between these variables. | |||
4.2.3.3 Potential For a Water-Logging Rupture | |||
For water-logging to occur, the fuel cladding must have a small pinhole. Pinholes are eliminated during production by 100% | |||
leak check of assemblies. The leak LSCS-UFSAR 4.2-29 REV. 18, APRIL 2010 detector system consists of a high vac uum system capable of attaining pressures less than 5 x 10 | |||
-3 torr, and a mass spectrometer capable of detecting leaks smaller than the design limit (1 x 10 | |||
-8 std. cc/sec). The fuel bundle or fuel rod is placed in the vacuum chamber and evacuated to less than 1 x 10 | |||
-4 torr. After the vacuum is attained, the mass spectrometer tuned to the helium mass range is switched into the system. The output meter of the mass spectrometer will indicate the presence of any helium gas in the chamber. The design basis for the fuel precludes the potential for a water-logging rupture throughout the fuel cycle. | |||
4.2.3.4 Potential For Hydriding | |||
The design basis for fuel in regard to the cladding hydriding mechanism is to assure, through a combination of engineering specifications and strict manufacturing controls, that production fuel will not contain excessive quantities of moisture or hydrogenous impurities. Anal ysis of BWR fuel performance on fuel manufactured since July 1972 has indicated that this failure mechanism has been eliminated in BWR fuel through the adoption of the changes in the fuel rod design and manufacturing processes described in Subsection 4.2.3.2 and in Reference 41. | |||
AREVA addressed internal hydriding in Reference 49. The absorption of hydrogen by the cladding can result in cladding failure due to reduced ductility and formation of hydride platelets. Careful moisture control during fuel fabrication reduces the potential for hydrogen absorption on the inside of the cladding. The fabrication limit for total hydrogen in the fuel pellets is less than 2.0 ppm (References 46 and 49). | |||
4.2.3.5 Dimensional Stability The fuel assembly and fuel components are designed to assure dimensional stability in service. The fuel cladding and channel specifications include provisions to preclude dimensional changes due to residual stresses. In addition, the fuel assembly has been designed to accommoda te dimensional changes that occur in service due to thermal differential expansio n and irradiation effects. For example, the fuel rods are free to expand lengthwise independent of each other, and the channel is free to expand relative to the fuel bundle. | |||
The differential thermal expansion betwee n the tie plates and spacer grid is calculated to introduce a bending stress of less than 400 psi at the end of the fuel tube. Additional information regarding this calculation is presented in Section 4 of Reference 1. | |||
LSCS-UFSAR 4.2-29a REV. 18, APRIL 2010 During shipment the fuel bundle is in a horizontal position with flexible packing separators installed between the fuel rod so that the weight of the fuel rods is supported by the shipping co ntainer rather than the spac er grids. AREVA fuel rods are supported by the fuel assembly spacers during shipment. AREVA has performed testing to verify that this is acceptable for the Atrium-9B fuel assembly. | |||
Fuel bundle shipping procedures are qualified by a test performed on each new design, and each individual bundle is inspected relative to important dimensional characteristics following shipment to verify that no dimensional deviations have occurred. | |||
LSCS-UFSAR 4.2-30 REV. 18, APRIL 2010 The two major handling loads of concern are (1) the loads due to maximum upward acceleration of the fuel assembly while grappled, and (2) the loads due to impact of the fuel assembly into the fuel support wh ile grappled. Analyses of these loading conditions have been performed and the resulting fuel assembly component stresses are within design limits. Additional in formation on fuel handling and shipping loads for GE fuel is presented in Sectio n 5 of Reference 1 and in Reference 41. | |||
AREVA addresses fuel assembly handling loads in Reference 49. The AREVA assembly design must withstand all norm al axial loads from shipping and fuel handling operations without permanent de formation. AREVA uses either a stress analysis or testing to demonstrate compliance. The analysis or test uses an axial load of 2.5 times the static fuel assembly weight. At this load, the fuel assembly structural components must not show any yielding. Because of the design, failure from axial loads will occur at the tie rod end caps rather than in the cladding or tie plates. The fuel rod plenum has a design criteria associated with handling requirements. The spring must maintain a force against the stack weight to prevent column movement during handling (Reference 49). | |||
4.2.3.6 Fuel Densification The amount of incore fuel densification in BWR Zircaloy clad UO 2 pellet fuel has been observed to be small and is not cons idered to have any significant effects on fuel performance. Detailed consideration of the occurrence and potential effects of incore fuel densification in General Electr ic BWR's is reported in Reference 5 and its supplements. See Section 4.2.1.2.3. | |||
2 for a similar discussion for AREVA Fuel. | |||
4.2.3.7 Fuel Cladding Temperatures | |||
Fuel cladding temperatures for 8x8R type fuel are shown in Figure 4.2-6 as a function of surface heat flux for beginning of life conditions. A core distribution of segment powers is developed. The value of Zircaloy-2 thermal conductivity used in these calculations is approximately 9.0 Btu/hr-ft F. Calculated fuel cladding temperatures for 8x8R type fuel for late-in-life conditions are shown on Figure 4.2-7 as a function of heat flux. Th e solid lines on Figure 4.2-7 represent the expected fuel cladding temperatures. The temperatures employed in mechanical design evaluations are calculated using a conservative design allowance for the degradation in fuel rod surface heat transfer coefficient due to the accumulation of system corrosion products on the surface of the rod (crud) and cladding corrosion (zirconium oxide formation). The expected fuel cladding temperatures are calculated employing a more realistic allowance for the effects of crud and oxide on the fuel rod surface heat transfer coefficient. The calculated peak cladding temperatures are used in the th ermal and mechanical design analyses addressed in Reference 41. The fuel cladding temperatures for other fuel types can be found in Reference 41. | |||
LSCS-UFSAR 4.2-31 REV. 18, APRIL 2010 AREVA also prevents the fuel rod cladding from overheating by minimizing the probability of exceeding thermal margin limits on limiting fuel rods during normal operation and anticipated operatio nal occurrences (Reference 49). | |||
4.2.3.8 Peaking Factors The typical power distribution is divided into several components: the radial peaking factor, local peaking, and axial peaking. The maximum radial peaking factor is defined as the total power produced in the most limiting fuel assembly divided by the core average fuel assembly power. The maximum local peaking factor is defined as the maximum fuel rod heat flux in a fuel assembly divided by the fuel assembly average fuel rod heat flux. The maximum axial peaking factor is defined as the maximum heat flux along the length of a given fuel rod divided by | |||
the average heat flux of that rod. The initial reactor core de sign employs typical power peaking factors shown in Table 4.4-1. Peaking factors for reload cores are such that margins to limits for LHGR, MCPR, and APLHGR remain within the COLR limits. | |||
4.2.3.8.1 Local Peaking Factors The enrichment distribution in each fuel assembly is selected to reduce the relative local peak-to-average fuel rod power ratio within each assembly. The local peaking factor used for the initial design is provided in Table 4.4-1. | |||
4.2.3.8.2 Axial and Gross Peaking Factors The axial and gross peaking factors used for the initial core design are provided in Table 4.41. Axial and gross peaking factors for reload cores are such that margins to limits for LHGR, MCPR, and APLHGR remain within the COLR limits. | |||
4.2.3.9 Temperature Transients with Waterlogged Fuel Element As indicated in Subsection 4.2.3.3, the po tential for water-logging is considered in the fuel design. For waterlogging to occu r, the fuel cladding must have a small pinhole. Pinholes are eliminated during production by 100% leak check of assemblies. The leak detector system employ ed is described in Subsection 4.2.3.3. Since waterlogging is not expected and since it has not been observed in commercial power BWR fuel, no specific analysis of the consequences is performed. | |||
In the unlikely event that a waterlogged fu el element does exist in a BWR core, it should not have a significant potential for cladding burst (due to internal pressure) during a transient power increase unless the transient started from a cold or very low power condition. Normal reactor heatup rates are sufficiently slow ( 100 F/hr increase in coolant temperature) such that water vapor formed inside a waterlogged fuel rod would be expected to evacuate the rod through the same passage it entered, LSCS-UFSAR 4.2-32 REV. | |||
13 allowing internal and external pressures to equilibrate as the coolant temperature and pressure rise to the rated conditions. | |||
Once the internal and external pressure s are at equilibrium, at rated coolant pressure and temperature, transient power increases should, in general, have the effect of only slightly reducing the internal fuel rod plenum volume due to differential thermal expansion between fuel and cladding, thus effecting a small, short-term increase in internal fuel rod pressure. The potential short-term increase in pressure due to this effect would, in general, be small, (e.g., a power increase from the cold condition to peak rated power would increase internal pressure less than 15% in the peak power fuel rod fuel rod). For the range of anticipated transients, the cladding primary membrane stress resulting from the temporary increase in internal pressure above the coolant pressure would not be expected to exceed the cladding stress design limits of Subsection 4.2.1.2.5. | |||
4.2.3.10 Potential Damaging Temperature Effects During Transients | |||
There are no predicted significant temperature effects during a power transient resulting from a single operator error or single equipment malfunction which result in fuel rod, control rod, or structural damage. The calculated fuel rod cladding strain for this class of transients is sign ificantly below the calculated damage limit. The predicted additional bowing deflection for this class of transients is small compared to the steady-state rod-to-channel clearance. | |||
4.2.3.11 Energy Release During Fuel Element Burnout | |||
The metal-water chemical reaction between zirconium and water is given by: | |||
where H = 140 cal/g-mole. The reaction rate is conservatively given by the familiar Baker-Just rate equation: | |||
where W is milligrams of zirconium reacted per cm 2 of surface area, is time (seconds), R is the gas constant, (cal/mol | |||
- K), and T is the temp erature of zirconium | |||
( K). This rate equation has been shown to be conservatively high by a factor of 2 (Reference 21). The above equation can be differentiated to give the rate at which the thickness of the cladding is oxidized. This becomes: | |||
3-4.2 | |||
T A- exp X A th 2 1 where: th = rate at which the cladding thickness is oxidizing, 2-4.2 RT 45,500exp 10 x 33.3W 6 2 1)(4.2 H2HZrO02HZr22 2 LSCS-UFSAR 4.2-33 REV. | |||
13 = oxidized cladding thickness, A 1 , A 2 = appropriate constants, and T = reaction temperature. | |||
The reaction rate is inversely proportional to the oxide buildup; therefore, at a given cladding temperature the reaction rate is self-limiting as the oxide builds up. The total energy release from this chemical reaction over a time period is given by: | |||
where: N rods = number of rods experiencing boiling transition (at temperature T), -H = heat of reaction, C = cladding circumferences, L = axial length rod experiencing boiling transition, and | |||
= density of zirconium. | |||
This equation can be integrated and compa red to the normal bundle energy release if the following conservative assumptions are made: | |||
: a. At an axial plane all the rods experience boiling transition and are at the same temperature. This is highly conservative since, if boiling transition occurs, it will normally occur on the high power rod(s). | |||
: b. Boiling transition is assumed to occur uniformly around the circumference of a rod. This generally occurs only at one spot. | |||
: c. The rods are assumed to reach some temperature T instantaneously and stay at this temperature for an indefinite amount of time. | |||
This integration has been performed per ax ial foot of bundle and the total energy release as a function of time has been comp ared to the total energy release of a high power bundle (6 MW) over an equal amo unt of time. The results are shown in Figure 4.2-8. For example, if the temperatu re of all rods along a 1-foot section of the bundle were instantly increased to 1500 F, the total amount of energy that has 4-4.2 | |||
Xdt CL H- N Q t rods T LSCS-UFSAR 4.2-34 REV. 20, APRIL 2014 been released at 0.1 seconds is 0.4% of the total energy that has been released by the bundle (6 MW x 0.1 second). Note th at the fractional energy release decreases rapidly with time even though a constant temperature is maintained. This is because the reaction is self-limiting as was discussed above with the Baker-Just equation. | |||
The amount of energy released is dependent on the temperature transient, and the surface area that has experienced heatup. This, of course, is dependent on the initiating transient. For example, if boiling transition were to occur during steady-state operating conditions, the cladding surface temperature would range from 1000 F to 1500 F depending on the heat fluxes and heat transfer coefficient. Even assuming all rods experience boiling transi tion instantaneously, the magnitude of the energy release is insignificant. Significant boiling transition is not possible at normal operating conditions because of th e thermal margins at which the fuel is operated. This is also true for abnormal tr ansients. It can, therefore, be concluded that the energy release and potential for a chemical reaction is not an important consideration during normal op eration or abnormal transients. | |||
4.2.3.12 Energy Release for Rupture of Waterlogged Fuel Elements Experiments have been performed to show that waterlogged fuel elements can fail at a lower damage threshold than nonwaterlogged fuel during rapid reactivity excursion from the cold condition (Referen ces 22 and 23), (i.e., 60 cal/g as compared to > 300 cal/g). No analysis of cladding stress has been performed by GE for such conditions. One can postulate that if such a failure occurred, the resultant energy release and pressure pulse would be much less than for a nonwaterlogged fuel rod which exceeded its damage threshold since the energy level required for damage is apparently much lower in the waterlogged fuel element. Any fuel dispersion that might result in such a case would further reduce the severity of such a transient. | |||
4.2.3.13 Fuel Rod Behavior Effects from Coolant Flow Blockage In Reference 24, GE evaluated the conseque nces of a fuel bundle flow blockage incident. The percent of flow blocked to the bundle reduces the MCPR margin, and must be considered when evaluating the effects of a known lost part. A portion of reference 24 also discusses the consequences associated with 100% blockage of a fuel bundle; however, this event was never reviewed and approved by the NRC, nor has it ever been made a licensing requirement. | |||
Reference 16 provides an updated discussion, applicable to GE9, GE14, AREVA ATRIUM-9B fuel, and AREVA ATRIUM-10 fuel, of the effects of flow blockages on MCPR margin. Reference 60 provides the flow blockage MCPR analysis for GNF2 fuel. This relationship is used to determine the impact of known lost parts. This document also discusses the potential for fuel fretting for parts small enough to migrate into the bundle. Fuel fretting may lead to fuel failures, which would be detected by the offgas system. If a blocked bundle becomes suddenly unblocked, the increase in reactivity is less than the delayed neutron fraction, and therefore a prompt critical excursion is avoided. | |||
LSCS-UFSAR 4.2-35 REV. 18, APRIL 2010 4.2.3.14 Channel Evaluation | |||
An evaluation of fuel channel loading due to internally applied pressure has been performed. Tests have been conducted to verify the applicability of the "fixed-fixed | |||
beam" analytical model under uniform load. | |||
To confirm the applicability of the analytical model, a channel section was pressurized and the resultant deflections were measured and compared with the deflections predicted by the analytical model. A 4-foot-long section of channel with welded end plates was used for the test. The channel section was pressurized at room temperature in steps up to a pressure which was equivalent to a calculated stress intensity of approximately three times the yield strength of the channel material. Measurements of channel deflection were made for each pressure step and at zero pressure following each step. The deflection of the channel walls was found to be linear with pressure in th e pressure range tested. The measured deflection was within approximately 5% to 10% of the deflection predicted by the analytical model. There was no measurable permanent deformation of the channel walls until the calculated stress in the wall had reached approximately 1.2 times the measured yield strength of the test channel. | |||
The good performance of the channels have been demonstrated by both in-reactor experience and tests. The preponderance of the experience has been with channels that are 5.278 inches inside width with 0.080-inch wall thickness. Channel sizes ranging from 4.290 inches inside width with 0.060-inch walls to 6.543 inches inside width with 0.100-inch. walls, are included. | |||
The LSCS channel is 5.278 inches inside width with either 0.100-inch or 0.080-inch walls, dependin g on the specific reload. Additional information regarding channel analyses is presented in Section 2 of | |||
Reference 1 and in References 17, 45 and 57. | |||
Channel Management | |||
Channels are not being reused at LaSalle. This is one of the assumptions that is used for the MCPR safety lim it calculations by AREVA. | |||
To preclude unacceptable fuel element channel box deflection, a channel verification program, as discussed below, is implemented at LaSalle. | |||
The following general guidelines are followed to detect and control the potential of channel bowing. | |||
: a. Records are kept of channel location and exposure for each operating cycle. | |||
: b. Channels are not retained in the outer row of the core for more than two successive operating cycles. | |||
LSCS-UFSAR 4.2-36 REV. 13 c. At the beginning of each fuel cycle, the combined outer row residence time for any two channels in any control rod cell should not exceed four peripheral cycles. | |||
Prior to the beginning of a new operating cycle, control rod drive friction tests shall be performed for those core cells exc eeding the above general guidelines or containing fuel channels with exposures greater than 30,000 MWd/T (associated | |||
fuel bundle exposures). | |||
In lieu of friction testing, fuel channel me asurements may be used to justify use of fuel channels exceeding 30,000 MWd/T exposu re for a maximum of four additional operating cycles. | |||
In the future, analytical channel lifetime prediction methods, benchmarked and backed by periodic measurements of a sample of the highest duty fuel channels, may be used to assure clearance between control rod blades and fuel channels without additional testing. | |||
4.2.3.15 Fuel Reliability The information in this section is historical GE data on fuel reliability experience. The fuel component characteristics which ca n influence fuel reliability include: (a) the fuel pellet thermal and mechanical pr operties, dimensions, density, and U-235 enrichment; (b) the Zircaloy cladding thermal and mechanical properties, dimensions, and defects; (c) the fuel rod internal void volume and impurities; (d) the fuel rod-to-rod and rod-to-channel spacing; and (e) the spring constants of the fuel rod spacer springs which maintain contact between the spacer and the fuel rods. | |||
Important fuel pellet, cladding, and associated hardware characteristics and dimensions for the 8x8R fuel design are provided in Table 4.2-4 and Figure 4.1-2. The characteristics of other fuel designs ma y be found in Reference 41, 47, 48, or 49. | |||
The large volume of irradiation experience to date with GE BWR fuel indicates only a few mechanisms which have actually had a direct impact on fuel reliability; namely, cladding defects, excessive deposition of system corrosion products, cladding hydriding resulting from hydrogen impurity, and pellet-cladding interaction. | |||
The cladding defects have been virtually eliminated through implementation of improved quality inspection equipment and more stringent quality control requirements during fuel fabrication. Ex cessive deposition of corrosion products has also been virtually eliminated throug h improved control of corrosion product impurities in the reactor feedwater and by manufacturing improvements. Cladding hydriding is the result of excessive amounts of hydrogenous impurities (moisture and/or hydrogenous material) inadvertently introduced into the rod LSCS-UFSAR 4.2-37 REV. 18, APRIL 2010 during the fuel fabrication process. An alysis of BWR fuel performance on fuel manufactured since July 1972 has indicated that this failure mechanism has been eliminated in BWR fuel through the adoption of the changes in the fuel rod design and manufacturing processes described in Subsection 4.2.3.2 and Reference 41. | |||
Pellet-cladding interaction is the fuel failure mechanism which currently has the greatest effects on reactor operation at LaSalle. It has been identified as resulting from the combination of two basic effects: | |||
(a) the observed variability in local cladding strains due to pellet-cladding interaction which can result in the random occurrence of higher-than-average local strain value; and (b) the statistical | |||
variability in postirradiation ductility of the cladding which can result in the random occurrence of tubing segments with ductility lower than average. The fuel design improvements described in Subsecti on 4.2.3.2 have been shown to virtually eliminate PCI as a major cause of fuel failures. When zirconium-barrier fuel replaces all initial cycle fuel, the ramp rate guidelines may be virtually eliminated as a restraint on reactor operations. However, administrative restrictions may still be maintained. | |||
The cladding liner material is an enhanced zirconium alloy. The purpose of the material enhancement to the liner is to reduce the potential for secondary hydriding following the intrusion of coolant into a fuel rod. | |||
4.2.3.16 Fuel Operating and Developmental Experience | |||
Production fuel rods employing gadolinia-urania fuel pellets have been in use since 1965. Fuel operating experience is docu mented in References 4, 19, 40 and 42. | |||
4.2.3.17 Fuel Assembly | |||
During shipment the fuel bundle is in a horizontal position with flexible packing separators installed between the fuel rods so that the weight of the fuel rods is supported by the shipping co ntainer rather than the spac er grids. AREVA fuel rods are supported by the fuel assembly spacers during shipment. AREVA has performed testing to verify that this is acceptable for the Atrium-9B and ATRIUM-10 fuel assemblies. Fuel bundle shipping procedures are qualified by a test performed on each new design, and each individual bundle is inspected relative to important dimensional characteristics following shipment to verify that no dimensional deviations have occurred. | |||
The two major handling loads of concern are (1) the loads due to maximum upward acceleration of the fuel assembly while grappled, and (2) the loads due to impact of the fuel assembly into the fuel support wh ile grappled. Analyses of these loading conditions have been performed and the resulting fuel assembly component stresses are within design limits. Additional in formation of fuel handling and shipping loads is presented in Section 5 of Reference 1 and in Reference 41. | |||
LSCS-UFSAR 4.2-38 REV. 18, APRIL 2010 AREVA has also evaluated their fuel for fu el handling and shipping concerns. The assembly design must withstand all norm al axial loads from shipping and fuel handling operations without permanent de formation. AREVA uses either a stress analysis or testing to demonstrate compliance (Reference 46 and 55). | |||
The rod plenum spring also has design criteria associated with handling requirements. The spring must maintain a force against the stack weight to prevent column movement during handling. | |||
4.2.3.17.1 Loads Assessment of Fuel Assembly Components | |||
The analytical methods and acceptance criteria applied to determine the fuel assembly response to externally applied forc es are both deemed to be in accordance with the requirements of Appendix A to SRP 4.2. LaSalle County Station fuel assembly capability has been evaluated accordingly with acceptable results. Information on load assessment of fuel assembly components is provided in Table 3.9-4. | |||
4.2.3.18 Spacer Grid and Channel Boxes Refer to Subsection 4.2.3.14. | |||
4.2.3.19 Burnable Poison Rods | |||
The failure rate of the gadolinia-urania fuel rods is negligible, from previous operating experience over the years. | |||
4.2.3.20 Control Rods | |||
4.2.3.20.1 Materials Adequacy Throughout Design Lifetime | |||
The adequacy of the materials throughout the design life was evaluated in the mechanical design of the control rods. The primary materials, B 4 C powder, Hafnium, and Type 304 and Type 316L austen itic stainless steel, have been found suitable in meeting the demands of the BWR environment. | |||
4.2.3.20.2 Dimensional and Tolerance Analysis | |||
Layout studies are done to ensure that, given the worst combination of extreme detail part tolerances at assembly, no interference exists which will restrict the movement of control rods. In addition, preoperational verification is made on each control blade assembly to show that the acceptable levels of operational performance are met. | |||
LSCS-UFSAR 4.2-39 REV. 13 4.2.3.20.3 Thermal Analysis of the Tendency to Warp | |||
All parts of the control rod assembly remain at approximately the same temperature during reactor operation, ne gating the problem of distortion or warpage. Differential thermal growth is allowed for in the mechanical design. A minimum axial gap is maintained between absorber rod tubes and the control rod frame assembly for this purpose. In addition, dissimilar metals are avoided. | |||
4.2.3.20.4 Forces for Expulsion | |||
An analysis was made to evaluate the maximum pressure forces which could tend to eject a control rod from the core. The results of this analysis are given in | |||
Subsection 4.6.2.3.1.2.2 under item "Rupture of Hydraulic Line(s) to Drive Housing Flange." In summary, if the collet were to remain open, which is unlikely, calculations indicate that the steady-state control rod withdrawal velocity would be 2 ft/sec for a pressure-under line break, the limiting case for rod withdrawal. | |||
4.2.3.20.5 Functional Failu re of Critical Components The consequences of a functional failure of critical components have been evaluated and the results are covered in Subsection 4.6.2.3.2. | |||
4.2.3.20.6 Precluding Excessive Rates of Reactivity Addition | |||
In order to preclude excessive rates of reactivity addition, analysis has been performed both on the velocity limiter device and the effect of probable control rod failures (Subsection 4.6.2.3.2). | |||
4.2.3.20.7 Effect of Fuel Rod Failure on Control Rod Channel Clearances | |||
The control rod drive mechanical design ensures a sufficiently rapid and forceful insertion of control rods so that no channe l misalignments or distortion could hinder reactor shutdown by impeding a significant number of rods from full insertion. | |||
4.2.3.20.8 Mechanical Damage | |||
Analysis has been performed for all areas of the control system showing that system mechanical damage does not affect the cap ability to continuously provide reactivity control. | |||
The following discussion summarizes the analysis performed on the control rod guide tube. | |||
LSCS-UFSAR 4.2-40 REV. 13 The guide tube can be subjected to any or all of the following loads: | |||
: a. inward load due to pressure differential, | |||
: b. lateral loads due to flow across the guide tube, c. dead weight, and | |||
: d. seismic. | |||
In all cases analysis was performed considering both a recirculation line break and a steamline break, events which result in the largest hydraulic loadings on a control rod guide tube. | |||
Two primary modes of failure were considered in the guide tube analysis: exceeding allowable stress and excessive elastic deformation. It was found that the allowable stress limit will not be exceeded and that the elastic deformations of the guide tube never are great enough to cause the free movement of the control rod to be jeopardized. | |||
4.2.3.20.8.1 First Mode of Failure | |||
The first mode of failure is evaluated by the addition of all the stresses resulting from the maximum loads for the faulted condition. This results in the maximum theoretical stress value for that conditio | |||
: n. Making a linear supposition of all calculated stresses and comparing this value to the allowable limit defined by the ASME Boiler and Pressure Vessel Code yields a factor of safety of approximately 3. For faulted conditions the factor of safety is approximately 4.2. | |||
4.2.3.20.8.2 Second Mode of Failure | |||
Evaluation of the second mode of failure is based on clearance reduction between the guide tube and the control rod. The minimum allowable clearance is about 0.1 inch. This assumes maximum ovality and minimum diameter of the guide tube and the maximum control rod dimension. The analysis showed that if the approximate 6000 psi for the faulted condition were entire ly the result of differential pressure, the clearance between the control rod and th e guide tube would reduce by a value of approximately 0.01 inch. This gives a design margin of 10 between the theoretically calculated maximum displacement and the minimum allowable clearance. | |||
4.2.3.20.9 Analysis of Guide Tube Design Two types of instability were considered in the analysis of guide tube design. The first was the classic instability associated with vertically loaded columns. The second was the diametral collapse when a circular tube experiences external to internal differential pressure. | |||
LSCS-UFSAR 4.2-41 REV. | |||
13 The limiting axially applied load is ap proximately 77,500 pounds resulting in a material compressive stress of 17,450 psi (code allowable stress). Comparing the actual load to the yield stress level gives a design margin greater than 20 to 1. From these values it can be concluded that the guide tube is not an unstable column. | |||
When a circular tube experiences external to internal differential pressure, two modes of failure are possible depending on whether the tube is long or short. In the analysis here the guide tube is taken to be an infinitely long tube with the maximum allowable ovality and minimum wall thickness. The conditions will result in the lowest critical pressure calcul ation for the guide tube (i.e., if the tube were short, the critical pressure calculation would give a higher number). The critical pressure is approximately 140 psi. | |||
However, if the maximum allowable stress is reached at a pressure lower than the critical pressure, then that pressure is limiting. This is the case for a BWR guide tube. The allowable stress of 17,450 psi will be reached at approximately 93 psi. Comparing the maximum possible pressure differential for a steamline break to the limiting pressure of 93 psi gives a design margin greater than 3 to 1. Ther efore, the guide tube is not unstable with respect to differential pressure. | |||
4.2.3.20.10 Evaluation of Control Rod Velocity Limiter | |||
The control rod velocity limiter limits the free fall velocity of the control rod to a value that cannot result in nuclear system process barrier damage. This velocity is evaluated by the rod-drop accident analysis in Chapter 15.0. | |||
4.2.3.21 Rod Bowing | |||
4.2.3.21.1 GE Evaluation | |||
Irradiation-induced bowing in fuel rods and assemblies is a phenomenon which is not, in itself, a failure mechanism. However, rod bowing must be addressed in the design analysis so as to establish operational tolerances. General Electric has indicated that boiling water reactor fuel operating experience, testing, and analysis indicate that there is no significant problem with rod bowing even at small rod-to-rod and rod-to-channel clearances. Specifically, General Electric noted that: (1) no gross bowing has been observed (excluding the rod bowing-related failures in an early design); (2) a very low frequency of minor bowing has been observed; (3) mechanical analysis indicates deflections within design bases; and (4) thermal-hydraulic testing has shown that small rod-to-rod and rod-to-channel clearances pose no significant problem. Based on those report observations and Reference 37, that address: (1) updates the General Electric rod bowing experience; (2) verifies the accuracy with which General Electric measures rod bowing; and (3) documents the overall General Electric rod bowing sa fety analyses, there is no reason to anticipate a problem with fuel rod or asse mbly bowing during operation of LaSalle. | |||
LSCS-UFSAR 4.2-42 REV. 18, APRIL 2010 4.2.3.21.2 AREVA Evaluation | |||
Differential expansion between the fuel rods, and lateral thermal and flux gradients can lead to lateral creep bow of the rods in the spans between sp acer grids. This lateral creep bow alters the pitch between the rods and may affect the peaking and local heat transfer. The AREVA design basis for fuel rod bowing is that lateral displacement of the fuel rods shall not be of sufficient magnitude to impact thermal margins. Extensive post-irridation examin ations have confirmed that such rod bow has not reduced spacing between adjacent rods by more than 50%, The potential effect of this bow on thermal margins is negligible. Rod bow at extended burnup does not affect thermal margins due to the lower powers achieved at high exposure (Reference 49). | |||
4.2.3.22 Fission Gas Release | |||
The information in this section is historical GE data on fuel reliability experience. | |||
In 1976, the NRC had questioned the validity of fission gas release calculations in most fuel performance codes, including GEGAP - III (Reference 34), for a burnup greater than 20,000 megawatt days per ton of uranium. The General Electric Company was informed of this concern (Ref. 28) and was provided with a method of correcting fission gas release calculat ions for burnups greater than 20,000 megawatt days per ton of uranium (Ref. 29). Subsequently, the General Electric Company provided (Ref. 30) a generic reanal ysis of fuel performance calculations using GEGAP - III with the NRC's fission correction factor for BWR 2/3/4 plants with 7x7 and 8x8 fuel assemblies. Although the reanalysis was not specifically performed for the LaSalle fuel, a referenc ed 8x8 reanalysis performed for early refloodings plants bounded the LaSalle case. In the generic reanalysis, fuel rod internal pressure was shown to remain below system pressure for rod peak burnups below 40,000 megawatt days per ton of ur anium. This conclusion remains unchanged for the prepressurized fuel design (Ref. 31). The generic reanalysis did, however, result in higher initial stored energy and rupture pressure in the loss-of-coolant accident conditions, the higher fission gas release results in a maximum increase of 85 degrees Fahrenheit in calculated peak cladding temperature at end-of-life (approximately 33,000 megawatt days per tons of uranium planar average exposure). This added temperature increment results in calculated peak cladding temperatures of less than 2100 degrees Fahrenheit for average burnups below 33,000 megawatt days per ton of uranium and thus would not violate the 2200 degrees Fahrenheit loss-of coolant accident peak cladding temperature limit required by 10 CFR 50.46. | |||
A full reanalysis of the effects of fission gas release prior to exceeding a peak local burnup of 20,000 megawatt days per ton of uranium was required by the NRC for LaSalle. General Electric proposed that credit for approved emergency core cooling system evaluation model changes be used to offset any detrimental effects of fission gas release at high burnups (Ref. 32). The proposal was accepted by the NRC LSCS-UFSAR 4.2-43 REV. | |||
13 provided the more recent generic analysis was applicable to LaSalle. Per reference 33 CECo stated the latter generic analysis is applicable to LaSalle. The issue of enhanced fission gas release at high burnup is satisfactorily resolved for LaSalle. | |||
4.2.3.23 Ballooning and Rupture 4.2.3.23.1 GE Evaluation | |||
The information in this section is historical GE data on fuel reliability experience. In another loss-of-coolant accident related area of concern, the NRC had been generically evaluating three fuel material models that are used in emergency core cooling system evaluations. These models predict cladding rupture temperature, cladding burst strain, and fuel assembly flow blockage. | |||
In a letter from L. O. DelGeorge to A. Schwencer dated May 21, 1981, CECo endorsed the results of a generic sensitivity study performed by General Electric submitted to the NRC by letter dated May 15, 1981. As reported in this generic study, General Electric has assessed the boiling water reactor emergency core cooling system sensitivity to rupture temperature by using three rupture temperature models: | |||
(1) the General Electric CHAS TE model, (2) the NUREG-0630 model, and (3) a proposed General Electric model termed the adjusted model. For the LaSalle type of 8 x 8 with 2 water rod fuel design (designated the "improved 8 x 8 design"), General Electric found that the use of the NUREG-0630 model resulted in an increased peak cladding temperature of up to 50 degrees Fahrenheit over that which was obtained with the CHASTE model. However, sensitivity studies performed on the adjusted model, which is a combination of the CHASTE and | |||
NUREG-0630 models and may be the bett er of the three models, found the maximum impact on peak cladding temperature to be 10 degrees Fahrenheit. | |||
With regard to the boiling water reactor emergency core cooling system sensitivity to burst strain, the General Electric submittal assessed the impact of using a burst strain model that bounds the burst strain model given in NUREG-0630. | |||
It is estimated from the impact (i.e., < 5 degrees Fahrenheit) of the reduced versus the CHASTE model comparison that if the comparison had been made against the unaltered NUREG-0630 strain model, the impact would have been < 115 degrees Fahrenheit. In light of the calculated 2009 degrees Fahrenheit loss-of-coolant accident peak cladding temperature for La Salle, sufficient margin exists between the 2200 degrees Farenheit peak cladding te mperature limit as required by 10 CFR 50.46 and the calculated 2009 degrees Fahrenheit LaSalle peak cladding temperature to accommodate an uncertainty of 115 degrees Fahrenheit in the peak cladding temperature. | |||
LSCS-UFSAR 4.2-44 REV. 18, APRIL 2010 4.2.3.23.2 AREVA Evaluation During a severe loss of coolant accident, the cladding swelling and burst strain can result in flow blockage. Therefore, the LOCA analysis must consider the cladding swelling and burst strain impacts on the flow. AREVA uses the models in NUREG 0630 for cladding rupture. There is no explicit limit on the deformation. However, the calculations with the deformation models must satisfy the event criteria given in 10CFR 50.46. This swelling and rupture model is an integral part of the LOCA evaluation and is not part of the mechanical design analysis (Reference 49). | |||
====4.2.4 Testing==== | |||
and Inspection Plan Rigid quality control requirements are enforced at every stage of fuel manufacturing to ensure that the design specifications are met. Written manufacturing procedures and quality control plans define the steps in the manufacturing process. The quality control plan is provided in Reference 43. Each fuel tube is subjected to dimensional inspection and ultrasonic inspection to reveal defects in the cladding wall. Destructive tests are performed on representative samples from each lot of tubing, including chemical analysis, tensile, and burst tests. Integrity of end plug welds is assured by standardization of weld processes based on radiographic and metallographic inspection of welds. Completed fuel rods are helium leak tested to detect the escape of helium through the tubes and end plugs or welded regions. The UO 2 powder characteristics and pellet densities, composition, and surface finish are controlled by regular sampling inspections. | |||
The UO 2 weights are recorded at every stage in manufacturing. Dimensional measurements and visual inspections of critical areas, such as fuel rod-to-rod clearances, are performed after assembly. | |||
Each separate pellet enrichment group has at times been characterized by a single stamp. Such a co ntrol has varied over time and varied among fuel vendors. Fuel rods are individually numbered prior to fuel loading: (a) to aid in identifying which pellet type is to be loaded in each fuel rod; (b) to aid in identifying which position in the fuel assembly each fuel rod is to be loaded; and (c) to facilitate total fuel material accountability for a given project. | |||
Prior to introduction of AREVA fuel, further iden tification of individu al fuel rod gadolinia concentrations and uranium enrichments is accomplished by symbolization on the upper end plug shank for each differing rod. Each upper end plug is ensured proper placement on a fuel rod by reference to the specific fuel rod type. Each fuel rod is ensured of proper placement within a fuel bundle by inspection of the fuel rod serial number on the lower end plug or clad bar code. For AREVA fuel beginning with AREVA ATRIUM-10 fuel loaded into LaSalle 2 Cycle 10, fuel rod identi ty was tracked by use of a bar code on the cladding. This facilitates proper tracking at the fuel fabrication factory including proper loading into the fuel bundle skeleton through automated controls. Computer software ensures that the correct rods are loaded into the proper locations in the fuel bundle. | |||
Fuel rod inspection includes metallographic and radiographic (not applicable to upset shape welded fuel rods) examination of fuel rods on a sample basis. Sample tests are performed for qualification of weld stations, weld parameters, and weld operators prior to application. Production samples are tested as a check on th e process and process controls. | |||
LSCS-UFSAR 4.2-45 REV. 18, APRIL 2010 Fuel assembly inspections consist of complete dimensional checks of channels and fuel bundles prior to shipment. A sample of fuel bundles is given another visual and dimensional inspection of significant dimensions at the reactor site prior to use. Comparable tests and inspections are used by AREVA. | |||
Onsite receipt of fuel rods and other reactor internals is the responsibility of EGC. General Electric and AREVA do provide recommendations to the purchaser for receipt, inspection, and handling of these components. General Electric and AREVA also perform audits to ensure that these activities are performed in compliance with General Electric and AREVA requirements. Such audits, however, are performed solely to satisfy General Electric and AREVA interests relative to warranty fulfillment. | |||
The sampling rate and method of the site fu el receiving inspection are outlined in Table 4.2-5. However, current LaSalle fuel inspection procedures meet the requirements outlined in the fuel vendor's Quality Plan, which may or may not be the same as the sampling rate shown in Table 4.2-5. | |||
Verification of enrichment and burnable poison concentrations is described in Subsection 4.2.4.1. | |||
4.2.4.1 Testing and Inspection (Enrichm ent and Burnable Poison Concentrations) | |||
The shutdown reactivity requirement is veri fied during initial fuel loading and at any time that core loading is changed. Nuclear limitations for control rod drives and SLC are verified by periodically testing the individual system. | |||
The following serves to identify the various test and inspections employed by the Fuel Vendor(s) in verifying the nuclear characteristics of the fuel and reactivity control systems. Comparable tests and inspections are used by AREVA. | |||
4.2.4.1.1 Enrichment Control Program GE uses emission spectrometry for determining impurities and mass spectrometry for verifying the U-235 enrichment in samples of UO 2 powder. AREVA verifies that samples of incoming UF 6 and the resultant UO 2 powder are within limits for impurities by emission spectroscopy. Th e U-235 content of a st atistical sample of UF 6 is verified by gamma counting and by mass spectroscopy measurement. | |||
A sample of the sintered pellets is also checked for impurities by emission spectroscopy. AREVA performs chemical verification of impurities and O/U measurements on sintered pellets by emission spectroscopy, wet chemistry and LSCS-UFSAR 4.2-45a REV. 14, APRIL 2002 inert gas fusion. GE verifies the O/U ratio of UO 2 pellets and gadolinia bearing pellets up to 6 w/o Gd203 concentration by gravimetric methods. The O/U ratio for gadolinia bearing pellets with concentrat ion above 6 w/o Gd203 is confirmed using a spectrophotometric method. (G E uses emission spectrometry) | |||
Each rod is gamma scanned to screen out any possible but unlikely misplaced pellet or enrichment deviations. | |||
LSCS-UFSAR 4.2-46 REV. 18, APRIL 2010 4.2.4.1.2 Gadolinia Inspections The same rigid quality control requirements observed for standard UO 2 fuel are employed in manufacturing gadolinia-urania fuel. Gadolinia-bearing UO 2 fuel pellets of a given enrichment and gadolinia concentration are maintained in separate groups throughout the manufacturi ng process. For General Electric, the percent enrichment and gadolinia concentration characterizing a pellet group are identified by a stamp on the pellet. ForAREVA, gadolinia pellets are uniquely identified with a symbol stamped on the pellet. | |||
Fuel rods are individually numbered prior to loading of fuel pellets into the fuel rods: (1) to identify which pellet group is loaded in each fuel rod; (2) to identify which position in the fuel assembly each fuel rod is load ed; and (3) to facilitate total material accountability for a given project. The correct location of all fuel rods in the bundle is ensured through the use of a computer-controlled, automated bundle | |||
assembly machine. | |||
The following quality control inspections are made: | |||
: a. Gadolinia concentration in the gadolinia-urania powder blend is verified. | |||
: b. Sintered pellet UO 2-Gd 2 O 3 solid-solution homogeneity across a fuel pellet is verified by examination of ceramographic | |||
specimens. | |||
: c. Gadolinia-urania pellet identification is verified. | |||
: d. Gadolinia-urania fuel rod identification is checked. | |||
: e. Each gadolinia - urania fuel rod is scanned to assure proper assembly. | |||
: f. Gadolinia content is verified by X-ray fluorescence measurements of each pellet or scanning the assembled rod. | |||
All assemblies and rods of a given project are inspected to ensure overall accountability of fuel quantity and placement for the project. | |||
AREVA uses similar practices and techniques for gadolinia inspection. | |||
LSCS-UFSAR 4.2-47 REV. 13 4.2.4.1.3 Reactor Control Rods Inspections and tests are conducted at various points during the manufacture of control rod assemblies to ensure that design requirements are being met. All boron carbide lots are analyzed and certified by the supplier. Among the items tested are: | |||
: a. chemical composition, b. boron weight percent, c. boron isotopic content, and | |||
: d. particle size distribution. | |||
Following receipt of the boron carbide and review of material certificates, additional samples from each lot are tested including those previously listed. Control is maintained on the B 4C powder through the remaining steps prior to loading into the absorber rod tubes. | |||
Certified test results are obtained on other control rod components. The absorber rod tubing is subjected to extensive test ing by the tubing supplier, including 100% ultrasonic examination. Metallographic examinations are conducted on several tubes randomly selected from each lot to verify cleanliness and absence of conditions resulting from improper fabricat ion, cleaning or heat treatment. Other checks are made on the subassemblies and final control rod assembly, including weld joints inspected and B 4 C loading. | |||
4.2.4.2 Surveillance Inspection an d Testing of Irradiated Fuel Rods | |||
General Electric has a cooperative program of surveillance of BWR fuel, both production and developmental, which operates beyond current production fuel experience as it becomes available for inspection. The schedule of inspection is, of course, contingent on the availability of the fuel as influenced by plant operation. | |||
This program is provided in Reference 41. | |||
The lead experience fuel rods (with respect to exposure, linear heat generation rate, and the combination of both) are selectively inspected. Inspection techniques used include: a. leak detection tests, such as "sipping;" | |||
: b. visual inspection with various aids such as binoculars, borescope, periscope, and/or underwater TV with a photographic record of observations as appropriate; LSCS-UFSAR 4.2-48 REV. 13 c. nondestructive testing of selected fuel rods by ultrasonic and eddy current test techniques; and | |||
: d. dimensional measurements of selected fuel rods. | |||
Unexpected conditions or abnormalities which may arise, such as distortions, cladding perforation, or surface disturbance s are analyzed. Resolution of specific technical questions indicated by site examinations may require examination of selected fuel rods in the Radioactive Material Laboratory (RML) facilities. | |||
The results of the program are used to evaluate the boiling water reactor fuel design methods and criteria used by General Electric. | |||
The results of the surveillanc e program are generally review ed with the Division of Reactor Licensing and documented in generic fuel experience licensing topical | |||
reports. Historical fuel performance results prior to 1979 on highly precharacterized lead test assemblies are provided in several report s listed in Reference 38. The lead test assemblies are utilized as one means of providing some confirmation of design adequacy or early warning of negative features of the design. Details on lead test assembly programs are pr ovided in Reference 39. | |||
In addition to fuel bundle inspection, the fuel channels are under surveillance in continuing programs. These surveillance programs are designed not only for the evaluation of present day products, but are also providing data in the areas of alternate materials and design modeling. | |||
4.2.4.3 Operating Experience with Gadolinia-Containing Fuel | |||
Production fuel rods employing gadolinia-urania fuel pellets have been in use since 1965. During this time, a substantial num ber of gadolina-urania rods have been successfully irradiated to appreciable exposures. Additional information on gadolinia-urania physical and irradiation characteristics, material properties, and operating experiences is provided in Reference 25. | |||
Temperature coefficients are virtually unchanged because of gadolinia. The gadolinia-bearing pellets act as thermally gr ay or black adsorbers, and their effect on moderator coefficients in the lattice is not essentially different from that of the control which they replace. Doppler response is unaffected because the gadolinia has essentially no effect on the resonance group flux or on the U-238 content of the core. The concentration of gadolinia has been selected so that the initial concentration of the high cross section isotopes, Gd-155 and | |||
-157, will be completely depleted by the end of the first cycle. The irradiation pr oducts of this process are other gadolinia LSCS-UFSAR 4.2-49 REV. 14, APRIL 2002 isotopes having low cross sections. Power in the gadolinium pins generally remains below 90% of the average bundle power. The control augmentation effect disappears on a predetermined schedule without changes in the chemical composition of the fuel or the physical makeup of the core. | |||
The thermal margins described by the steady-state operating limits (LHGR, APLHGR and MCPR) are easily maintained in a gadolinia core because additional power shaping is possible through spatial va riation of the burnable poison loading. The damage limits on gadolinia-urania fuel rods are designed with similar margins as maintained for the UO 2 rods. 4.2.5 References | |||
: 1. NEDO-20948-P, "BWR/6 Fuel Design," December 1975. | |||
: 2. WAPD-TM-283, "Effects of High Burnup on Zircaloy-clad, Bulk UO 2 Plate Fuel Element Samples," September 1962. | |||
: 3. WAPD-TM-629, "Irradiation Behavior of Zircaloy-Clad Fuel Rods Containing Dished End UO 2 Pellets," July 1967. | |||
: 4. H. E. Williamson and D. C. Di tmore, "Experience with BWR Fuel Through September 1971," NEDO-10505, May 1972. | |||
: 5. D. C. Ditmore and R. B. Elkins, "Densification Considerations in BWR Fuel Design and Performance," NEDM-10735, December 1972. | |||
: 6. J. A. Christensen, "Melting Point of Irradiated Uranium Dioxide," WACP-6065, February 1965. | |||
: 7. "Thermal Conductivity of Uranium Dioxide," Technical Report series No. 59, IACA, Vienna, 1966. | |||
: 8. Supplement 1 to the Technical Re port on Densification of General Electric Reactor Fuels, December 1973. | |||
: 9. This reference has been deleted. | |||
: 10. J. P. Fritz, "Testing of Cruciform Control Rods for BWR/6" NEDO-10565, April 1972. | |||
: 11. R. J. Benche, "Visual and Photographic Examination of Dresden-1 High Exposure Control Rod B-87" NEDO-10541, April 1972. | |||
LSCS-UFSAR 4.2-50 REV. 19, APRIL 2012 12. W. F. O'Donnel and B. F. Langer, "Fatigue Design Basis for Zircaloy Components," Nuclear Science and Engineering, Vol. 20, pp. 1-12, 1964. 13. E. P. Quinn, "Vibration of Fu el Rods in Parallel Flow," GEAP-4059, July 1962. | |||
: 14. NEDO-20360-1P, Revision 4, "General Electric Boiling Water Reactor Generic Reload Application for 8x8 Fuel," March 1976. | |||
: 15. "Design Safety Standards for Boiling Water Reactors," complied by Safety and Standards Unit, NEDE-10370, June 1971. | |||
: 16. Design Analysis No. L-003508, Rev. | |||
0, "LaSalle Lost Parts Analysis," July 2010. | |||
: 17. "BWR Fuel Channel Mechanical Design and Deflection," NEDO-21354, September 1976. | |||
: 18. This reference has been deleted. | |||
: 19. R. B. Elkins, "Experience with BWR Fuel Through September 1974," NEDO-20922, June 1975. | |||
: 20. H. E. Williamson and D. C. Ditmore, "Current State of Knowledge High Performance BWR Zircaloy-Clad UO 2 Fuel," NEDO-10173, May 1970. | |||
: 21. "Thermal Response and Cladding Performance of an Internally Pressurized, Zircaloy Clad, Simulated BWR Fuel Bundle Cooled by | |||
Spray Under Loss-of-Coolant Conditions," GEAP-13112, April 1971. | |||
: 22. L. A. Stephan, "The Response of Waterlogged UO 2 Fuel Rods to Power Burst," IDO-ITR-105, April 1969. | |||
: 23. L. A. Stephan, "The Effects of Cladding Material and Heat Treatment on the Response of Waterlogged UO 2 Fuel Rods to Power Bursts," IN-ITR-111, January 1970. | |||
: 24. "Consequences of a Postulated Flow Blockage Incident in a Boiling Water Reactor," NEDO-10174, Re vision 1, October 1977. | |||
: 25. G. A. Potts, "Urania-Gadolinia Nuclear Fuel Physical and Irradiation Characteristics and Material Proper ties," NEDE-20943 (proprietary), | |||
NEDO-20943 (nonproprietary), January 1977. | |||
LSCS-UFSAR 4.2-51 REV. 13 26. "Creep Collapse Analysis of BWR Fuel Using Safe-Collapse Model," NEDE-20606 (Proprietary), NEDO-20606A (Non-Proprietary) August 1976. 27. This reference has been deleted. | |||
: 28. D. F. Ross (NRC) Letter to G.G. Sherwood (GE) dated November 23, 1976. | |||
: 29. NUREG-0418, "Fission Gas Release from Fuel at High Burnup," March, 1978. | |||
: 30. G.G. Sherwood (GE) Letter to D.F. Ross (NRC) dated December 22, 1976. | |||
: 31. General Electric report, NEDO - 23786-1, "Fuel Rod Pressurization - | |||
Amendment 1," May 1978. | |||
: 32. Letters from R. E. Engel to T.A. Ippolito dated May 6, 1981 and May 23, 1981. | |||
: 33. L. O Del George letter to A. Schwencer (NRC) dated September 21, 1981. | |||
: 34. General Electric topical report NEDO-20181, "GEGAP-III: A Model for the Predictions of Pellet - Cladding Thermal Conductance in BWR Fuel Rods," November 1973 | |||
: 35. Letter from W.R. Butler (NRC) To I. Stuart (GE), dated April 4, 1975. | |||
: 36. General Electric report, NEDO - 21156, "Supplemental Information for Plant Modification's to Eliminate Significant In-core Vibration," | |||
January, 1976. | |||
: 37. General Electric topical report, NEDE - 24284, "Fuel Rod Bowing in General Electric Boiling Water Reactors," Dated August 1980. | |||
: 38. NUREG 0633, "Fuel Performance Annual Report," December 1979. | |||
: 39. General Electric report, NEDC 24609, "Boiling Water Reactor Fuel Rod Performance Evaluation Program," February 1979. | |||
: 40. General Electric report NEDE , 21660-P, "Experience with BWR Fuel through December 1976," July 1977. | |||
LSCS-UFSAR 4.2-52 REV. 18, APRIL 2010 41. General Electric report NEDE-2 4011-P-A, "General Electric Standard Application For Reactor Fuel (GESTAR II)", (latest approved revision). | |||
: 42. Letter Number MFN-078-086/JSC-067-08 6, J. S. Channley (GE) to C. H. Berlinger (NRC), "1985 Fuel Experi ence Report," August 13, 1986. | |||
: 43. "Nuclear Energy Business Group BWR Quality Assurance Program Description," NEDO-11209-04A, March 1978. | |||
: 44. NEDE-31152, "GE Fuel Bundle Designs," (latest revision). | |||
: 45. General Electric Report GE-NE-523-A191-1294, "Final Report of the Impact of Using GE9 80-Mil Fuel Channel for the LaSalle Units 1 and 2", latest revision and supplements. | |||
: 46. Advanced Nuclear Fuels Corporation Generic Mechanical Design for Advanced Nuclear Fuels 9x9-IX and 9x9-9X Reload Fuel, ANF-89-014(A), Revision 1 and Suppl ements 1-2, Advanced Nuclear Fuels Corporation, Richland, WA, October 1991. | |||
: 47. Generic Mechanical Design for Exxon Nuclear Jet Pump BWR Reload Fuel, XN-NF-85-67(A), Exxon Nuclear Fuels Corporation, September 1986. | |||
: 48. AREVA document, Cycle Specific Fuel Design Report. | |||
: 49. Generic Mechanical Design Criteria for BWR Fuel Designs, ANF-89-098(P)(A), Revision 1, Suppl ement 1, Advanced Nuclear Fuels Corporation, May 1995. | |||
: 50. RODEX2A (BWR) Fuel Rod Thermal-Mechanical Evaluation Model, XN-NF-85-74(P)(A), Exxon Nuclear Company, August 1986. | |||
: 51. RODEX2 Fuel Rod Thermal-Mechanical Response Evaluation Model, XN-NF-81-58(P)(A) Supplements 1 and 2 Revision 2, Exxon Nuclear Company, May 1986. | |||
: 52. General Electric Report NEDE-22290 Su pplement 2, Production of Advanced Longer Life Control Rod in BWR 2-6 Plants. | |||
: 53. General Electric Report NEDE-3 1578PA, Class III, GE Marathon Control Rod Assembly, October 1991. | |||
: 54. General Electric Report NLM-CR-4505, Revision C, Class III, Tubricast Velocity Limits for Dura life and Marathon Control Blades, September 1998. | |||
LSCS-UFSAR 4.2-52a REV. 20, APRIL 2014 | |||
: 55. Fuel Design Evaluation for Siemens Power Corporation ATRIUM-10 BWR Reload Fuel, EMF-95-52(P), Revision 2, December 1998. | |||
: 56. Mechanical Design Evaluation for ATRIUM-10 BWR Reload Fuel to 54 MWd/KgU Assembly Exposure, EMF-98-006(P), January 1998. | |||
: 57. Mechanical Design for BWR Fuel Channels, EMF-93-177(P)(A) Revision 1 and Supplement 1, August 2005. | |||
: 58. GE14 Compliance with Amendment 22 of NEDE-24011-P-A (GESTAR II), NEDC - 32868P, latest approved revision. | |||
: 59. Design Analysis L-003664, Revision 1, "GNF2 Advantage Generic Compliance with NEDE-24011-P-A (GESTAR II), GEH Report NEDC-33270P" 60. Design Analysis L-003697, Revision 0, "GNF S-0000-0142-0455, Analysis of Flow Blockage Consequences for LaSalle Units 1 and 2 GNF2 New Fuel Introduction" LSCS-UFSAR TABLE 4.2-1 TABLE 4.2-1 REV. 13 TYPICAL LIMITING LHGR'S FOR GADOLINIA-URANIA FUEL RODS (kW/ft) | |||
EXPOSURE (MWd/tU) | |||
INCIPIENT CENTER MELTING 1% PLASTIC STRAIN | |||
OF CLADDING EXPECTED OPERATING MAXIMUM (4 wt% Gd 2 0 3) 0 18.4 23.0 ~ 4 20,000 17.8 21.4 ~ 11 40,000 16.7 18.2 ~ 8 | |||
LSCS-UFSAR TABLE 4.2-2a. TABLE 4.2-2 REV. 13 General Electric STRESS INTENSITY LIMITS YIELD STRENGTH (S y) ULTIMATE TENSILE STRENGTH (S u) Primary membrane stress 2/3 1/2 Primary membrane plus bending stress intensity 1 1/2 to 3/4 Primary plug secondary stress | |||
intensity 2 1.0 to 1.5 | |||
LSCS-UFSAR TABLE 4.2-2b TABLE 4.2-2b REV. 15, APRIL 2004 FANP STRESS INTENSITY LIMITS* | |||
Stress Intensity Limits** | |||
YIELD STRENGTH (y) ULTIMATE TENSILE STRENGTH (u) Primary membrane stress 2/3 y 1/2 y Primary membrane plus | |||
bending stress intensity 1.0 y 1/2 u Primary plug secondary stress | |||
intensity 2.0 y 1.0 u | |||
* Characteristics of the stress categories are defines as follows: | |||
a) Primary stress is a stress developed by the imposed loading which is necessary to satisfy the laws of equilibrium between external and internal forces and moments. The basic characteristics of a primary stress is that it is not self-limiting. If a primary stress exceeds the yield strength of the material through the entire thickness, the pr evention of failure is entirely dependent on the strain-hardening properties of the material. | |||
b) Secondary stress is a stress developed by the self-constraint of a structure. It must satisfy an imposed strain pattern rather than being in equilibrium with an external load. The basic characteristic of a secondary stress is that it is self-limiting. Local yielding and minor distortions can satisfy the discontinuity conditions due to thermal expansions which cause the stress to occur. | |||
** The stress intensity is defined as twice the maximum shear stress and is equal to the largest algebraic difference between any two of the three principal stresses. | |||
LSCS-UFSAR TABLE 4.2-3 TABLE 4.2-3 REV. 0 - APRIL 1984 CONDITIONS OF DESIGN RESULTING FROM IN-REACTOR PROCESS CONDITIONS COMBINED WITH EARTHQUAKE LOADING | |||
CONDITIONS OF DESIGN REACTOR INITIAL CONDITIONS PERCENT OF SAFE SHUTDOWN EARTHQUAKE IMPOSED 0% 50% 100% Startup Testing Upset -- -- Normal Normal Upset Faulted Abnormal Upset -- -- | |||
LSCS-UFSAR TABLE 4.2-4(a) | |||
TABLE 4.2-4(a) REV. 15, APRIL 2004 DATA FOR THE 8x8R FUEL DESIGN Core (Full Core Data) Fuel cell spacing (control rod pitch), in. 12Number of fuel assemblies 764 Total number of fueled rods 47368Core power density (rated power), kW/l 50.0 Total core heat transfer area, ft 2 74872 Fuel Assembly Data Overall length, in. | |||
176Nominal active fuel length, in. | |||
150Fuel rod pitch, in. | |||
0.640 Space between fuel rods, in. | |||
0.157Fuel channel wall thickness, in. | |||
0.100 Fuel bundle heat transfer area, ft 2 98.0Channel width (inside), in. | |||
5.278 Fuel Rod Data Outside diameter, in. | |||
0.483 Cladding inside diameter, in. | |||
0.419Cladding thickness, in. | |||
0.032Fission gas plenum length, in. | |||
10.0Pellet immersion density, % T.D. | |||
95 Pellet outside diameter, in. | |||
0.410 Pellet length, in. | |||
0.410 Water Rod Data Outside diameter, in. | |||
0.591 Inside diameter, in. | |||
0.531 LSCS-UFSAR TABLE 4.2-4(b) | |||
TABLE 4.2-4(b) REV. 15, APRIL 2004 DATA FOR THE GE8x8NB (GE9B) FUEL DESIGN Core (Full Core Data) Fuel cell spacing (control rod pitch), in. | |||
Number of fuel assemblies Total number of fueled rods Core power density (rated power), kW/l Total core heat transfer area, ft 2 12 764 45840 50.0 approximately 71816 Fuel Assembly Data Nominal active fuel length, in. Fuel rod pitch, in. | |||
Space between fuel rods, in. Fuel channel wall thickness, in. | |||
Fuel bundle heat transfer area, ft 2 Channel width (inside), in. | |||
150 0.640 0.157 0.100* approximately 94 5.278 Fuel Rod Data Outside diameter, in. Cladding inside diameter, in. Cladding thickness, in. Pellet immersion density, % T.D. | |||
Pellet outside diameter, in. | |||
Pellet length, in. | |||
0.486 0.419 0.032 96.5 0.411 0.410 Water Rod Data Outside diameter, in. | |||
Inside diameter, in. | |||
1.340 1.260 | |||
* Either 100 or 80 mil channels are used, depending on the reload. | |||
LSCS-UFSAR TABLE 4.2-4(c) TABLE 4.2-4(c) REV. 15, APRIL 2004 DATA FOR THE FANP ATRIUM-9B FUEL DESIGN Core (Full Core Data) Fuel cell spacing (control rod pitch), in. 12Number of fuel assemblies 764Total number of fueled rods 55008Core power density (rated power), kW/l 50.0 | |||
Total core heat transfer area, ft 2 77426 Fuel Assembly Data Overall length, in. | |||
176Nominal active fuel length, in. | |||
149Fuel rod pitch, in. | |||
0.569 Space between fuel rods, in. | |||
0.136Fuel channel wall thickness, in. | |||
0.08*Fuel bundle heat transfer area, ft 2 101.343Channel width (inside), in. | |||
5.278 Fuel Rod Data Outside diameter, in. | |||
0.433 Cladding inside diameter, in. | |||
0.3807Cladding thickness, in. | |||
0.026Fission gas plenum length, in. | |||
10.578Pellet immersion density, % T.D. | |||
96 Pellet outside diameter, in. | |||
0.3737 Pellet length, in. | |||
Enriched, in. Natural, in. 0.393 0.545 Water Box Data Outside dimension, in. | |||
1.516 Water box wall thickness, 0.0285 | |||
* Either 100 or 80 mil channels ar e used, depending on the reload. | |||
LSCS-UFSAR TABLE 4.2-4(d) TABLE 4.2-4(d) REV. 17, APRIL 2008 DATA FOR THE AREVA ATRIUM-10 FUEL DESIGN Core (Full Core Data) Fuel cell spacing (control rod pitch), in. 12Number of fuel assemblies 764 Total number of fueled rods 69524Core power density (rated power), kW/l 52.0 Total core heat transfer area, ft 2 86315 Fuel Assembly Data Overall length, in. | |||
176.386Nominal active fuel length, in. | |||
* Full length fuel rods 149* Part length fuel rods 90Fuel rod pitch, in. | |||
0.510 Space between fuel rods, in. | |||
0.114Fuel channel wall thickness, in. | |||
0.100 Fuel bundle heat transfer area, ft 2 113.0Channel width (inside), in. | |||
5.278 Fuel Rod Data Outside diameter, in. | |||
0.3957 Cladding inside diameter, in. | |||
0.3480Cladding thickness, in. | |||
0.024Fission gas plenum length, in. | |||
* Full length fuel rod 11.52(TIG)/ 11.53 (USW) | |||
* Part length fuel rod 5.26(TIG)/ 5.42 (USW) Pellet immersion density, % T.D. (typical, pellet enrichment dependent) 96.26Pellet outside diameter, in. 0.3413 Pellet length, in. Enriched, in. Natural, in. | |||
0.413 0.551 Water Box Data Outside dimension, in. 1.378Water box wall thickness, 0.0285 LSCS-UFSAR TABLE 4.2-4(e) | |||
TABLE 4.2-4(e) REV. 16, APRIL 2006 DATA FOR THE GE14 FUEL DESIGN Core (Full Core Data) | |||
Fuel cell spacing (control rod pitch), in. | |||
12 Number of fuel assemblies 764 Total number of fueled rods 70288 92*764Core power density (rated power), kw/l 53.01 NEDE 31152P Rev8 Total core heat transfer area, ft 2 86332 113*764 Fuel Assembly Data Nominal active fuel length, in. | |||
* Full length fuel rods 150 GE Dwg 217C1442 | |||
* Part length fuel rods 84 GE Dwg 217C1444Fuel rod pitch, in. 0.510 NEDE 31152P Rev 8 Space between fuel rods, in. 0.106 NFM DIR-00-081, Nov 30, 2000, GE14 Design ReviewFuel channel wall thickness (corner/median), in. 0.120/0.075 NEDE 31152P Rev 8 Fuel bundle heat transfer area, ft 2 113 NEDE 31152P Rev 8Channel width (inside), in. 5.278 NEDE 31152P Rev 8 Fuel Rod Data Outside diameter, in. 0.404 NEDE 31152P Rev 8Cladding inside diameter, in. 0.352 NEDE 31152P Rev 8Cladding thickness, in. 0.026 NEDE 31152P Rev 8Fission gas plenum length, in. | |||
* Full length fuel rod 9.64 GE Dwg 217C1442 | |||
* Part length fuel rod 10.94 GE Dwg 217C1444Pellet immersion density, %T.D. (typical, pellet enrichment dependent) 97.0 NEDE 31152P Rev 8 Pellet outside diameter (cold), in. 0.345 GE Dwg 137C9061 Pellet length, in. 0.370 GE Dwg 137C9061 Water Rod Data Outside diameter, in. 0.980 NEDE 31152P Rev 8Inside diameter, in. 0.920 NEDE 31152P Rev 8 LSCS-UFSAR TABLE 4.2-4(f) | |||
TABLE 4.2-4(f) REV. 20, APRIL 2014 DATA FOR THE GE14 FUEL DESIGN Core (Full Core Data) | |||
Fuel cell spacing (control rod pitch), in. | |||
12 Number of fuel assemblies 764 Total number of fueled rods 70288 92*764 Core power density (rated power), kw/l 52.5 Total core heat transfer area, ft2 86332 113*764 Fuel Assembly Data Nominal active fuel length, in. | |||
* Full length fuel rods 150 GE Dwg 217C1442 | |||
* Part length fuel rods 84 GE Dwg 217C1444 Fuel rod pitch, in. | |||
0.510 NEDE 31152P Rev 8 Space between fuel rods, in. | |||
0.106 NFM DIR-00-081, Nov 30, 2000, GE14 Design Review Fuel channel wall thickness (corner/median), in. | |||
0.120/0.075 NEDE 31152P Rev 8 Fuel bundle heat transfer area, ft2 113 NEDE 31152P Rev 8 Channel width (inside), in. | |||
5.278 NEDE 31152P Rev 8 Fuel Rod Data Outside diameter, in. | |||
0.404 NEDE 31152P Rev 8 Cladding inside diameter, in. | |||
0.352 NEDE 31152P Rev 8 Cladding thickness, in. | |||
0.026 NEDE 31152P Rev 8 Fission gas plenum length, in. | |||
* Full length fuel rod 9.64 GE Dwg 217C1442 | |||
* Part length fuel rod 10.94 GE Dwg 217C1444 Pellet immersion density, %T.D. (typical, pellet enrichment dependent) 97.0 NEDE 31152P Rev 8 Pellet outside diameter (cold), in. | |||
0.345 GE Dwg 137C9061 Pellet length, in. | |||
0.370 GE Dwg 137C9061 Water Rod Data Outside diameter, in. | |||
0.980 NEDE 31152P Rev 8 Inside diameter, in. | |||
0.920 NEDE 31152P Rev 8 | |||
LSCS-UFSAR TABLE 4.2-5 (SHEET 1 OF2) TABLE 4.2-5 REV. 13 SITE FUEL RECEIVING INSPECTION | |||
*, ** FUEL INSPECTION OBJECTIVES CHARACTERISTIC INTENDED METHOD E XPECTED FREQUENC Y Container Damage and Leak Visual 100% Bundle Damage Visual 100% Shipping Separators Visual 100% | |||
Removed Cleanliness Visual 100% | |||
Rod Integrity Visual, gauge when required 100% Lock Tab Washers Visual 100% Channel Integrity Visual 100% | |||
Channel Cleanliness Visual 100% | |||
Guard Integrity and Installation Visual and Torque Wrench 100% Spacer Damage Visual 100% for first 5 | |||
bundles and every | |||
20th thereafter, otherwise the middle 3 spacers. Rod to Rod Feeler gauge 100% of first 5 bundles | |||
and every 20th thereafter, otherwise two sections, all spaces, alternate the sections. Rod-to-Simulated | |||
Channel Simulated Channel and Feeler Gauge 100% of first 5 bundles | |||
and every 20th thereafter, otherwise 2 sections, 4 sides per section, alternate sections excluding end sections. | |||
LSCS-UFSAR TABLE 4.2-5 (SHEET 2 OF 2) TABLE 4.2-5 REV. 15, APRIL 2004 CHARACTERISTIC INTENDED METHOD EXPECTED FREQUE NSpring Length Visual 100% for all bundles Gauge 100% for first 5 | |||
bundles and every fourth thereafter, otherwise visual inspection. | |||
Finger Spring Seated in Pocket Visual 100% for all bundles Gauge 100% for first 5 | |||
bundles and every | |||
fourth thereafter, otherwise visual inspection. | |||
NOTE: Deviations require 100% inspection of the next 5 bundles for that characteristic. Two deviations for a characteristic within 6 consecutive bundles require revision of the AQL (acceptable quality level) with the General El ectric, Wilmington, North Carolina, U.S.A., facility. | |||
Where a reduced inspection was performed, a ll inspection steps shall be designated S OK (stamped OK). | |||
* Current LaSalle fuel inspection procedures meet the requirements outlined in the fuel vendor's Quality plan, which may or may not be the same as the sampling rate in Table 4.2-5. | |||
** These inspection objectives are specific to GE fuel. FANP fuel has similar inspection objectives for the FANP designs. | |||
LSCS-UFSAR 4.4-1 REV. 14, APRIL 2002 | |||
===4.4 THERMAL=== | |||
AND HYDRAULIC DESIGN | |||
====4.4.1 Design==== | |||
Bases | |||
4.4.1.1 Safety Design Bases | |||
Thermal hydraulic design of the LaSalle County Station (LSCS) core is established | |||
and based upon the following design bases: | |||
: a. Actuation limits for the devices of the nuclear safety systems are employed such that no fuel damage occurs as a result of | |||
abnormal transients (Chapter 15.0). Specifically, the minimum | |||
critical power ratio (MCPR) operating limit is specified such that | |||
at least 99.9% of the fuel rods in the core will not experience | |||
boiling transition during the most severe abnormal operational | |||
transient. A 1% plastic strain limit is specified to ensure that | |||
clad overstraining does not occur. b. Thermal hydraulic safety limits are used in setting safety margins and the consequences of fuel barrier failure to public | |||
safety. c. The nuclear system must meet the requirements in 10CFR50, Appendix A, General Design Criterion 12 - Suppression of | |||
Reactor Power Oscillations. | |||
4.4.1.2 Power Generation Design Bases | |||
The thermal-hydraulic design of the core provides the following operational | |||
characteristics: | |||
: a. ability to achieve rated core power output throughout the design life of the fuel without sustaining premature fuel failure, and b. flexibility to adjust core output over the range of plant load and load maneuvering requirements in a stable, predictable manner | |||
without sustaining fuel damage. | |||
4.4.1.3 Requirements for Steady-State Conditions | |||
Steady-State Limits | |||
For purposes of maintaining adequate thermal-hydraulic margin during normal steady-state operation, the minimum critical power ratio must not be less than the | |||
required MCPR operating limit, the operational linear heat generation rate (LHGR) | |||
is maintained below the LHGR limit for the fuel type, and the maximum average LSCS-UFSAR 4.4-2 REV. 14, APRIL 2002 planar linear heat generation rate (M APLHGR) must be maintained below the limits for the plant. This does not specify the operating power nor does it specify peaking factors. These parameters are determined subject to a number of | |||
constraints, including the thermal limits given previously. The core and fuel | |||
thermal-hydraulic design basis for steady-state operation, has been defined to provide margin between the steady-state operating condition and any fuel damage | |||
condition to accommodate uncertainties and to ensure that no fuel damage results, even during the worst anticipated transient conditions at any time in life. For | |||
LSCS, the operating limits for all three fuel thermal design limits are contained in the Core Operating Limits Report. | |||
4.4.1.4 Requirements for Transient Conditions | |||
Transient Limits | |||
The transient thermal-hydraulic limits are established such that no fuel damage is | |||
expected to occur during the most severe abnormal operating transient. Fuel | |||
damage is defined as perforation of the cladding that permits release of fission | |||
products (Section 4.2). Mechanisms that cause fuel damage in reactor transients | |||
are: | |||
: a. severe overheating of fuel cladding caused by inadequate cooling, and b. fracture of the fuel cladding caused by relative expansion of the uranium dioxide pellet inside the fuel cladding. | |||
For design purposes, the transient thermal-hydraulic limit requirement is met if at | |||
least 99.9% of the fuel rods in the core do not experience boiling transition during | |||
any abnormal operating transient. No fuel damage is expected to occur even if a | |||
fuel rod actually experiences a boiling transition. | |||
A value of 1% plastic strain of Zircaloy cladding has been established as the limit | |||
below which fuel damage from overstrainin g the fuel cladding is not expected to occur. Available data indicate that the threshold for damage is in excess of this value. The linear heat generation rate required to cause this amount of cladding | |||
strain decreases with burnup. | |||
4.4.1.5 Summary of Design Bases | |||
In summary, the steady-state thermal-hydraulic operating limits have been | |||
established to ensure that the design basis is satisfied for the most severe abnormal | |||
operational situation, whether a transient or an accident. Transient analyses are | |||
performed that demonstrate compliance with overpower transient limits assuming | |||
steady-state operation has been in compliance with steady state operating limits. | |||
An overpower which occurs during an abnormal operational transient must not LSCS-UFSAR 4.4-3 REV. 14, APRIL 2002 result in violation of the MCPR safety limit for the plant. Demonstration that the transient limits are not exceeded is sufficient to conclude that the thermal hydraulic | |||
design basis is satisfied. | |||
The MCPR, LHGR and MAPLHGR limits are su fficiently general so that no other limits need to be stated. The cladding and fuel bundle integrity criterion is assured as | |||
long as MCPR, LHGR and MAPLHGR limits are met. There are no additional design criteria on coolant void fraction, core coolant flow-velocities, or flow distribution, nor are they needed. Core design and target rod patterns ensure CPRs remain above the | |||
MCPR limits, thereby ensuring bundle para meters (e.g., flow, power, void fraction) remain within prescribed ranges. The coolant flow velocities and void fraction become | |||
constraints upon the mechanical and physics design of reactor components and are | |||
partially constrained by stability and control requirements. | |||
4.4.1.5.1 Fuel Cladding Integrity The fuel cladding integrity is defined in Subsection 4.2.1. The fuel cladding integrity | |||
from a thermal hydraulic viewpoint is assured by the operating and transient MCPR | |||
requirements. | |||
4.4.1.5.2 Fuel Assembly Integrity The fuel channel provides adequate lateral structural support for the fuel bundle and protects the fuel rods and spacers from impact and abrasion. The upper tie-plate | |||
handle is capable of supporting the weight of the fuel assembly. Specific design characteristics are given in Section 4.2. | |||
4.4.1.5.3 Fuel-Cladding Gap Characteristics | |||
The subject of fuel to cladding gap characteristics is covered in Section 4.2. | |||
====4.4.2 Description==== | |||
of Thermal Hydraulic Design of Reactor Core 4.4.2.1 Summary Comparison An evaluation of plant performance from a thermal and hydraulic standpoint is | |||
provided in Subsection 4.4.4. | |||
Transient evaluations are given in Chapter 15. A tabulation of thermal and hydraulic | |||
parameters of the LSCS reactor initial core, along with a comparison to the initial core | |||
of other reactors of a similar design, are given in Table 4.4-1. | |||
4.4.2.2 Critical Power Ratio There are three different types of boiling heat transfer in water forced convection | |||
systems: nucleate boiling, transition boiling, and film boiling. Nucleate boiling, at | |||
lower heat transfer rate, is an extremely efficient mode of heat transfer, allowing LSCS-UFSAR 4.4-4 REV. 18, APRIL 2010 large quantities of heat to be transferred with a very small temperature rise at the heated wall. As heat transfer rate is increased the boiling heat transfer surface | |||
alternates between film and nucleate boiling, leading to fluctuations in heated wall | |||
temperatures. The point of departure from the nucleate boiling region into the | |||
transition boiling region is called the boiling transition. Transition boiling begins at | |||
the critical power, and is characterized by fluctuations in cladding surface | |||
temperature. Film boiling occurs at the highest heat transfer rates; it begins as | |||
transition boiling comes to an end. Film boiling heat transfer is characterized by | |||
stable wall temperatures which are higher than those experienced during nucleate | |||
boiling. | |||
4.4.2.2.1 Boiling Correlations | |||
4.4.2.2.1.1 GE Fuel | |||
The occurrence of boiling transition is a function of the local steam quality, boiling | |||
length, mass flow rate, pressure, flow geom etry, and local peaking pattern. General Electric has conducted extensive experimental investigations of these parameters. | |||
These parametric studies encompass the enti re design range of these variables. In the experimental investigations, a boiling transition event was associated with a | |||
25º F rise in rod surface temperature. The (critical) quality at which boiling | |||
transition occurs as a function of the distance from the equilibrium boiling | |||
boundary is predicted by the GEXL (G eneral E lectric Critical Quality X - Boiling Length) correlation. This correlation is based on accurate test data of full-prototype simulations of reactor fuel assemblies operating under conditions duplicating those in actual reactor designs. The GEXL correlation is a best fit to the data and is used | |||
together with a statistical analysis to assure adequate reactor thermal margins (References 1 and 11). | |||
The figure of merit used for reactor design and operation is the critical power ratio (CPR). This is defined as the ratio of the bundle power which would produce | |||
equilibrium quality equal to but not exceeding the correlation value (critical | |||
quality), to the bundle power at the reactor condition of interest (i.e., the ratio of | |||
critical bundle power to operating bundle power). In this definition, the critical | |||
power is determined at the same mass flux, inlet temperature, and pressure which | |||
exist at the specified reactor condition. | |||
The core is sized with sufficient coolant flow to assure that the MCPR is maintained | |||
greater than the operating limit at rated conditions. | |||
4.4.2.2.1.2 AREVA Fuel | |||
In the AREVA methodology, the fuel assembly critical power corresponding to a particular reactor operating state is determined from the SPCB (References 25 | |||
and 26) or ANF-B (Reference 17) critical power correlations. | |||
LSCS-UFSAR 4.4-5 REV. 18, APRIL 2010 The ANF-B and SPCB correlations provides a generic tool for evaluating critical power and to assess thermal margin for all current domestic AREVA BWR fuel designs. It is based on a data base characteristic of AREVA product designs. The database contains data for AREVA fuel designs with both axially uniform and nonuniform power profiles. | |||
The ANF-B and SPCB critical power correlations are an empirical representation of | |||
planar average thermal-hydraulic fluid conditions at which boiling transition has | |||
been experimentally determined. The minimum heat flux required to produce | |||
boiling transition is predicted from fluid conditions of pressure, mass velocity, and | |||
enthalpy averaged over the plane of interest. The correlation contains correction | |||
factors for the effects of boiling transition due to a nonuniform axial heat flux | |||
profile and the grouping of relatively high-powered rods. | |||
The test assemblies include full-length rods; typical BWR grid spacers; 4x4, 5x5, and 9x9 rod configurations; and a variety of rod diameters, assembly hydraulic | |||
diameters, rod-to-wall spacings, and rod-to-rod spacings. The database was | |||
compiled from data taken at two test l aboratories: Columbia University and the ATLAS facility. The uniform axial data was used to develop the correlation, while | |||
the nonuniform axial data was used to validat e the correlation with the Tong factor. | |||
Therefore, the correlation has been checked against independent test data. | |||
The correlations address the effects of op erating pressure, mass velocity, enthalpy, axial power profile, local power peaking and distribution, rod diameter, and fuel | |||
assembly hydraulic diameter and heated length on boiling transition. | |||
The ANF-B and SPCB correlations have also been used to predict the number of | |||
rods experiencing boiling transition (predi ct multiple indications) for the test database. The probability of boiling transition for each rod in a test section was | |||
determined from the critical power prediction based on that rod. The probabilities | |||
for all the rods in the test assembly, as predicted by ANF-B and SPCB, were then | |||
summed to yield the prediction of the total number of rods experiencing boiling | |||
transition. The ANF-B and SPCB correlations were found to conservatively | |||
overpredict the expected number of rods that experience boiling transition (References 16, 17, 25 and 26). | |||
4.4.2.3 Maximum Average Planar Linear Heat Generation Rate (MAPLHGR) | |||
The MAPLHGR limit for fuel assures that the peak cladding temperature of fuel | |||
following a postulated design basis loss-of-coolant accident (LOCA) will not exceed | |||
the peak cladding temperature (PCT) and maximum oxidation limits specified in | |||
10CFR50.46. The calculational procedure used to establish the MAPLHGR limits is | |||
based on a LOCA analysis. The analysis is performed using calculational models LSCS-UFSAR 4.4-6 REV. 18, APRIL 2010 which are consistent with the requirements of Appendix K to 10CFR50. The models are described in Reference 20 for AREVA and Reference 12 for GE. | |||
The PCT following a postulated LOCA is primarily a function of the average heat | |||
generation rate of all the rods of a fuel assembly at any axial location and not strongly influenced by the rod-to-rod power distribution within the assembly. | |||
The MAPLHGR limits for two-loop operation for a particular cycle are specified in | |||
the COLR. | |||
For single-loop operation, an APLHGR limit corresponding to the product of the | |||
two-loop limit and a reduction factor specified in the COLR can be conservatively | |||
used to ensure that the PCT for single-loop operation is bound by the PCT for two- | |||
loop operation. | |||
4.4.2.3.1 Design Power Distribution Thermal-hydraulic design of the reactor -- including the selection of the core size and effective heat transfer area, the design steam quality, the total recirculation | |||
flow, the inlet subcooling, and the specification of internal flow distribution -- is | |||
based on the concept and application of a design power distribution. The design | |||
power distribution represents a conservative thermal operating state at rated | |||
conditions and includes design allowances for the combined effects (on the fuel rod, and the fuel assembly heat flux and temper ature) of the gross and local steady-state power density distributions and adjustments of the control rods. | |||
The design power distribution is used in conjunction with flow and pressure drop | |||
distribution computations to determine the thermal conditions of the fuel and the | |||
enthalpy conditions of the coolant throughout the core. | |||
The design power distribution is based on detailed calculations of the neutron flux | |||
distribution. | |||
LSCS-UFSAR 4.4-7 REV. 18, APRIL 2010 The core average and maximum void fractions are dependent on the reactor operating state and power distributions. Typical average and maximum void | |||
fraction results for AREVA fuel can be found in Reference 13. | |||
4.4.2.4 Void Fraction Distribution | |||
The core average and maximum void fractions for the initial core at rated condition | |||
are given in Table 4.4-1. The typical axial distribution of core void fractions for the | |||
average radial channel and the maximum radial channel (end of node value) is | |||
given in Table 4.4-2. The core average and maximum exit value are also provided. | |||
Similar distributions for steam quality are provided in Table 4.4-3. The core | |||
average axial power distributions used to produce these tables are given in | |||
Table 4.4-2a. | |||
4.4.2.5 Core Coolant Flow Distribution | |||
Correct distribution of core coolant flow among the fuel assemblies is accomplished | |||
by the use of an accurately calibrated fixed or ifice at the inlet of each fuel assembly. | |||
The orifices are located in the fuel support piece. They serve to control the flow | |||
distribution and, hence, the coolant conditions within prescribed bounds throughout | |||
the design range of core operation. | |||
The core is divided into two orificed flow zones. The outer zone is a narrow, reduced-power region around the periphery of the core. The inner zone consists of | |||
the core center region. No other control of flow and steam distribution other than | |||
that incidentally supplied by adjusting the power distribution with the control rods, is used or needed. The orifices can be changed during refueling, if necessary. | |||
The sizing and design of the orifices ensure stable flow in each fuel assembly during all phases of operation at normal operating conditions. | |||
Design core flow distribution calculations are made using the design power | |||
distribution which consists of a hot and average powered assembly in each of the two orifice zones. Typical design bundle powers and resulting relative flow | |||
distributions are given in Table 4.4-4. | |||
The flow distribution to the fuel assemblies is calculated on the assumption that the | |||
pressure drop from lower plenum to upper plenum (across all fuel assemblies) is the same. This assumption has been confirmed by measuring the flow distribution in a modern boiling water reactor as reported in Reference 2. | |||
LSCS-UFSAR 4.4-8 REV. 18, APRIL 2010 There is reasonable assurance, therefore, that the calculated flow distribution throughout the core is in close agreement with the actual flow distribution of an operating reactor. The use of the design power distribution discussed previously | |||
ensures that the chosen orificing covers the range of normal operation. The expected | |||
shifts in power production during core life are less severe and are bounded by the design | |||
power distribution. | |||
4.4.2.6 Core Pressure Drop and Hydraulic Loads | |||
The pressure drop across various core compon ents under steady-state design conditions is included in Table 4.4-1 for the initial core. Initial Cycle analyses for the most limiting | |||
conditions, the recirculation line break and the steamline break, are reported in Chapter | |||
: 15. For SAFER/GESTR information, see Reference 12. For core pressure drop | |||
information for AREVA fuel, see Reference 13. | |||
The components of bundle pressure drop considered are friction, local, elevation, and | |||
acceleration pressure drops. Pressure drop measurements made in operating reactors | |||
confirm that the total measured core pressure drop and calculated core pressure drop | |||
are in good agreement. | |||
Subsections 4.4.2.6.1 through 4.4.2.6.4 describe the pressure drop models that were used | |||
by GE for the initial core. AREVA utilizes similar pressure drop correlations and methodology. For more detail on these correlations and methodologies see References 14 | |||
and 15. | |||
4.4.2.6.1 Friction Pressure Drop | |||
Friction pressure drop is calculated using the relationship: | |||
where: p f = friction pressure drop, psi, w = mass flow rate, g = acceleration of gravity, = water density, D H = channel hydraulic diameter, A ch = channel flow area, L = length, ()1-4.4 AD fL 2g w P2 TPF 2chH 2 f= | |||
LSCS-UFSAR 4.4-9 REV. 14, APRIL 2002 f = friction factor, and 2 TPF = two phase friction multiplier. | |||
This formulation is similar to that used throughout the nuclear power industry. The | |||
formation for the two-phase multiplier is based on data which compares closely to that | |||
found in the open literature (Reference 3). | |||
4.4.2.6.2 Local Pressure Drop | |||
The local pressure drop is defined as the i rreversible pressure loss associated with an area change such as the orifice, lower tie-plate, and spacers of a fuel assembly. | |||
The general local pressure drop model is similar to the friction pressure drop and is | |||
given by: | |||
where: | |||
P L = local pressure drop, psi; K = local pressure drop loss coefficient; A = reference area for local loss coefficient; and 2 TPL = two-phase local multiplier and w, g, and are defined the same as for friction. This basic calculation is similar to that used throughout the nuclear power industry. The formulation for the two-phase multiplier is similar to that reported in the open literature (Reference 4) with the addition of empirical constants to adjust the results to fit data taken at General | |||
Electric Company for the specific designs of the BWR fuel assembly. | |||
4.4.2.6.3 Elevation Pressure Drop | |||
() 2-4.4 A K 2g w P 2 TPL 2 2 L= ()()3-4.4 -1 L; P g f E+== | |||
LSCS-UFSAR 4.4-10 REV. 14, APRIL 2002 The elevation pressure drop is based on the well-known relationship where: | |||
P E = elevation pressure drop, psi; L = incremental length: | |||
= average coolant density: | |||
= average void fraction over length -L; and f , g = saturated water and vapor density, respectively. | |||
4.4.2.6.4 Acceleration Pressure Drop | |||
A reversible pressure change occurs when an area change is encountered, and an | |||
irreversible loss occurs when the fluid is accelerated through the boiling process. | |||
The basic formulation for the reversible pressure change resulting from a flow area | |||
change is given by: | |||
where: P ACC = acceleration pressure drop, A 2 = final flow area, and A 1 = initial flow area | |||
and other terms are as previously defined. The basic formulation for the | |||
acceleration pressure change due to density change is: | |||
where: ()()4-4.4 | |||
A A ;A 2 w 1P 1 2 2 2 2 2 ACC==g ()5-4.4 | |||
1 1A g w P IN M OUT M 2 ch 2 ACC= ()(), -1 x1g x1 f 2 2 M+= | |||
LSCS-UFSAR 4.4-11 REV. 18, APRIL 2010 M = momentum density, and x = steam quality and other terms are as previously defined. The total acceleration pressure drop in | |||
boiling water reactors is on the order of a few percent of the total pressure drop. | |||
4.4.2.7 Correlation and Physical Data | |||
The General Electric Company has obtained substantial amounts of physical data | |||
in support of the pressure drop and thermal hydraulic loads discussed in Subsection | |||
4.4.2.6. Correlations have been developed to fit this data to the formulations | |||
discussed. | |||
Subsection 4.4.2.7.1 through 4.4.2.7.3 de scribe the thermal hydraulic correlations | |||
used by GE for the initial core. AREVA has also qualified their thermal hydraulic correlations for use in calculating pressure drop, void fraction, and heat transfer in | |||
References 14 and 15. | |||
4.4.2.7.1 Pressure Drop Correlations | |||
The General Electric Company has taken significant amounts of friction pressure | |||
drop data in multirod geometries representative of modern BWR plant fuel bundles | |||
and correlated both the friction factor and two-phase multipliers on a best fit basis | |||
using the pressure drop formulations reported in Subsections 4.4.2.6.1 and | |||
4.4.2.6.2. Tests are performed in single-phase water to calibrate the orifice and the | |||
lower tie-plate, and in both single- and two-phase flow to arrive at best fit design | |||
values for spacer and upper tie-plate pressure drop. | |||
The range of test variables is specified to include the range of interest to boiling | |||
water reactors. New data are taken whenever there is a significant design change | |||
to ensure the most applicable methods are in use at all times. | |||
Applicability of the single-phase and two-phase hydraulic models discussed in | |||
Subsections 4.4.2.6.1 and 4.4.2.6.2 is conf irmed by prototype (64-rod bundle) flow tests. The typical range of the test data is summarized in Table 4.4-5. | |||
4.4.2.7.2 Void Fraction Correlation | |||
The void fraction correlation is similar to models used throughout the nuclear | |||
power industry and includes effects of pressure, flow direction, mass velocity, quality, and subcooled boiling. | |||
LSCS-UFSAR 4.4-12 REV. 18, APRIL 2010 4.4.2.7.3 Heat Transfer Correlation The Jens-Lottes (Reference 5) heat transfer correlation is used in fuel design to | |||
determine the cladding-to-coolant heat transfer coefficient for nucleate boiling. | |||
4.4.2.8 Thermal Effects of Operational Transients | |||
The evaluation of the core's capability to withstand the thermal effects resulting | |||
from anticipated operational transients is covered in Chapter 15 and Appendix G. | |||
In summary, all transients due to normal operation and to single operator error or | |||
equipment malfunction result in MCPR greater than the transient MCPR limit. | |||
4.4.2.9 Uncertainties in Estimates | |||
Uncertainties in thermal-hydraulic parameters are considered in the statistical | |||
analysis which is the basis for setting the transient MCPR limit such that at least | |||
99.9% of the fuel rods in the core are expected not to experience boiling transition | |||
during any abnormal operating transient. The statistical model and analytical | |||
procedure are described in detail in References 1 and 11. The uncertainties | |||
considered and their input values for the analysis are given in References 1 and 11. | |||
For AREVA fuel, the statistical models and the methodology for calculating the MCPR safety limit are described in References 16, 17, 21, 25 and 26. | |||
4.4.2.9.1 Transition Boiling Uncertainties | |||
The fuel cladding employed for the nuclear fuel is Zircaloy. This material is | |||
selected primarily for its nuclear properties. Zircaloy also has good corrosion and | |||
strength properties at normal operating conditions. However, continued operation | |||
at the elevated temperatures possible in the transition and film boiling regimes | |||
could cause gradual reduction in strength and accelerated corrosion, resulting in | |||
damage to the cladding. | |||
The boiling transition does not necessarily correspond to the fuel damage threshold, especially in the high steam-quality range. Boiling transition is identified as the | |||
heat transfer rate below which cladding overheating does not occur. Damage would | |||
not actually occur until well into the film boiling regime. For example, during inpile | |||
tests (Reference 6), Zircaloy-clad uranium dioxide fuel was purposely operated at | |||
heat transfer rates well into film boiling for a total time exceeding 5 minutes, then | |||
operated at typical boiling water reactor conditions for 10 days. Post-irradiation | |||
examination showed evidence of overheating but no cladding failure. To ensure | |||
good performance and long life of the cladding, conservative limits have been | |||
established to ensure that normal operations remain well below the transition | |||
boiling regime. | |||
LSCS-UFSAR 4.4-13 REV. 14, APRIL 2002 4.4.2.9.2 Variation of Fuel Damage Limit Incipient center melting of the uranium dioxide pellet occurs at a higher kW/ft than | |||
the peak LHGR during any abnormal operating transient. If UO 2 center melting occurs and the molten uranium dioxide is redistributed and densified, the damage | |||
limit for strain can reduce to a lower value. The redistribution and densification | |||
phenomena are functions of time and temperature. Plant transients of short | |||
duration in the molten range do not result in appreciable redistribution or | |||
densification. For the plant events that meet the transient MCPR limit, there is no | |||
appreciable change in the kW/ft damage limit. | |||
4.4.2.9.3 Effects of Misoriented Fuel Bundle | |||
The concern with a misoriented assembly is primarily that the redistribution of | |||
power among the fuel pins could lead to higher local powers than indicated by the | |||
core monitoring system. In addition, a miso rientation could lead to slightly higher assembly powers as well. A detailed descri ption of this evaluation may be found in section 15.4.7. | |||
4.4.2.10 Flux Tilt Considerations | |||
For flux tilt considerations, refer to Subsection 4.3.2.2.7. | |||
====4.4.3 Description==== | |||
of the Thermal and Hydraulic Design of the Reactor Coolant System | |||
The thermal and hydraulic design of the reactor coolant system is described in this | |||
subsection. | |||
4.4.3.1 Plant Configuration Data | |||
The descriptive summary of the reactor coolant system is given in Section 5.1. That | |||
overview describes the reactor coolant pressure boundary and the reactor coolant | |||
equipment used for the various coolant requirements encountered in both normal | |||
and abnormal operations. The engineered safety functions are described in | |||
Chapter 6.0 with system details and analysis shown there. The reactor | |||
recirculation loops are described in detail in Subsection G.2.3 of Appendix G; The | |||
main steam and feedwater systems are treated in Section 5.4. Plant configuration | |||
data are included in these chapters. | |||
Table 4.4-7 provides the flow path length, height, liquid level, minimum elevations, and minimum flow areas for each major flow path volume within the reactor vessel | |||
and recirculation loops of the reactor coolant system. Table 4.4-8 provides the | |||
lengths and sizes of all safety injectio n lines to the reactor coolant system. | |||
LSCS-UFSAR 4.4-14 REV. 14, APRIL 2002 4.4.3.2 Operating Restrictions on Pumps See Subsection G.2.2 of Appendix G. | |||
4.4.3.3 Power-Flow Operating Map | |||
See Subsection G.2.3 of Appendix G. | |||
4.4.3.4 Temperature-Power Operating Map (PWR) | |||
Not applicable. | |||
4.4.3.5 Load-Following Characteristics | |||
See Subsection G.2.4 of Appendix G. | |||
4.4.3.6 Thermal and Hydraulic Characteristics Summary Table | |||
A summary of the thermal and hydraulic char acteristics of the reactor coolant system for the initial core and the initial cores of other reactors of similar design is included in | |||
Table 4.4-1. | |||
====4.4.4 Evaluation==== | |||
The thermal-hydraulic design of the reactor core and reactor coolant system is based | |||
upon an objective of no fuel damage during normal operation or during abnormal | |||
operational transients. This design objective is demonstrated by analysis in the | |||
following sections. | |||
4.4.4.1 Critical Heat Flux | |||
Table 4.4-1 provides data on maximum heat flux, average heat flux, heat transfer areas, and other parameters affecting heat transfer of the initial core. The concept of | |||
critical heat flux has been used in the determination of operationally significant power | |||
distribution constraints. These are given in terms of the linear heat generation rate and minimum critical power ratio as discussed in the following subsections. | |||
4.4.4.2 Core Hydraulics | |||
See Subsection G.2.3 of Appendix G. | |||
4.4.4.3 Influence of Power Distribution | |||
The design constraints imposed by the maximum average planar linear heat | |||
generation rate, the core power density, and the local peaking factor limit the gross LSCS-UFSAR 4.4-15 REV. 18, APRIL 2010 peaking factor (radial x axial). There are many combinations of radial and axial peaking factors that satisfy this design constraint, but each will have a different effect | |||
on the MCPR. In general, the MCPR decreas es as the radial peaking (bundle power) increases and as the axial peak location moves to the top of the core. For example, for | |||
a 1.96 gross factor, a flat (1.0) axial and a 1.96 radial would give a relatively low CPR, whereas a 1.0 radial and a 1.96 axial peaked in the bottom of the core would give a | |||
relatively high CPR. These extremes are obviously not suited to design because they | |||
are not representative of realistic reactor behavior. Therefore, the design radial | |||
peaking factor is selected higher than that likely to be encountered in reactor | |||
operation, and the combination of this radial with the design axial profile is also more | |||
limiting than that expected during operating conditions. | |||
4.4.4.4 Core Thermal Response | |||
The thermal response of the core evaluated for expected transient conditions is | |||
covered in Chapter 15. All expected abnormal operational transients are | |||
conservatively evaluated to ensure that the integrity of the vessel and fuel is not | |||
compromised. These transients are analyzed at varying power and flow conditions | |||
within the analyzed power-to-flow map. | |||
4.4.4.5 Analytical Methods | |||
The analytical methods, thermodynamic data, and hydrodynamic data used in | |||
determining the thermal and hydraulic characteristics of the core are similar to those | |||
used throughout the nuclear power industry. | |||
Core thermal-hydraulic analyses are performed with the aid of a digital computer | |||
program. This program models the reacto r core through a hydraulic description of orifices, lower tie-plates, fuel rods, fuel rod spacers, upper tie-plates, fuel channel, and | |||
the core bypass flow paths. | |||
The methods discussed in section 4.4.4.5.1 through 4.4.4.5.3 describe the analytical | |||
methods for GE. However, the descriptions below are typical for the nuclear industry. | |||
These descriptions apply generally to AREVA methods. Further detailed descriptions of AREVA methods can be found in References 14, 17, 25 and 26. | |||
4.4.4.5.1 Reactor Model | |||
The orifice, lower tie-plate, fuel rod spacers, and upper tie-plate are hydraulically | |||
represented as being separate, distinct local losses of zero thickness. The fuel channel cross section is represented by a square section with enclosed area equal to the | |||
unrodded cross-sectional area of the actual fu el channel. The fuel channel assembly consists of three basic axial regions. The first and most important is the active fuel | |||
region which consists of the fuel rods, nonfueled rods, and fuel-rod spacers. The | |||
second is the nonfueled region consisting of nonfueled rods and the upper tie-plate. | |||
LSCS-UFSAR 4.4-16 REV. 14, APRIL 2002 The third region represents the unrodded portion of the fuelchannel above the upper tie-plate. The active fuel region is considered in independent axial segments | |||
or nodes over which fuel thermal properties are assumed constant and coolant | |||
properties are assumed to vary linearly. The code can handle 12 fuel channel types and 10 types of bypass flow paths. In normal analyses the fuel assemblies are | |||
modeled by four channel types--a hot centra l orifice region channel type, an average central orifice region channel type, a hot peripheral orifice region type and an | |||
average peripheral orifice region type. | |||
Usually there is one fuel assembly representing each of the hot types. The average types then make up the balance of the core. | |||
The computer program iterates on flow through each flow path (fuel assemblies and | |||
bypass paths) until the total differential pr essure (plenum to plenum) across each path is equal, and the sum of the flows thro ugh each path equals the total core flow. | |||
Orificing is selected to optimize the core flow distribution between orifice regions as | |||
discussed in Subsection 4.4.2.5. The core design pressure is determined from the | |||
required turbine throttle pressure, the steamline pressure drop, steam dryer | |||
pressure drop, and the steam separator pre ssure drop. The core inlet enthalpy is determined from the reactor and turbine heat balances. The core power | |||
distribution is determined as per Subsection 4.4.2.3. The required core flow is then | |||
determined by applying the procedures of this section and specifications such that | |||
the thermal limits of Reference 11 are satisfied and the nominal expected bypass | |||
flow fraction is approximately 10%. The results of applying these methods and | |||
specifications are: | |||
: a. flow for each bundle type, | |||
: b. flow for each bypass path, | |||
: c. core pressure drop, | |||
: d. fluid property axial distribution for each bundle type, and | |||
: e. CPR calculations for each bundle type. | |||
4.4.4.5.2 System Flow Balances | |||
The basic assumption used by the code in performing the hydraulic analysis is that the flow entering the core will divide itself between the fuel bundles and the bypass flow paths such that each assembly and bypass flow path experience the same | |||
pressure drop. | |||
LSCS-UFSAR 4.4-17 REV. 14, APRIL 2002 The bypass flow paths considered are described in Table 4.4-9 and shown in Figure 4.2-2. Due to the large flow area , the pressure drop in the bypass region above the core plate is essentially all elevation head. Thus, the sum of the core | |||
plate differential pressure and the bypass region elevation head is equal to the core | |||
differential pressure. | |||
The total core flow less the control rod cooling flow enters the lower plenum through | |||
the jet pumps. A fraction of this passe s through the various bypass paths. The remainder passes through the orifice in the fuel support (experiencing a pressure | |||
loss) where more flow is lost through the fit-up between the fuel support and the lower tie-plate into the bypass region. The majority of the flow continues through | |||
the lower tie-plate (experiencing a pressure loss) where some flow is lost through | |||
the flow path defined by the fuel channel and lower tie-plate, and restricted by the finger springs, into the bypass region. | |||
The flow through the bypass flow paths are expressed by the form: | |||
Full scale tests have been performed to establish the flow coefficients for the major | |||
flow paths. These tests simulate actual plant configurations which have several | |||
parallel flow paths and, therefore, the fl ow coefficients for the individual paths could not be separated. However, analytical models of the individual flow paths | |||
were developed as an independent check of the tests. The models were derived for | |||
actual BWR design dimensions and considered the effects of dimensional variations. | |||
These models predicted the test results when the "as-built" dimensions were | |||
applied. When using these models for hydraulic design calculations, nominal | |||
drawing dimensions are used. This is done to yield the most accurate prediction of | |||
the expected bypass flow. With the larg e number of components in a typical BWR core, deviations from the nominal dimensions will tend to statistically cancel, resulting in a total bypass flow best represented by that calculated using nominal | |||
dimensions. | |||
The balance of the flow enters the fuel bundle from the lower tie plate and passes | |||
through the fuel rod channel spaces. A small portion of the in-channel flow enters | |||
the non-fueled rod through orifice holes just above the lower tie-plate. This flow, normally referred to as the water-rod flow, remixes with the active coolant channel | |||
flow below the upper tie-plate. | |||
4.4.4.5.3 System Heat Balances | |||
Within the fuel assembly, heat balances on the active coolant are performed | |||
nodally. Fluid properties are expressed as the bundle average at the particular | |||
node of interest and are based on Reference 7. In evaluating fluid properties a | |||
constant pressure model is used. | |||
()6-4.4 P C PCP CW 2 3 4 C 221 1++= | |||
LSCS-UFSAR 4.4-18 REV. 14, APRIL 2002 The core power is divided into two parts: an active coolant power and a bypass flow power. The bypass flow is heated by neutron-slowing down and gamma heating in the | |||
water and by heat transfer through the channe l walls. Heat is also transferred to the bypass flow from structures and control elements which are themselves heated by gamma absorption and by (n, ) reactions in the control material. The fraction of total reactor power deposited in the bypass region is very nearly 2%. A similar phenomena occurs with the fuel bundle to the active coolant and the water rod flows. The net | |||
effect is that approximately 96% of the core power is conducted through the fuel | |||
cladding and appears as heat flux. | |||
The power is allocated to the individual fuel bundles using a relative power factor. | |||
The power distribution along the length of th e fuel bundle is specified with axial power factors which distribute the bundle's power among the axial nodes. A nodal location | |||
power or peaking factor is used to establish the peak heat flux at each nodal location. | |||
Relative, axial, and local peaking factors are more thoroughly discussed in Subsection | |||
4.3.2. | |||
The relative (radial) and axial power distributions when used with the bundle flow | |||
determine the axial coolant property distribution resulting in sufficient information to | |||
calculate the pressure drop components with in each fuel assembly type. Once the equal pressure drop criterion has been sati sfied, the critical bundle power (the power which would result in critical quality existing at some point in the bundle using the | |||
correlation expressed in References 1 and 11) is determined by an iterative process for | |||
each fuel type. | |||
In applying the above methods to core design, the number of bundles (for a specified | |||
core thermal power) and bundle geometry (8 x 8, rod diameter, etc.) are selected based | |||
on power density and linear heat generation rate limits. | |||
4.4.4.5.4 Uncertainties in Design Analyses | |||
The effects of uncertainties in design values and on calculational results are accounted | |||
for in the statistical analysis on which the MCPR limits are based. | |||
4.4.4.6 Reactor Stability Analysis | |||
4.4.4.6.1 Introduction | |||
There are many definitions of stability, but for feedback processes and control systems | |||
it can be defined as follows: a system is stable if, following a disturbance, the | |||
transient settles to a steady, noncyclic state. | |||
A system may also be acceptably safe even if oscillatory, provided that any limit cycle of the oscillations is less than a prescribed magnitude. Instability then, is either a | |||
continual departure from a final steady-state value or greater-than-prescribed limit | |||
cycle about the final steady-state value. | |||
LSCS-UFSAR 4.4-19 REV. 14, APRIL 2002 The mechanism for instability can be explained in terms of frequency response. | |||
Consider a sinusoidal input to a feedback control system which for the moment has | |||
the feedback disconnected. If there were no time lags or delays between input and | |||
output, the output would be in phase with the input. Connecting the output so as to | |||
subtract from the input (negative feedback or 180º out-of-phase connection) would | |||
result in stable closed loop operation. However, natural laws can cause phase shift | |||
between output and input and should the phase shift reach 180º, the feedback | |||
signal would be reinforcing the input signal rather than subtracting from it. If the | |||
feedback signal were equal to or larger than the input signal (loop gain equal to one | |||
or greater), the input signal could be disconnected and the system would continue to | |||
oscillate. If the feedback signal were less than the input signal (loop gains less than | |||
one), the oscillations would die out. | |||
The design of the BWR is based on the premise that power oscillations can be | |||
readily detected and suppressed. | |||
4.4.4.6.2 Description | |||
Three types of stability considered in the design of boiling water reactors are (1) | |||
reactor core (reactivity) stability, (2) channel hydrodynamic stability, and (3) total | |||
system stability. Reactivity feedback instability of the reactor core could drive the | |||
reactor into power oscillations. Hydrodynamic channel instability could impede | |||
heat transfer to the moderator and drive the reactor into power oscillations. The | |||
total system stability considers control system dynamics combined with basic | |||
process dynamics. The criteria is demonstrated if it is analytically demonstrated | |||
that no divergent oscillation develops within the system as a result of calculated | |||
step disturbances of any critical variable, such as steam flow, pressure, neutron | |||
flux, and recirculation flow, or that the divergent oscillation can be detected and suppressed. | |||
Stability is expressed in terms of two compatible parameters. First is the decay | |||
ratio x 2/x 0 , designated as the ratio of the magnitude of the second overshoot to the first overshoot resulting from a step perturbation. A plot of the decay ratio is a | |||
graphic representation of the physical responsiveness of the system, which is readily evaluated in a time-domain analysis. Second is the damping coefficient n, the definition of which corresponds to the pole pair closest to the j axis in the s-plane for the system closed loop transfer function. This parameter also applies to the frequency-domain interpretation. The damping coefficient is related to the decay ratio as shown in Figure 4.4-1. | |||
4.4.4.6.3 Solution Description for Thermal-Hydraulic Stability | |||
BWR cores may exhibit thermal-hydraulic instabilities in certain portions of the | |||
core power and recirculation flow operating domain. The instabilities and the | |||
solutions devised to detect and suppress them are discussed in Reference 22 and 23. | |||
LSCS-UFSAR 4.4-20 REV. 17, APRIL 2008 LSCS has adopted the solution Option III, designated as the Oscillation Power Range Monitor (OPRM). The OPRM complies with GDC-12, as discussed in | |||
Section 3.1.2.2.3. | |||
The overall design philosophy of the OPRM is to generate an alarm in the control | |||
room if it detects core instabilities (based on period-based algorithm only), and to generate an automatic suppression system trip if the instabilities reach an | |||
amplitude that could threaten the fuel safety limits. | |||
The overall objective of the oscillation detection algorithm is to reliably detect expected instabilities at a low magnitude such that mitigation can occur well before the MCPR Safety Limit is exceeded, while avoiding spurious trips during expected neutron flux transients. The algorithm is based on the detection of the three known characteristics that BWR neutron flux oscillations exhibit. These characteristics are the amplitude or absolute magnitude, growth rate, and periodic behavior. Only the period based detection algorithm is used in the safety analysis. The other algorithms provide defense in depth and additional protection against unanticipated oscillations. Details of the algorithm can be found in References 22 and 23. | |||
The OPRM consists of a micoprocessor that analyzes signals from LPRMs. Since LPRMs are evenly distributed throughout the reactor core, they are capable of | |||
responding to any neutron flux oscillations that can create an MCPR concern. | |||
Individual LPRMs readily respond to a wide variety of normal operating maneuvers | |||
and expected events, and are also subject to electrical interference. For these | |||
reasons, each OPRM may use multiple LPRMs as a means of maintaining a strong | |||
response to a neutron flux oscillation while minimizing the susceptibility to false | |||
signals associated with a single LPRM, or may utilize a detection algorithm designed to achieve the same objective. The OPRM is automatically bypassed at | |||
high flow or low power conditions, where core instabilities are unlikely to occur, to | |||
avoid spurious actuation. | |||
4.4.4.6.4 Stability Criteria | |||
The following discussion on stability is based on the original design bases, which did not assume an inherent tendency towards oscillations. They are presented here | |||
for historic perspective. The new design, in compliance with the NRC Generic | |||
Letter 94-02, is based on the detection and suppression methodology, and is | |||
discussed above in Section 4.4.4.6.3. | |||
Stability criteria are established to demonstrate compliance with the requirements | |||
set forth in 10CFR50 Appendix A, General Design Criterion (GDC) 12. | |||
These stability compliance criteria consider potential limit cycle response within the | |||
limits of safety system and/or operator in tervention and the OPRM assures that for BWR fuel designs this operating mode does not result in specified acceptable LSCS-UFSAR 4.4-21 REV. 17, APRIL 2008 fuel design limits being exceeded. The onset of power oscillations for which corrective actions are necessary is reliably and readily detected and suppressed by | |||
operator actions and/or automatic system functions. | |||
To ensure compliance of the GE BWR design with GDC 12 requirements, the following stability acceptance criteria have been established. | |||
(1) Neutron flux limit cycles which oscillate up to the 120% APRM high neutron flux scram set point or up to the LPRM upscale | |||
alarm trip (without initiating scram) prior to operator | |||
mitigating action, shall not result in exceeding specified | |||
acceptable fuel design limits. | |||
(2) The individual channels shall be designed and operated to be hydrodynamically stable or more stable than the reactor core for | |||
all expected operating conditions. | |||
Calculations which predict that core-wide limit cycles will not occur (decay ratio | |||
< 0.8) also demonstrate compliance with GDC-12. | |||
This criteria is presently used for LSCS two recirculation loop operation. For single | |||
recirculation loop operation, the plant is monitored per General Electric SIL-380 (Reference 8). | |||
These criteria shall be satisfied for all attainable conditions of the reactor that may | |||
be encountered in the course of plant operation. For stability purposes, the most | |||
severe conditions to which these criteria will be applied correspond to natural | |||
circulation flow at a power corresponding to the extrapolated APRM rod block | |||
intercept condition. | |||
The licensing basis is to generate a trip signal during oscillations of sufficiently low | |||
amplitude to provide margin to the MCPR safety limits for all expected modes of BWR oscillations. The OPRM oscillation recognition algorithm is intended to | |||
discriminate between true stability-related neutron flux oscillations and other flux | |||
variations that may be expected during plant operation. Extensive evaluation of | |||
operating plant data is done to determine the combination of algorithm and OPRM | |||
setpoints, which meet the design objectives. The final algorithm/setpoint design is | |||
subjected to in-plant testing with the trip function disabled. | |||
The OPRM assures that for BWR fuel designs, this operating mode does not result | |||
in specified acceptable fuel design limits being exceeded. The onset of power | |||
oscillations for which corrective actions are necessary is reliably and readily | |||
detected and suppressed by operator actions and/or automatic system functions. | |||
LSCS-UFSAR 4.4-22 REV. 17, APRIL 2008 4.4.4.6.5 Expected Oscillation Modes | |||
The OPRM is capable of responding to the expected modes of BWR stability-related | |||
oscillations. The expected oscillation modes are as follows (Reference 13, Section 6.1): | |||
* Core-wide, in which the average neutron flux in all fuel assemblies oscillates | |||
in phase. | |||
* First Order Side-by-Side or a regional oscillation where the neutron flux on | |||
one side of the reactor oscillates 180 o out of phase with the flux on the other side. | |||
* First Order Precession a regional oscillation where the axis of zero oscillation amplitude rotates azimuthally, or the two reactor regions of peak oscillation | |||
amplitude shift from one location to another at a frequency lower than the | |||
oscillation frequency. | |||
Other modes of oscillation are not expected in a BWR. | |||
4.4.4.6.6 Analysis Approach | |||
The total system stability analysis evaluates the relative stability of the total | |||
system, from time responses generated by applying step changes to the input variables to the total system stability model. The observed time response of an | |||
output variable of a high order dynamic system represents a superposition of the | |||
system's several response modes. The relative intensity of each particular mode in | |||
the time response is determined by the zeroes (the roots of the numerator) of the | |||
transfer function relating a given output variable to a particular input. Therefore, in judging the relative stability of the sy stem, the observer should separate the distinct modes in the time response and apply the stability criterion to each modal | |||
response. The approach used here, of disturbing one input variable and applying | |||
the stability criterion to the resulting system response is a good approximation to | |||
modal separation. It is particularly applicable in calculating stability since, as a | |||
system tends toward instability a single oscillatory mode tends to dominate the | |||
observed time response (Reference 9). | |||
LSCS-UFSAR 4.4-23 REV. 17, APRIL 2008 Reference 15 describes the process used to calculate a conservative final MCPR value for an anticipated stability-related oscillation. It involves the determination of | |||
initial MCPR by a cycle-specific evaluation and the calculation of hot bundle | |||
oscillation magnitude. The licensing crite rion is met when the final MCPR is greater than the MCPR safety limit. Appr opriate reload parameters are checked every cycle to determine the initial MCPR. This methodology provides a | |||
conservative means of demonstrating with a high probability and confidence that | |||
the MCPR safety limits will not be violated for anticipated oscillations. The use of the MCPR safety limit to provide protection against possible fuel damage is | |||
exceedingly conservative (Reference 24, Section 4.5.2). | |||
4.4.4.6.7 Mathematical Model | |||
This mathematical model applies to the initial core analysis. The mathematical | |||
model representing the core examines the linearized reactivity response of a reactor | |||
system with density-dependent reactivity feedback caused by boiling. In addition, the hydrodynamics of various hydraulically coupled reactor channels or regions are | |||
examined separately on an axially multin oded basis by grouping various channels that are thermodynamically and hydraulically similar. This interchannel | |||
hydrodynamic interaction or coupling exists through pressure variations in the inlet plenum, such as can be caused by disturbances in the flow distribution between | |||
regions or channels. This approach provides a reasonably accurate, three-dimensional representation of the reactor's hydrodynamics. | |||
The core model, shown in block diagram fo rm in Figure 4.4-2, solves the dynamic equations that represent the reactor core in the frequency domain. From the | |||
solution of these dynamic equations, the reactivity and individual channel | |||
hydrodynamic stability of the boiling water reactor is determined for a given reactor | |||
flow rate, power distribution, and total power. This gives the most basic | |||
understanding of the inherent core behavior (and hence the system behavior) and is | |||
the principal consideration in evaluating the stable performance of the reactor. As LSCS-UFSAR 4.4-24 REV. 17, APRIL 2008 new experimental or reactor operating data are obtained, the model is refined to improve its capability and accuracy. | |||
The plant model considers the entire reactor system, neutronics, heat transfer, hydraulics, and the basic processes, as well as associated control systems such as | |||
the flow controller, pressure regulator, feedwater controller, etc. Although the | |||
control systems may be stable when analyzed individually, final control system | |||
settings must be made in conjunction with the operating reactor so that the entire | |||
system is stable. The plant model yields results that are essentially equivalent to | |||
those achieved with the core model and allows the addition of the controllers, which | |||
have adjustable features permitting the attainment of the desired performance. | |||
The plant model solves the dynamic equations that present the BWR system in the | |||
time domain. The variables, such as steam flow and pressure, are represented as a | |||
function of time. The extensiveness of this model (Reference 10) is shown in block diagram form in Figure 4.4-3. Many of the blocks are extensive systems in | |||
themselves. The model is periodically refined as new experimental or reactor | |||
operating data are obtained to improve its capability and accuracy. | |||
4.4.4.6.8 Initial Core Analysis Results | |||
The results of the two recirculation pump operation core and channel stability | |||
analysis is given in the Reload Licensing Package for each cycle. The plant stability | |||
analysis is performed only for the initial core and is described below. | |||
The plant stability analysis was performed by assuming that the reactor is initially | |||
operating at the most sensitive condition, corresponding to natural circulation flow | |||
and a power level at the rod block limit. The nuclear system is then subjected to | |||
step disturbances from control rods, pressure regulator setpoint, and level controller | |||
setpoint. These time responses are shown in Figures 4.4-6 through 4.4-8. It is clear | |||
that the decay ratio is less than the stability criterion. | |||
For expected normal operating modes, the time response of each of the important | |||
variables of the reactor system (neutron flux, pressure, and steam flow) to small step disturbances can be underdamped, but must analytically show a decay ratio of | |||
less than 0.25 in order to satisfy the operational design guide limit. Using final | |||
design parameters each of the following disturbances are analytically imposed, one | |||
at a time, using the model previously described for time domain analysis: | |||
: a. a pressure setpoint change of at least 5 psi, | |||
: b. a control rod position change equivalent to a local power change of at least 5% of point (of the magnitude of power at the time of the | |||
disturbance), | |||
LSCS-UFSAR 4.4-25 REV. 17, APRIL 2008 c. a load demand change of at least 5% of point, and d. a reactor water level setpoint change of at least 6 inches. | |||
Using actual design parameters, calculated responses of important nuclear system variables to step disturbances from control rod reactivity, pressure regulator | |||
setpoint, level controller setpoint, and turbine load setpoint are tested for rated | |||
power-flow conditions and at the nominal power corresponding to the lower end of | |||
the automatic power-flow control path. | |||
Results of the analysis for 105% rated power and 100% rated flow are shown in | |||
Figures 4.4-9 through 4.4-12. It is evident that the response meets the stability | |||
criterion. Figures 4.4-13 through 4.4-16 show the results of analysis at the low | |||
limit of the automatic flow-control range. | |||
It is concluded that for all normal operating points over the flow-control range the | |||
decay ratio of the total system response s is less than one-fourth, good dynamic performance is expected, and the ratio conforms with the stability criterion. | |||
====4.4.5 Testing==== | |||
and Verification | |||
See Subsection G.4.3 of Appendix G. | |||
The OPRM, which is installed to detect and suppress thermal-hydraulic Instabilities, is extensively tested using available data from several BWR plants. | |||
After installation, the plant is operated for a period of time with the OPRM trip function disabled while OPRM performance is monitored for susceptibility to spurious trips. The OPRM trip function is enabled following approval of the associated Technical Specification. | |||
====4.4.6 Instrumentation==== | |||
Requirements See Subsections 7.7.3.2 and 7.6.3.4 of Chapter 7. | |||
4.4.6.1 Loose Parts Monitoring System (Deleted) | |||
LSCS-UFSAR 4.4-26 REV. 17, APRIL 2008 This page intentionally left blank. | |||
LSCS-UFSAR 4.4-27 REV. 17, APRIL 2008 | |||
====4.4.7 References==== | |||
: 1. "General Electric BWR Thermal Analysis Basis (GETAB): Data, Correlation, and Design Application," January 1977 (NEDE-10958-PA | |||
and NEDO-10958A). | |||
: 2. "Core Flow Distribution in a Modern Boiling Water Reactor as Measured in Monticello," NEDO-10299, AEC Topical Report | |||
NEDO-10299, January 1971. | |||
: 3. R. C. Martinelli and D. E. Nelson, "Prediction of Pressure Drops During Forced Connection Boiling of Water," ASME Trans., 70, pp. 695-702, 1948. | |||
: 4. C. J. Baroozy, "A Systematic Correlation for Two-Phase Pressure Drop," Heat Transfer Conference, Preprint No. 37, AICLE, Los | |||
Angeles, 1966. | |||
: 5. W. H. Jens and P. A. Lottes, "Analysis of Heat Transfer, Burnout, Pressure Drop, and Density Data for High Pressure Water," USAEC | |||
Report - 4627, 1972. | |||
LSCS-UFSAR 4.4-28 REV. 18, APRIL 2010 | |||
: 6. S. Levy et al., "Experience with BWR Fuel Rods Operating Above Critical Flux," Nucleonics , April 1965. | |||
: 7. 1967 International Standard Steam Water Properties. | |||
: 8. General Electric Service Information Letter (SIL) No. 380, "BWR Core Thermal Hydraulic Stability," Rev. 1, dated February 10, 1984. | |||
: 9. Zadeh and Desoer, "Linear System Theory," McGraw-Hill Book Co., 1963. | |||
: 10. "Analytical Methods of Plant Transient Evaluations for General Electric Boiling Water Reactor," NEDO-10802, General Electric | |||
Company, BWR Systems Department, February 1973. | |||
: 11. "General Electric Standard Application for Reactor Fuel," NEDE-P-A, (Latest approved revision). | |||
: 12. GE Document, "SAFER/GESTR-LOCA, Loss-of-Coolant Accident Analysis, LaSalle County Station Units 1 & 2," NEDC-31510P, as | |||
amended & revised. | |||
: 13. Cycle specific Fuel Design Report (AREVA fuel only). | |||
: 14. Exxon Nuclear Methodology for Boiling Water Reactors: THERMEX Thermal Limits Methodology, Summary Description , XN-NF-80-19(A), Volume 3, Revision 2, Exxon Nuclear Company, Inc., Richland, WA (January 1987). | |||
: 15. Generic Mechanical Design Criteria for BWR Fuel Design , ANF-89-98(P)(A), Revision 1 and Supplement 1, Advanced Nuclear | |||
Fuels Corporation, Richland, WA (May 1995). | |||
: 16. Advanced Nuclear Fuels Critical Power Methodology for Boiling Water Reactors, ANF-524(P)(A), Revision 2 and Supplements, Advanced | |||
Nuclear Fuels Corporation, November 1990. | |||
: 17. ANFB Critical Power Correlation , ANF-1125 (P)(A) and Supplements 1 and 2, Advanced Nuclear Fuels Corporation, April 1990; ANFB Critical Power Correlation Application for Co-Resident Fuel , EMF-1125(P)(A), Supplement 1, Appendix C, Siemens Power Corporation, August 1997; | |||
and ANFB Critical Power Correlation Determination of ATRIUM-9B Additive Constant Uncertainties , ANF-1125(P)(A) Supplement 1, Appendix E, Siemens Power Corporation, September 1998. | |||
LSCS-UFSAR 4.4-29 REV. 18, APRIL 2010 | |||
: 18. Siemens Power Corporation Methodology for Boiling Water Reactors: | |||
Evaluation and Validation of CASMO-4 / MICROBURN-B2 EMF-2158(P)(A) Revision 0, Siemens Power Corporation, October 1999. | |||
: 19. Exxon Nuclear Methodology for Boiling Water Reactors - Neutronic Methods for Design and Analysis , XN-NF-80-19(P)(A) Volume 1 and Supplements 1 and 2, Exxon Nuclear Company, March 1983. | |||
: 20. EXEM BWR-2000 ECCS Evaluation Model EMF-2361(P)(A) | |||
Revision 0, AREVA NP Inc., May 2001. | |||
: 21. MICROBURN-B2 Based Impact of Failed / Bypassed LPRMs and TIPs, Extended LPRM Calibration Interval, and Single Loop Operation Measured Radical Bundle Power Uncertainty , EMF-2493(P) | |||
Revision 0, Siemens Power Corporation, December 2000. | |||
: 22. NEDO-31960, "BWR Owners Group Long-Term Stability Solutions Licensing Methodology," June 1991. | |||
: 23. NEDO 31960, "BWR Owners Group Long-Term Stability Solutions Licensing Methodology," Supplement 1, March 1992. | |||
: 24. NEDO-32465-A, "BWR Owners' Group Reactor Stability Detect and Suppress Solution Licensing Basis Methodology and Reload | |||
Application," August 1996. | |||
: 25. SPCB Critical Power Correlation , EMF-2209(P)(A) Revision 3, AREVA NP, September 2009. | |||
: 26. Application of Siemens Power Corporation's Critical Power Correlation to Co-Resident Fuel , EMF-2245(P)(A), Revision 0, Siemens Power Corporation, August 2000. | |||
LSCS-UFSAR TABLE 4.4-1 (SHEET 1 OF 2) TABLE 4.4-1 REV. 13 THERMAL AND HYDRAULIC DESIGN CHARACTERISTICS OF THE REACTOR CORE (INITIAL CORE DATA) 238-732 218-592 218-560 251-764 251-784 251-764 BWR/6 BWR/6 ZPS-1 WPPSS NP No. 2 BWR/6 LSCS GENERAL OPERATING CONDITIONS Reference design thermal output, MWt 3579 2894 2436 3323 3833 3323 Power level for engineered safety features, MWt 3758 3039 2550 3489 4025 3489 Steam flow rate, at 420° F final feedwater temperature, millions lb/hr 15.396 12.451 10.477 14.295 16.488 14.166 Core coolant flow rate, millions | |||
lb/hr 105.0 84.5 78.5 108.5 113.5 108.5 Feedwater flow rate, millions | |||
lb/hr 15.358 12.42 10.448 14.256 16.488 14.127 System pressure, nominal in steam dome, psia 1040 1040 1020 1020 1040 1020 System pressure, nominal core design, psia 1055 1055 1035 1035 1055 1035 Coolant saturation temperature at core design pressure, °F 551.1 551.1 548.8 548.8 551.1 548.8 Average power density, kW/liter 56 56 50.51 51.2 56.0 48.17 Specific power, kW/kg (U total) 25.9 25.9 23.7 23.7 25.9 23.7 Maximum thermal output, kW/ft 13.4 13.4 13.4 13.4 13.4 13.4 Average thermal output, kW/ft 6.04 6.04 5.45 5.45 6.04 5.33 Core total heat transfer area, ft 2 73,409 59,369 55,401 75,582 78,624 74,871 Maximum heat flux, Btu/hr-ft 2 354,000 354,000 354,000 354,000 354,000 361,000 Average heat flux, Btu/hr-ft 2 159,550 159,550 143,900 143,920 159,550 143,740 LSCS-UFSAR TABLE 4.4-1 (SHEET 2 OF 2) TABLE 4.4-1 REV. 13 238-732 218-592 218-560 251-764 251-784 251-764 BWR/6 BWR/6 ZPS-1 WPPSS NP No. 2 BWR/6 LSCS GENERAL OPERATING CONDITIONS Core inlet enthalpy, at 420° F FFWT, Btu/lb 527.8 527.8 527.4 527.6 528.1 527.5 Core inlet temperature, at 420° F FFWT, °F 533.0 533.0 532.6 532.8 533.3 532.8 Core maximum exit voids within assemblies, % 76 76 75 75 76 76 Core average void fraction, active coolant 0.428 0.429 0.418 0.415 0.427 0.418 Active coolant flow area per | |||
assembly, in 2 15.50 15.50 15.50 15.50 15.50 15.82 Core average inlet velocity, ft/sec 7.2 7.2 7.0 7.1 7.2 6.77 Maximum inlet velocity, ft/sec 7.6 7.6 7.4 7.5 7.6 7.2 Total core pressure drop, psi 25.7 25.5 27.3 27.5 25.8 24.8 Core support plate pressure drop, psi 21.3 21.1 22.9 23.1 21.4 19.61 Average orifice pressure drop Central region, psi 8.6 8.5 11.2 11.4 8.7 8.13 Peripheral region, psi 17.3 17.2 19.6 19.8 17.5 16.66 Maximum channel pressure loading, psi 14.5 14.5 13.7 13.7 14.6 12.84 TYPICAL POWER PEAKING FACTOR Maximum relative assembly power 1.40 1.40 1.40 1.40 1.40 1.40 Local peaking factor 1.13 1.13 1.24 1.15 1.13 1.15 Axial peaking factor 1.40 1.40 1.40 1.40 1.40 1.40 Gross peaking factor 1.96 1.96 1.96 1.96 1.96 1.96 Total peaking factor 2.22 2.22 2.43 2. 2.22 2.25 LSCS-UFSAR TABLE 4.4-2 TABLE 4.4-2 REV. 13 TYPICAL VOID DISTRIBUTION (INITIAL CORE) | |||
NODE CORE AVERAGE (AVERAGE NODE VALUE) MAXIMUM CHANNEL (END OF NODE VALUE) | |||
Bottom 1 0.000 0.0 2 0.001 0.032 3 0.018 0.122 4 0.065 0.230 5 0.136 0.325 6 0.212 0.401 7 0.281 0.462 8 0.341 0.511 9 0.391 0.552 10 0.433 0.587 11 0.469 0.616 12 0.499 0.641 13 0.525 0.662 14 0.547 0.681 15 0.566 0.696 16 0.582 0.708 17 0.595 0.719 18 0.606 0.728 19 0.616 0.736 20 0.624 0.742 21 0.631 0.748 22 0.637 0.753 23 0.643 0.757 Top 24 0.647 0.761 | |||
Core average value = 0.419 M aximum exit value = 0.761 Active fuel length = 150 inches LSCS-UFSAR TABLE 4.4-2a TABLE 4.4-2a REV. 4 - APRIL 1988 AXIAL POWER DISTRIBUTION USED TO GENERATE VOID AND QUALITY DISTRIBUTIONS (TYPICAL) | |||
AXIAL NODE POWER-FACTOR Bottom of core 1 0.54 2 0.83 3 1.02 4 1.17 5 1.26 6 1.33 7 1.37 8 1.39 9 1.40 10 1.39 11 1.38 12 1.34 13 1.29 14 1.21 15 1.10 16 0.99 17 0.89 18 0.79 19 0.71 20 0.64 21 0.58 22 0.52 23 0.46 Top of core 24 0.40 | |||
LSCS-UFSAR TABLE 4.4-3 (SHEET 1 OF 2) TABLE 4.4-3 REV. 13 FLOW QUALITY DISTRIBUTION (TYPICAL) | |||
* Core average value = 0.074 Maximum exit value = 0.281 Active fuel length = 150 inches NODE CORE AVERAGE (AVERAGE NODE VALUE) MAXIMUM CHANNEL (END OF NODE VALUE) | |||
BOTTOM 1 0.00 0.00 2 0.000 0.001 3 0.000 0.006 4 0.002 0.017 5 0.006 0.032 6 0.013 0.049 7 0.022 0.067 8 0.032 0.085 9 0.044 0.103 10 0.053 0.121 11 0.063 0.139 12 0.073 0.157 13 0.083 0.173 14 0.093 0.189 15 0.101 0.203 16 0.109 0.216 17 0.117 0.228 18 0.123 0.238 19 0.129 0.248 20 0.134 0.256 21 0.138 0.263 | |||
LSCS-UFSAR TABLE 4.4-3 (SHEET 2 OF 2) | |||
TABLE 4.4-3 REV. 16, APRIL 2006 NODE CORE AVERAGE (AVERAGE NODE VALUE) MAXIMUM CHANNEL (END OF NODE VALUE) 22 0.142 0.270 23 0.146 0.276 TOP 24 0.150 0.281 | |||
* These flow quality distribution valu es are typical for the initial core. | |||
The GE9 and GE14 fuel has an active fuel length of 150 inches. The | |||
ATRIUM-9B and ATRIUM-10 fuel have an active fuel length of 149.0 inches. This design characteristic difference in combination with changes in power distribution and reactor core state produce different flow quality distributions. These differences are included in transient and core design methodology. | |||
LSCS-UFSAR TABLE 4.4-4 TABLE 4.4-4 REV. 4 - APRIL 1988 CORE FLOW DISTRIBUTION (TYPICAL) | |||
OFFICE ZONEDESCRIPTION CENTRAL HOT CENTRAL AVERAGE PERIPHERAL HOT PERIPHERAL AVERAGE Relative Assembly Power 1.4 1.04 0.95 0.70 Relative Assembly Flow 0.93 1.06 0.55 0.57 | |||
LSCS-UFSAR TABLE 4.4-5 TABLE 4.4-5 REV. 0 - APRIL 1984 TYPICAL RANGE OF TEST DATA MEASURED PARAMETER TEST CONDITIONS ADIABATIC TESTS Spacer single-phase loss coefficient N Re* = 0.5 x 10 5 to 3.5 x 10 5 Lower tie plate + orifice | |||
single-phase loss | |||
coefficient T = 100 to 500°F Upper tie plate single-phase friction factor Spacer two-phase loss | |||
coefficient P = 800 to 1400 psia Two-phase friction | |||
multiplier G = 0.5 x 10 6 to 1.5 x 10 6 lb/h-ft 2 X - 0 to 40% | |||
DIABATIC TESTS Heated bundle pressure drop P = 800 to 1400 psia G = 0.5 x 10 6 to 1.5 x 10 6 lb/h-ft 2 ___________________ | |||
* Reynolds Number | |||
LSCS-UFSAR TABLE 4.4-6 TABLE 4.4-6 REV. 4 - APRIL 1988 | |||
THIS PAGE LEFT INTENTIONALLY BLANK. | |||
LSCS-UFSAR TABLE 4.4-7 TABLE 4.4-7 REV. 0 - APRIL 1984 REACTOR COOLANT SYSTEM GEOMETRICAL DATA FLOW PATH LENGTH (in.) HEIGHT AND LIQUID LEVEL (in.) ELEVATION OF BOTTOM OF EACH VOLUME* (in.) MINIMU M FLOW AREAS (ft 2) A. Lower Plenum 216 216 216 -172.5 71.5 B. Core 164 164 164 44 142.0 C. Upper Plenum and Separators 178 178 208 49.5 D. Dome (Above Normal Water Level) 312 312 386.0 343.5 E. Downcomer Area 321 321 321 -51.0 79.5 F. Recirculation Loops and Jet Pumps (one loop) 108.5 ft (one loop) 403 -394.5 132.5 in 2 | |||
*Reference point is recirculat ion nozzle outlet centerline. | |||
LSCS-UFSAR TABLE 4.4-8 TABLE 4.4-8 REV. 0 - APRIL 1984 LENGTHS AND SIZES OF SAFETY INJECTION LINES LINE OD (inches) | |||
LINE LENGTH (feet | |||
)I. HPCS Line A. Pump discharge to valve 16 146.0 B. Inside containment to RPV 12 101.5 Total 247.5 II. LPCI Lines A. Loop A | |||
: 1. Pump discharge to valve* 18/12 182.0 2. Inside containment to RPV 12 101.5 Total 283.5 B. Loop B | |||
: 1. Pump discharge to valve* 18/12 388.5 | |||
: 2. Inside containment to RPV 12 84.5 Total 473.0 C. Loop C | |||
: 1. Pump discharge to valve* 18 344.0 2. Inside containment to RPV 12 77.0 Total 421 0 III. LPCS Line A. Pump discharge to valve* 16 282.5 B. Inside containment to RPV 12 84.5 Total 367.0 | |||
___________________ | |||
* Valve located as near as possible to outside of containment wall. | |||
LSCS-UFSAR TABLE 4.4-9 TABLE 4.4-9 REV. 0 - APRIL 1984 BYPASS FLOW PATHS FLOW PATH DESCRIPTION DRIVING PRESSURE NUMBER OF PATHS 1a. Between Fuel Support and the Control Rod Guide Tube (Upper Path) Core Plate Differential One/Control Rod 1b. Between Fuel Support and the Control Rod Guide Tube (Lower Path) Core Plate Differential One/Control Rod 2. Between Core Plate and the Control Rod Guide Tube Core Plate Differential One/Control Rod | |||
: 3. Between Core Support and the Incore Support Instrument Guide Tube Core Plate Differential One/Instrument 4. Between Core Plate and Shroud Core Plate Differential One | |||
: 5. Between Control Rod Guide Tube and Control Rod Drive Housing Core Plate Differential One/Control Rod 6. Between Fuel Support and Lower Tie-Plate Channel Wall Differential Plus Lower Tie-Plate Differential One/Channel 7. Control Rod Drive Coolant Independent of of Core One/Control Rod | |||
: 8. Between Fuel Channel and Lower Tie-Plate Channel Wall Differential One/Channel 9. Holes in Lower Tie-Plate Lower Tie-Plate/ Bypass Region Differential Two/Assembly | |||
LSCS-UFSAR 4.5-1 REV. 13 4.5 REACTOR MATERIALS | |||
====4.5.1 Control==== | |||
Rod System Structural Materials | |||
4.5.1.1 Material Specifications The following material listing applies to the control rod drive mechanism supplied for this application. The position in dicator and minor nonstructural items are | |||
omitted. a. Cylinder, Tube and Flange Assembly Flange ASME SA 182 Grade F304 Plugs ASME SA 182 Grade F304 Cylinder ASTM A269 Grade TP 304 Outer Tube ASTM A269 Grade TP 304 Tube ASTM A351 Grade CF-3 Spacer ASTM A351 Grade CF-3 | |||
: b. Piston Tube Assembly Piston Tube ASTM A479 Grade XM-19 Stud ASTM A276 Type 304 Head ASME SA 182 Grade F304 Ind. Tube ASME SA 312 Type 316 Cap ASME SA 182 Grade F304. | |||
: c. Drive Assembly Coupling Spud Inconel X-750 Index Tube ASTM A479 Grade XM-19 Piston Head Armco17-4 PH Coupling ASME SA 312 Grade TP 304 or ASTM A511 Grade MT 304 Magnet Housing ASME SA 312 Grade TP 304 or ASTM A511 Grade MT 304. | |||
: d. Collet Assembly Collet Piston ASTM A269 Grade TP 304 or ASME SA 312 Grade TP 304 Finger Inconel X-750 Retainer ASTM A260 Grade TP 304 or ASTM A511 Grade MT 304 Guide Cap ASTM A269 Grade TP 304. | |||
LSCS-UFSAR 4.5-2 REV. 13 | |||
: e. Miscellaneous Parts Stop Piston ASTM A276 Type 304 Connector ASTM A276 Type 304 O-Ring Spacer ASME SA 240 Type 304 Nut ASME SA 193 Grade B8 Barrel ASTM A269 Grade TP 304 or ASME SA 312 Grade TP 304 or | |||
ASME SA 240 Type 304 Collet Spring Inconel X-750 Ring Flange ASME SA 182 Grade F304. | |||
The materials listed under ASTM specific ation number are all in the annealed condition (with the exception of the outer tube in the cylinder, tube and flange assembly), and their properties are re adily available. The outer tube is approximately 1/8 hard, and has a te nsile of 90,000/125,000 psi, yield of 50,000/85,000 psi, and minimum elongation of 25%. | |||
The coupling spud, collet fingers and collet spring are fabricated from Inconel X-750 in the annealed or equalized condition, and heat treated to produce a tensile of 165,000 psi minimum, yield of 105,000 ps i minimum and elongation of 20% | |||
minimum. The piston head is Armco 17-4 PH in condition H-1100, with a tensile of 140,000 psi minimum, yield of 115,000 ps i minimum and elongation of 15% | |||
minimum. | |||
These are widely used materials, whose properties are well known. All have been successfully used for the past 10 to 15 years in similar drive mechanisms. The parts are readily accessible for inspec tion, and replaceable if necessary. | |||
4.5.1.2 Special Materials | |||
No cold worked austenitic stainless steel s with a yield strength greater than 90,000 psi are employed in the control rod drive system. Hardenable martensitic stainless steels are not used. Armco 17-4 PH (precipitation hardened stainless steel) is used for the piston head. This material is aged to the H-1100 condition to produce resistance to stress corrosion cracking in the BWR environments. Armco 17-4 PH (H-1100) has been successfully used for the past 10 to 15 years in BWR drive mechanisms. | |||
4.5.1.3 Processes, Inspections and Tests All austenitic stainless steel used in the control rod drive system is solution annealed material with one exception, the outer tube in the cylinder, tube, and | |||
flange assembly (Subsection 4.5.1.1). Proper solution annealing is verified by LSCS-UFSAR 4.5-3 REV. 13 testing per ASTM-A262, "Recommended Prac tices for Detecting Susceptibility to Intergranular Attack in Stainless Steels." Two special processes are employed which subject selected components to temperatures in the sensitization range: | |||
: a. The cylinder (cylinder, tube and flange assembly) and the retainer (collet assembly) are hard surfaced with Colmonoy 6. | |||
: b. The following components are nitrided to provide a wear resistant surface: | |||
: 1. tube (cylinder, tube and flange assembly), | |||
: 2. piston tube (piston tube assembly), 3. index tube (drive line assembly), and | |||
: 4. collet piston and guide cap (collet assembly). | |||
Colmonoy hard surfaced components have performed successfully for the past 10 to 15 years in drive mechanisms. Nitrided components have accumulated 8 years of BWR service. It is normal practice to remove some control rod drives at each refueling outage. At this time, both the Colmonoy hard surfaced parts and nitrided surfaces are accessible for visual examination. In addition, dye penetrant examinations have been performed on nitrided surfaces of the longest service drives. This inspection program is adequate to detect any incipient defects before they could become serious enough to cause operating problems. | |||
4.5.1.4 Control of Delta Ferrite Content | |||
All Type 308 weld metal is purchased to a specification which requires a minimum of 5% delta ferrite. This amount of ferrite is adequate to prevent any microfissuring (hot cracking) in austenitic stainless steel welds. | |||
4.5.1.5 Protection of Materials Duri ng Fabrication, Shipping and Storage | |||
All the control rod drive parts listed previo usly (Subsection 4.5.1.1) are fabricated under a process specification which limits contaminants in cutting, grinding and tapping coolants and lubricants. It also restricts all other processing materials (marking inks, tape, etc.) to those which are completely removable by the applied cleaning process. All contaminants are then required to be removed by the appropriate cleaning process prior to any of the following: | |||
: a. any processing which increases part temperature above 200° F, LSCS-UFSAR 4.5-4 REV. 13 | |||
: b. assembly which results in decre ase of accessibility for cleaning, or c. release of parts for shipment. | |||
The specification for packaging and shipping the control rod drive provides the following. | |||
The drive is rinsed in hot deionized water and dried in preparation for shipment. | |||
The ends of the drive are then covered with a vapor-tight barrier with desiccant. | |||
Packaging is designed to protect the drive and prevent damage to the vapor barrier. The planned storage period considered in the design of the container and packaging is 2 years. This packaging has been qualified and in use for a number of years. Periodic audits have indicated satisfactory protection. | |||
Site or warehouse storage specifications require inside heated storage comparable to level B of ANSI 45.2.2. | |||
====4.5.2 Reactor==== | |||
Internals Materials 4.5.2.1 Material Specifications | |||
Materials used for steam dryer and core structure are as follows: | |||
Plate, Sheet and Strip ASTM A240 Type 304 Bolts ASTM A193 Grade B8 Nuts ASTM A194 Grade 8 Forgings ASTM A182 Grade F304 Bar ASTM A276 Type 304 Bar ASTM A479 Type 304 Pipe ASTM A312 Grade TP 304 Tube ASTM A269, A249, or A213 Grade TP 304 Pipe Fittings ASTM A403 Grade WPW 304 or WP 304 Pipe Fittings (cast) ASTM A351 Grade CF8 | |||
The following materials are employed in other reactor internal structures: | |||
: a. Steam Separator. All materials are Type 304, 304L, or 316L stainless steel Plate, Sheet and Strip ASTM A240, Type 304 LSCS-UFSAR 4.5-5 REV. 14, APRIL 2002 Forgings ASTM A182, Grade F304 Bars ASTM A479 Type 304 Pipe ASTM A312 Grade TP 304 Tube ASTM A269 Grade TP 304 Bolting Material ASTM A193 Grade B8 Nuts ASTM A194 Grade 8 Castings ASTM A351 Grade CF8 | |||
: b. Jet Pump Assemblies. The components in the jet pump assemblies are a riser, inlet, mixer, diffuser, adaptor, and brackets. All these components are fabricated with Type 304 stainless steel to the following specifications: | |||
Castings ASTM A351 Grade CF8 Bars ASTM A276 Type 304 Bolts ASTM A193 Grade B8 or B8M Sheet and Plate ASTM A240 Type 304 Tubing ASTM A269 Grade TP 304 Pipe ASTM A358 Type 304 and ASTM A312 Grade TP304 Weld Coupling ASTM A403 Grade WP304 Forgings ASTM A182 Grade F304 Auxiliary Wedges The frames ar e fabricated from Type 304, 304L, 316, or 316L stainless steel. | |||
The sliding components are fabricated from XM-19 or Alloy X-750. | |||
Slip Joint Clamps The clamp frames are fabricated per ASTM A-182 Grade F XM-19. The sub-components are fabricated per ASTM B-637 UNS N07750 Type 3. Due to damage repaired during L1R08, the following unique features are associated with Unit 1 jet pump 9. | |||
* The damaged "Stelllite-6" hard faced surface on the restrainer bracket pad was removed. | |||
* Two auxiliary wedges are located on the riser restrainer bracket. The frames are fabricated from ASTM A-240 or A-479 Type 304 stainless steel (0.02% | |||
max. | |||
LSCS-UFSAR 4.5-5a REV. 16, APRIL 2006 | |||
* carbon) and the sliding component is fabricated from Alloy X-750 in accordance with ASTM B-637 UNS N077550 Type 3; | |||
* The replacement inlet mixer wedge is fabricated from ASTM A-240 or A-479 Type 304 stainless steel (0.02% max ca rbon). Both of the wedge bearing surfaces are hard faced with "Stellite-21". | |||
* In L2R10, inlet-mixer wedges and mounting hardware fabricated from Alloy X-750 and solution heat treated 300 series austenitic stainless steel (0.02% max. carbon) materials were installed in all of the Unit 2 jet pumps. | |||
* During L1R11, jet pump riser brace clamps were installed on Unit 1 jet pumps 5/6 and 9/10 to mitigate crack indications by structurally replacing the upper and lower riser brace yoke to riser pipe welds designated as RS-8 and RS-9. The clamp components are fabricated from ASME SA-479/ASTM A479, ASME SA-240/ASTM A240, or ASME SA-182/ASTM A182 Type 316 stainless steel. The bolting components are fabricated from ASME SA-479/ASTM A479, or ASME SA-240/ASTM A240 Type XM-19 stainless steel. | |||
The ratchet springs and nuts are fabricated from ASME SB-670/ASTM B-637 Grade UNS N07750, Type 3 Alloy X-750. | |||
LSCS-UFSAR 4.5-6 REV. 15, APRIL 2004 Identification and justification for using materials in the jet pump assemblies which are not included in Appendix I to Section III of ASME B&PV Code are provided as follows: a. The inlet mixer adaptor casting, the wedge casting, bracket casting adjusting screw, and the diffuser collar casting are Type 304 hard surfaced with Stellite 6 for slip fit joints. | |||
: b. The adaptor is a bimetallic component made by welding a Type 304 forged ring to a forged Inconel 600 ring, made to Specification ASTM B166. | |||
: c. The inlet contains a pin, insert, and beam made of Inconel X-750 to Specification ASTM B637 Grade 688 or UNS N07750 Type 3 (beam), and ASTM A370 Grade E 38 and E55 (pin and insert). | |||
: d. The jet pump beam bolt is stainless steel Type 316L. | |||
: e. The jet pump beam keeper, s crews, plate and pins are 304L, XM-19, or X-750. | |||
4.5.2.2 Controls on Welding | |||
All welding of the reactor internals is performed in accordance with the ASME Section IX B&PV Code. Interpass temperature does not exceed 370° F. Processes used are GTAW, SMAW, GMAW, and SAW. All welds except intermittent and tack welds are | |||
examined by liquid penetrant in accordance with ASME Section III. All welding filler material has a minimum of 5% ferrite as determined by the Schaeffler diagram. | |||
4.5.2.3 Nondestructive Examination of Wrought Seamless Tubular Products | |||
Wrought seamless tubular products were supplied in accordance with the applicable ASTM/ASME material specifications. These specifications require a hydrostatic test on each length of tubing. No special NDT was performed on the tubes. | |||
4.5.2.4 Fabrication and Processing of Austenitic Stainless Steel | |||
All materials have been solution heat treated and either water or air quenched. | |||
Where an air cool was used, a sample of each heat and heat treatment lot was tested in accordance with ASTM A262 practice A or E. There was no heating above 800° F after the final heat treatment, except for thermal cutting or welding. | |||
4.5.2.5 Regulatory Guide Conformance Assessment | |||
This information is addressed in Appendix B of the FSAR. | |||
LSCS-UFSAR 4.6-1 REV. 13 4.6 FUNCTIONAL DESIGN OF REACTIVITY CONTROL SYSTEMS | |||
====4.6.1 Information==== | |||
for Control Rod Drive Systems (CRDS) | |||
4.6.1.1 Control Rod Drive System Design 4.6.1.1.1 Design Bases | |||
4.6.1.1.1.1 General Design Bases | |||
4.6.1.1.1.1.1 Safety Design Bases | |||
The control rod drive mechanical system meets the following safety design bases: | |||
: a. Design provides for a sufficiently rapid control rod insertion so that no fuel damage results from any abnormal operating transient. | |||
: b. Design includes positioning devices, each of which individually supports and positions a control rod. | |||
: c. Each positioning device: | |||
: 1. prevents its control rod from initiating withdrawal as a result of a single malfunction; collet piston stuck in upper position or stuck open withdraw valve will allow drive to continue withdrawal if initiating signal already given (Subsection 4.6.2.3); | |||
: 2. is individually operated so that a failure in one positioning device does not affect the operation of any | |||
other positioning device; | |||
: 3. is individually hydraulically energized when rapid control rod insertion (scram) is signaled so that failure of power sources external to the positioning device does not prevent other positioning devices' control rods from being inserted; and 4. is locked to its control rod to prevent undesirable separation. | |||
LSCS-UFSAR 4.6-2 REV. 14, APRIL 2002 4.6.1.1.1.1.2 Power Generation Design Basis The control rod system drive design provid es for positioning the control rods to control power generation in the core. | |||
4.6.1.1.2 Description The control rod drive system (CRDS) controls gross changes in core reactivity by incrementally positioning neutron absorbing control rods within the reactor core in response to manual control signals. It is also required to quickly shut down the reactor (scram) in emergency situations by rapidly inserting withdrawn control rods into the core in response to a manual or automatic signal. The control rod drive system consists of locking piston, control rod drive mechanisms, and the CRD | |||
hydraulic system (including hydraulic control units, interconnecting piping, instrumentation, and electrical controls). | |||
4.6.1.1.2.1 Control Rod Drive Mechanisms | |||
The CRD mechanism (drive) used for positioning the control rod in the reactor core is a double-acting, mechanically latched, hydraulic cylinder using water as its operating fluid. (See Figures 4.6-1, 4.6-2, 4.6-3, and 4. | |||
6-4.) The individual drives are mounted on the bottom head of the reactor pressure vessel. The drives do not interfere with refueling and are operative even when the head is removed from the reactor vessel. The drives are also readily accessible for inspection and servicing. The bottom location makes maximum utilization of the water in the reactor as a neutron shield and gives the least possible neutron exposure to the drive components. Using water from the condensate storage tank as the operating fluid eliminates the need for special hydraulic fluid. Drives are able to utilize simple piston seals whose leakage does not contaminate the reactor water and does cool the drive mechanisms and their seals. | |||
The drives are capable of inserting or withdrawing a control rod at a slow, controlled rate, as well as providing rapid insertion when required. A mechanism on the drive locks the control rod in 6-inch increments of stroke over the length of the core. | |||
A coupling spud at the top end of the drive index tube (piston rod) engages and locks into a mating socket at the base of the control rod. The weight of the control rod is sufficient to engage and lock this coupling. Once locked, the drive and rod form an integral unit that must be manually unl ocked by specific procedures before components can be separated. | |||
The drive holds its control rod in distinct latch positions until the hydraulic system actuates movement to a new position. Withdrawal of each rod is limited by the seating of the rod in its guide tube. Withdrawal to the overtravel limit can be LSCS-UFSAR 4.6-3 REV. 18, APRIL 2010 accomplished only if the rod and drive ar e uncoupled and will re sult in a control room alarm. | |||
The individual rod indicators, grouped in one large core map control panel display, correspond to relative rod locations in the core. For display purposes the control rods are considered in groups of four adjacent rods centered around a common core volume. Each group is monitored by fo ur LPRM strings (Subsection 7.7.6). Rod groups at the periphery of the core may have less than four rods. | |||
A Rod Select Display and a Status Display are located below the core map display. The rod select display is a touch-screen LC D that provides the operational interface used to select and perform the movement of a control rod. The status display is a touchscreen LCD display mounted directly below the core map display. The status display is capable of providing the same indications as the rod select display. The status display also serves as a back-up to the flat panel touchscreen rod select display in the event of component failure. A selected rod is indicated on all three displays. | |||
4.6.1.1.2.2 Drive Components | |||
Figure 4.6-2 illustrates the operating principle of a drive. Figures 4.6-3 and 4.6-4 illustrate the drive in more detail. The main components of the drive and their functions are described in the following paragraphs. | |||
4.6.1.1.2.2.1 Drive Piston The drive piston is mounted at the lower end of the index tube. This tube functions as a piston rod. The drive piston and index tube make up the main moving assembly in the drive. The drive piston operates between positive end stops, with a hydraulic cushion provided at the upper end only. The piston has both inside and outside seal rings and operates in an annular space between an inner cylinder (fixed piston tube) and an outer cylinder (drive cy linder). Because the type of inner seal used is effective in only one direction, the lower sets of seal rings are mounted with one set sealing in each direction. | |||
A pair of nonmetallic bushings prevents metal-to-metal contact between the piston assembly and the inner cylinder surface. The outer piston rings are segmented step-cut seals with expander springs ho lding the segments against the cylinder wall. A pair of split bushings on the outside of the piston prevents piston contact with the cylinder wall. The effective piston area for downtravel, or withdrawal, is approximately 1.2 in 2 vs. 4.1 in 2 for uptravel, or insertion. This difference in driving area tends to balance the control rod weight and assures a higher force for insertion than for withdrawal. | |||
LSCS-UFSAR 4.6-4 REV. 13 4.6.1.1.2.2.2 Index Tube The index tube is a long hollow shaft made of nitrided Type 304 stainless steel. Circumferential locking grooves, spaced every 6 inches along the outer surface, transmit the weight of the control rod to the collet assembly. | |||
4.6.1.1.2.2.3 Collet Assembly | |||
The collet assembly serves as the index tube locking mechanism. It is located in the upper part of the drive unit. This as sembly prevents the index tube from accidentally moving downward. The assembly consists of the collet fingers, a return spring, a guide cap, a collet housing (part of the cylinder, tube, and flange), and the collet piston. LaSalle is the first domestic facility which contains the redesigned collet retainer tube. The collet retainer tube is fabricated from cast American Society for Testing and Materials A 351 CF-3 alloy with Colmonoy hardfacing, and | |||
the index tube and piston tube ar e fabricated from XM-19 alloy. | |||
Locking is accomplished by fingers mounted on the collet piston at the top of the drive cylinder. In the locked or latched position the fingers engage a locking groove in the index tube. | |||
The collet piston is normally held in the latched position by a force of approximately 150 pounds supplied by a spring. Metal piston rings are used to seal the collet | |||
piston from reactor vessel pressure. Th e collet assembly will not unlatch until the collet fingers are unloaded by a short, automatically sequenced, drive-in signal. A | |||
pressure, approximately 180 psi above reactor vessel pressure, must then be applied to the collet piston to overcome spring force, slide the collet up against the conical surface in the guide cap, and spread the fingers out so they do not engage a locking | |||
groove. A guide cap is fixed in the upper end of th e drive assembly. This member provides the unlocking cam surface for the collet fingers and serves as the upper bushing for the index tube. | |||
If reactor water is used during a scram to supplement accumulator pressure, it is drawn through a filter on the guide cap. | |||
4.6.1.1.2.2.4 Piston Tube | |||
The piston tube is an inner cylinder, or column, extending upward inside the drive piston and index tube. The piston tube is fixed to the bottom flange of the drive and remains stationary. Water is brought to the upper side of the drive piston through this tube. A series of orif ices at the top of the tube provides progressive water shutoff to cushion the drive piston at the end of its scram stroke. | |||
LSCS-UFSAR 4.6-5 REV. 13 4.6.1.1.2.2.5 Stop Piston A stationary piston, called the stop piston, is mounted on the upper end of the piston tube. This piston provides the seal between reactor vessel pressure and the space above the drive piston. It also functions as a positive end stop at the upper limit of control rod travel. A stack of spring washers just below the stop piston helps absorb the final mechanical shock at th e end of control rod travel. The piston rings are similar to the drive piston outer rings. A bleed-off passage to the center of the piston tube is located between the two pairs of rings. This arrangement allows seal leakage from the reactor vessel (during a scram) to be bled directly to the discharge line. The lower pair of seals is used only during the cushioning of the drive piston at the upper end of the stroke. | |||
The center tube of the drive mechanism forms a well to contain the position indicator probe. This probe is an aluminum extrusion attached to a cast aluminum housing. Mounted on the extrusion are he rmetically sealed, magnetically operated, position indicator switches. Each switch is sheathed in a braided glass sleeve, and the entire probe assembly is protected by a thin-walled stainless steel tube. The switches are actuated by a ring magnet located at the bottom of the drive piston. | |||
The drive piston, piston tube, and indicator tube are all of nonmagnetic stainless steel, allowing the individual switches to be operated by the magnet as the piston passes. One switch is located at each position corresponding to an index tube groove, thus allowing indication at each la tching point. An additional switch is located at each midpoint between latching points to indicate the intermediate positions during drive motion. Thus, indica tion is provided for each 3 inches of travel. Duplicate switches are provided for the full-in and full-out postions. One additional switch (an overtravel switch) is located at a position below the normal full-out position. Because the limit of do wntravel is normally provided by the control rod itself as it reaches the backseat position, the drive can pass this position and actuate the overtravel switch only if it is uncoupled from its control rod. A convenient means is thus provided to verify that the drive and control rod are coupled after installation of a drive or at any time during plant operation. | |||
4.6.1.1.2.2.6 Flange and Cylinder Assembly | |||
A flange and cylinder assembly is made up of a heavy flange welded to the drive cylinder. A sealing surface on the upper face of this flange forms the seal to the | |||
drive housing flange. The seals contain reactor pressure and the two hydraulic control pressures. Teflon coated, stainless steel rings are used for these seals. The drive flange contains the integral ball, or two-way, check (ball-shuttle) valve. This valve directs either the reactor vessel pressure or the driving pressure, whichever is higher, to the underside of the drive piston. Reactor vessel pressure is admitted to this valve from the annular space between the drive and drive housing through passages in the flange. | |||
LSCS-UFSAR 4.6-6 REV. 13 Water used to operate the collet piston passes between the outer tube and the cylinder tube. The inside of the cylinder tube is honed to provide the surface required for the drive piston seals. | |||
Both the cylinder tube and outer tube are welded to the drive flange. The upper ends of these tubes have a sliding fit to allow for differential expansion. | |||
The upper end of the index tube is threaded to receive a coupling spud. The coupling (Figure 4.6-1) accommodates a small amount of angular misalignment between the drive and the control rod. Six spring fingers allow th e coupling spud to enter the mating socket on the control rod. A plug then enters the spud and prevents uncoupling. | |||
4.6.1.1.2.2.7 Lock Plug | |||
Two means of uncoupling are provided. With the reactor vessel head removed, the lock plug can be raised against the spring force of approximately 50 pounds by a rod | |||
extending up through the center of the control rod to an unlocking handle located above the control rod velocity limiter. The control rod, with the lock plug raised, can then be lifted from the drive. | |||
The lock plug can also be pushed up from below, if it is desired to uncouple a drive without removing the reactor pressure vessel head for access. In this case, the central portion of the drive mechanism is pushed up against the uncoupling rod assembly, which raises the lock plug and allows the coup ling spud to disengage the socket as the drive piston and index tube are driven down. | |||
The control rod is heavy enough to force the spud fingers to enter the socket and push the lock plug up, allowing the spud to enter the socket completely and the plug to snap back into place. Therefore, the drive can be coupled to the control rod using only the weight of the control rod. However, with the lock plug in place, a force in excess of 50,000 pounds is required to pull the coupling apart. | |||
4.6.1.1.2.3 Materials of Construction Factors that determine the choice of construction materials are discussed in the following subsections. | |||
4.6.1.1.2.3.1 Index Tube | |||
The index tube must withstand the locking and unlocking action of the collet fingers. A compatible bearing combinat ion must be provided that is able to withstand moderate misalignment forces. | |||
The reactor environment limits the choice of materials suitable for corrosion resistance. The column and tensile loads LSCS-UFSAR 4.6-7 REV. 13 can be satisfied by an anne aled AISI-300 series stainless steel. The wear and bearing requirements are provided by Malc omizing the complete tube. To obtain suitable corrosion resistance, a carefully controlled process of surface preparation is employed. | |||
4.6.1.1.2.3.2 Coupling Spud The coupling spud is made of Inconel-750 that is aged for maximum physical strength and the required corrosion resistance. Because misalignment tends to cause chafing in the semispherical contact area, the part is protected by a thin chromium plating (Electrolized). This plating also prevents galling of the threads attaching the coupling spud to the index tube. | |||
4.6.1.1.2.3.3 Collet Fingers | |||
Inconel-750 is used for the collet fingers, which must function as leaf springs when cammed open to the unlocked position. Colmonoy 6 hard facing provides a long wearing surface, adequate for design life, to the area contacting the index tube and unlocking cam surface of the guide cap. | |||
4.6.1.1.2.3.4 Seals and Bushings | |||
Graphitar 14 is selected for seals and bushings on the drive piston and stop piston. The material is inert and has a low friction coefficient when water lubricated. | |||
Because some loss of Graphitar strength is experienced at higher temperatures, the drive is supplied with cooling wate r to hold temperatures below 250 | |||
° F. The Graphitar is relatively soft, which is adva ntageous when an occasional particle of foreign matter reaches a seal. The resulting scratches in the seal reduce sealing efficiency until worn smooth, but the drive design can tolerate considerable water leakage past the seals into the reactor vessel. | |||
4.6.1.1.2.3.5 Summary | |||
All drive components exposed to reactor vessel water are made of AISI-300 series stainless steel except the following: | |||
: a. Seals and bushings on the drive piston and stop piston are Graphitar 14. | |||
: b. All springs and members requiring spring action (collet fingers, coupling spud, and spring washer s) are made of Inconel-750. | |||
: c. The ball check valve is a Haynes Stellite cobalt-base alloy. | |||
: d. Elastomeric O-ring seals are ethylene propylene. | |||
LSCS-UFSAR 4.6-8 REV. 13 e. Collet piston rings are Haynes 25 alloy. | |||
: f. Certain wear surfaces are hard-faced with Colmonoy 6. | |||
: g. Nitriding by a proprietar y new Malcomizing process and chromium plating are used in certain areas where resistance to abrasion is necessary. | |||
: h. The drive piston head is made of Armco 17-4PH. | |||
Pressure-containing portions of the drives are designed and fabricated in accordance with requirements of Sect ion III of the ASME Boiler and Pressure Vessel Code. | |||
4.6.1.1.2.4 Control Rod Drive Hydraulic System | |||
The control rod drive hydraulic system (Drawing Nos. M-100 and M-146) controls the pressure and flow to and from the driv es through hydraulic control units (HCU). | |||
The water discharged from the drives during a scram flows through the HCU's to the scram discharge volume. The water discharged from a drive during a normal control rod positioning operation flows through the HCU into the exhaust header, a reverse flow then occurs from the exhaust header through the insert/exhaust directional solenoid valves (121) into the latched CRD's. There are as many HCU's as the number of control rod drives. | |||
4.6.1.1.2.4.1 Hydraulic Requirements The CRD hydraulic system design is sh own in Drawing Nos. M-100 and M-146 and Figures 4.6-5 and 4.6-6. The hydraulic requirements, identified by the function they perform, are as follows: | |||
: a. An accumulator hydraulic charging pressure of approximately 1400 to 1500 psig is required. Fl ow to the accumulators is required only during scram reset or system startup. | |||
: b. Drive pressure of approximately 250 psi above reactor vessel pressure is required. A flow rate of approximately 4 gpm to insert a control rod and 2 gpm to withdraw a control rod is required. | |||
: c. Cooling water to the drives is required at approximately 15 psi above reactor vessel pressure and at a flow rate of 0.20 to 0.34 gpm per drive unit. (Cooling water can be interrupted for short periods without damaging the drive.) | |||
LSCS-UFSAR 4.6-9 REV. 14, APRIL 2002 d. The scram discharge volume is sized to receive and contain all the water discharged by the drives during a scram; a minimum volume of 3.34 gallons per drive is required. | |||
: e. The CRD System provides approximately 0.05 gpm to the condensing chambers reference legs for the narrow range, wide range, and fuel zone reactor vessel level instrumentation (UFSAR Section 7.7.1.2.2). | |||
4.6.1.1.2.4.2 System Description The CRD hydraulic systems provide the required functions with the pumps, filter, valves, instrumentation, and piping shown in Drawing Nos. M-100 and M-146 and described in the following paragraphs. | |||
Duplicate components are included, where necessary, to ensure continuous system operation if an inservice component requires maintenance. | |||
The control rod drive hydraulic system also supplies a purge flow to the reactor water cleanup pumps to prevent settling of sediment in the base of each of the two pumps. This flow is taken from the char ging water header and becomes part of the RWCU process fluid once it enters the pump. It is not returned to the CRD | |||
hydraulic system. (Drawings M-97 an d M-143, Sheet 1, and M-100 and M-146, Sheet 1). This purge flow is not required for operation of the pumps. | |||
4.6.1.1.2.4.2.1 Supply Pump One supply pump pressurizes the system. | |||
The condensate system is the normal source of water from the hotwell reject lin | |||
: e. However, during shutdown conditions, the pump suction is from the condensate st orage tank. One spare pump is provided for standby. A discharge check valve prevents backflow through the nonoperating pump. A portion of the pump discharge fl ow is diverted through a minimum flow bypass line to the condensate st orage tank. This flow is controlled by an orifice and is sufficient to prevent immediate pump damage if the pump discharge is inadvertently closed. An additional recirculation line is provided for the supply pumps. This line provides a means of maintaining the pump manufacturer's recommended minimum flow, during unit ou tage time periods when CRD system flow demand is minimal. Flow in this line is controlled by a severe service manual control valve, which is closed during normal plant operation. This line is used concurrently with the previously mentioned minimum flow bypass line to the condensate storage tank. | |||
Condensate water is processed by two filters in the system. The pump suction filter is a cleanable element type with a 25-micron absolute rating. The drive water filter downstream of the pump is a cleanable element type with a 50-micron absolute rating. A differential pressure indicator and control room alarm monitor the filter element as it collects foreign material. | |||
LSCS-UFSAR 4.6-10 REV. 13 4.6.1.1.2.4.2.2 Accumulator Charging Pressure Accumulator charging pressure is established by the discharge pressure of the system supply pump. During scram the scram inlet (and outlet) valves open and permit the stored energy in the accumulato rs to discharge into the drives. The resulting pressure decrease in the charging water header allows the CRD supply pump to run out (i.e., flow rate to increase substantially) into the control rod drives via the charging water header. The fl ow sensing system upstream of the accumulator charging header detects high flow and closes the flow control valve. This action maintains increased flow through the charging water header. | |||
Pressure in the accumulator charging header is monitored in the control room with a pressure indicator and a low/high pre ssure alarm. An automatic scram is initiated when the charging water header pressure drops below 1157 psig for more than approximately 10 seconds. | |||
The automatic scram on low pressure in the charging water header is not active in the run mode because the accumulators are not required for scram at operating pressures. The automatic scram is also no t active in the shutdown mode since no control rods may be withdrawn in this mode. In all other modes, the automatic scram on low charging-water-header pressure remains active. | |||
During normal operation the flow control valve maintains a constant system flow rate. This flow is used for drive flow, drive cooling, and system stability. | |||
4.6.1.1.2.4.2.3 Drive Water Pressure Drive water pressure required in the dr ive header is maintained by the drive pressure control valve, which is manually adjusted from the control room. A flow rate of approximately 6 gpm (the sum of the flow rate required to insert and withdraw a control rod) normally passe s from the drive water pressure stage through two solenoid-operated stabilizing valves (arranged in parallel) and then goes into the cooling water line. The fl ow through one stabilizing valve equals the drive insert flow; that of the other stabilizing valve equals the drive withdrawal flow. When operating a drive, the required flow is diverted to that drive by closing the appropriate stabilizing valve. Thus, flow through the drive pressure control valve is always constant. | |||
Flow indicators in the drive water head er and in the line downstream from the stabilizing valves allow the flow rate thro ugh the stabilizing valves to be adjusted when necessary. Diff erential pressure between the reactor vessel and the drive pressure stage is indicated in the control room. | |||
LSCS-UFSAR 4.6-11 REV. 13 4.6.1.1.2.4.2.4 Cooling Water Header The cooling water header is located downstream from the drive pressure control valve. When not moving a CRD, all system flow returns to vessel through the cooling water header. | |||
The flow through the flow control valve is virtually constant. Therefore, once adjusted, the drive pressure control va lve maintains the required pressure independent of reactor pressure. Changes in setting of the pressure control valves are required only to adjust for changes in the cooling requirements of the drives, as their seal characteristics change with time. A flow indicator in the control room monitors cooling water flow. A differential pressure indicator in the control room indicates the difference between reactor vessel pressure and drive cooling water pressure. Although the drives can function without cooling water, seal life is shortened by long term exposure to reactor temperatures. The temperature of each drive is recorded in the control room, an d excessive temperatures are annunciated. | |||
4.6.1.1.2.4.2.5 Return Line | |||
The H 2O discharged from the HCU during a normal control rod positioning operation is discharged back to the RPV through the insert/exhaust directional solenoid valves of adjoining HCUs. | |||
4.6.1.1.2.4.2.6 Scram Discharge Volume | |||
The scram discharge volume consists of header piping which connects to each HCU and drains into an instrument volume. The header piping is sized to receive and | |||
contain all the water discharged by the drives during a scram, independent of the instrument volume. Each header pipe is designed with a hydrolazing port having 3/4" threaded plugs to allow the lines to be flushed occasionally, to prevent radiation build-up. During normal plant operation the scram discharge volume is empty and vented to atmosphere through its open vent and drain valves. When a scram occurs, upon a signal from the safety circuit, these vent and drain valves are closed to conserve reactor water. Lights in the control room indicate the position of these | |||
valves. During a scram, the scram di scharge volume partly fills with water discharged from above the drive pistons. While scrammed, the control rod drive seal leakage from the reactor continues to flow into the scram discharge volume until the discharge volume pressure equals the reactor vessel pressure. A check valve in each HCU prevents reverse flow from the scram discharge header volume to the drive. When the initial scram signal is cleared from the reactor pr otection system, the scram discharge volume signal is overridden with a keylock override switch, and the scram discharge volume is drained and returned to atmospheric pressure. | |||
LSCS-UFSAR 4.6-12 REV. 15, APRIL 2004 Remote manual switches in the pilot valve solenoid circuits allow the discharge volume vent and drain valves to be tested without disturbing the reactor protection system. Closing the scram discharge volume valves allows the outlet scram valve seats to be leak tested by timing the accumulation of leakage inside the scram discharge volume. | |||
There are two instrument volumes associat ed with the scram discharge volume. Four level switches and two analog trip systems connected to each instrument volume to monitor the volume for abnormal water level. Each analog trip system consists of a transmitter and a trip unit. The level switches are set at three different levels. At the lowe st level, a level switch actuates to indicate that the volume is not completely empty during post scram draining or to indicate that the volume starts to fill through leakage accumulation at other times during reactor operation. At the second level, one leve l switch produces rod withdrawal block to prevent further withdrawal of any cont rol rod when leakage accumulates to approximately half the capacity of the instrument volume. The remaining two level switches and the trip units are interconnected with the reactor protection system (RPS) trip channels and will initiate a reactor scram should water accumulation fill the instrument volume. The liquid level switches are float type and transmitters are differential pressure type. Each di fferential pressure transmitter/trip unit combinations are powered from separate ESS Division sources that are independent of the Reactor Protection system power supply. | |||
Redundant Vent & Drain Valves, placed in series, are located in the vent and drain piping for the scram discharge volume. | |||
This system configuration addresses the co ncerns identified in IE Bulletin No. 80- | |||
: 17. 4.6.1.1.2.4.3 Hydraulic Control Units | |||
Each hydraulic control unit (HCU) furnishes pressurized water on signal to a drive unit. The drive then positions its cont rol rod as required. Operation of the electrical system that supplies scram and normal control rod positioning signals to the HCU is described in Subsection 7.7.2. | |||
Operation of the electrical system which supplies ATWS signals to the HCU is described in Subsection 7.6.5. | |||
The basic components in each HCU are: manual, pneumatic, and electrical valves; an accumulator; related piping; electrical connections; filters; and instrumentation (Drawing Nos. M-100 and M-146 and Figure 4.6-7). | |||
The components and their functions are described in the following paragraphs. | |||
4.6.1.1.2.4.3.1 Insert Drive Valve | |||
The insert drive valve is solenoid-operated and opens on an insert signal. The valve supplies drive water to the bottom side of the main drive piston. | |||
LSCS-UFSAR 4.6-13 REV. 18, APRIL 2010 4.6.1.1.2.4.3.2 Insert Exhaust Valve | |||
The insert exhaust valve also opens by so lenoid on an insert signal. The valve discharges water from above the drive piston to the exhaust water header. | |||
4.6.1.1.2.4.3.3 Wi thdraw Drive Valve | |||
The withdraw drive valve is solenoid-opera ted and opens on a withdraw signal. The valve supplies drive water to the top of the drive piston. | |||
4.6.1.1.2.4.3.4 Withdraw Exhaust Valve The solenoid-operated withdraw exhaust valve opens on a withdraw signal and discharges water from below the main drive piston to the exhaust header. It also serves as the settle valve. The valve op ens following any normal drive movement (insert or withdraw) to allow the control rod and its drive to settle back into the | |||
nearest latch position. | |||
4.6.1.1.2.4.3.5 Speed Control Valves The speed control valves regulate the control rod insertion and withdrawal rates during normal operation. They are manually adjustable flow control valves used to regulate the water flow to and from the volume beneath the main drive piston. A correctly adjusted valve does not require readjustment except to compensate for changes in drive seal leakage. | |||
4.6.1.1.2.4.3.6 Scram Pilot Valves | |||
The scram pilot valves are operated from the reactor protection system trip system. Either a single scram pilot valve with dual solenoid operated pilot assemblies or two single scram pilot valve assemblies control both the scram inlet valve and the scram exhaust valve. The scram pilot valve (either with dual solenoid operated pilot assemblies or with a single pilot solenoid assembly) are solenoid-operated, normally energized valves. On loss of electrical sign al to the scram pilot valve solenoids, such as the loss of external a-c power, the inlet port(s) close and the exhaust port(s) open on both scram pilot valve solenoids. Th e scram pilot valves (Drawing M-100 and M-146) are arranged so that the trip system signal must be removed from both scram pilot valve solenoids before air pressure can be discharged from the scram valve operators. This prevents the inadvertent scram of a single drive in the event of a failure of one of the scram pilot valve solenoids. | |||
LSCS-UFSAR 4.6-13a REV. 18, APRIL 2010 4.6.1.1.2.4.3.7 Scram Inlet Valve | |||
The scram inlet valve opens to supply pressurized water to the bottom of the drive piston. This quick opening globe valve is op erated by an internal spring and system pressure. It is closed by air pressure applied to the top of its diaphragm operator. | |||
LSCS-UFSAR 4.6-14 REV. 18, APRIL 2010 A position indicator switch on this valve pr ovides indication in the control room as soon as the valve starts to open. | |||
As the scram inlet valve and the scram exhaust valve start to open, position indication switches on the valves initiate "valve open" indication in the main control room. 4.6.1.1.2.4.3.8 Scram Exhaust Valve | |||
The scram exhaust valve opens slightly before the scram inlet valve, exhausting water from above the drive piston. The exhaust valve opens faster than the inlet valve because of a high air pressure spring setting in the valve operator. Otherwise the valves are similar. | |||
4.6.1.1.2.4.3.9 Scram Accumulator | |||
The scram accumulator stores sufficient energy to fully insert a control rod at lower vessel pressures. At higher vessel pressures the accumulator pressure is assisted or supplanted by reactor vessel pressure. The accumulator is a hydraulic cylinder with a free-floating piston. The piston separates the water on top from the nitrogen below. A check valve in the accumulator charging line prevents loss of water pressure in the event supply pressure is lost. | |||
During normal plant operation, the accumulator piston is seated at the bottom of its cylinder. Loss of nitrogen decreases th e nitrogen pressure, which actuates a pressure switch and sounds an alarm in the control room. To ensure that the accumulator is always able to produce a sc ram, it is continuously monitored for water leakage. A float-type level switch actuates an alarm if water leaks past the piston barrier and collects in the accumulator instrumentation block. | |||
4.6.1.1.2.4.3.10 Alternate Rod Insertion Scram Valves | |||
The alternate rod insertion (ARI) scram valves are redundant to the existing RPS scram backup valves C11-F110A&B, and scram discharge volume vent and drain pilot valves C11-F379 & F387. The ARI valves provide an alternate means of initiating control rod insertion during an ATWS event. The ARI valves have direct current solenoid dual coil operators. The valves are provided with position switches to indicate valve open/closed status in the main control room. The valves perform three functions during an ATWS trip: | |||
: 1. Block the instrument air supply line to the pilot scram valves. | |||
: 2. Exhaust the air from the pilot scram air header to 5 psig in 15 seconds. | |||
LSCS-UFSAR 4.6-15 REV. 13 3. Exhaust air header to the scram discharge volume vent and drain valves, permitting these valves to close. | |||
4.6.1.1.2.5 Control Rod Drive System Operation The control rod drive system performs rod insertion, rod withdrawal, and scram. | |||
These operational functions are described as follows. | |||
4.6.1.1.2.5.1 Rod Insertion Rod insertion is initiated by a signal from the operator to the insert valve solenoids. This signal causes both insert valves to open. The insert drive valve applies reactor pressure plus approximately 90 psi to the bottom of the drive piston. The insert exhaust valve allows water from above the drive piston to discharge to the exhaust header. | |||
As is illustrated in Figure 4.6-3, the lock ing mechanism is a ratchet-type device and does not interfere with rod insertion. The speed at which the drive moves is determined by the flow through the insert speed control valve, which is set for approximately 4 gpm for a shim speed (non scram operation) of 3 in/sec. During normal insertion, the pressure on the downstream side of the speed control valve is 90 to 100 psi above reactor vessel pressure. However, if the drive slows for any reason, the flow through and pressure drop across the insert speed control valve will decrease; the full differential pressure (260 psi) will th en be available to cause continued insertion. With 260-psi differential pressure acting on the drive piston, the piston exerts an upward force of 1040 pounds. | |||
4.6.1.1.2.5.2 Rod Withdrawal Rod withdrawal is, by design, more involved than insertion. The collet finger (latch) must be raised to reach the unlocked position (Figure 4.6-3). The index tube notches and the collet fingers are shaped so that the downward force on the index tube holds the collet fingers in place. The index tube must be lifted before the collet fingers can be released. This is done by opening the drive insert valves (in the manner described in the preceding paragr aph) for approximately 1 second. The withdraw valves are then opened, applying driving pressure above the drive piston and opening the area below the piston to the exhaust header. Pressure is simultaneously applied to the collet piston. As the piston raises, the collet fingers | |||
are cammed outward, away from th e index tube, by the guide cap. | |||
The pressure required to release the latch is set and maintained at a level high enough to overcome the force of the latch return spring plus the force of reactor pressure opposing movement of the collet piston. When this occurs, the index tube is unlatched and free to move in the withdraw direction. Water displaced by the drive piston flows out through the withdraw speed control valve, which is set to give LSCS-UFSAR 4.6-16 REV. 18, APRIL 2010 the control rod a shim speed of 3 in/sec. The maximum control rod drive withdrawal speed is 6.0 in/sec when the Operating Limit MCPR established in the Core Operating Limits Report (COLR) is set greater than or equal to the value corresponding to a RWE - at Power analysis for an "unblocked" condition (References 6 and 7). Otherwise, the maximum control rod drive withdrawal speed is 3.6 in/sec. See subsection 15.4.2.3 fo r additional details. The entire valving sequence is automatically controlled and is initiated by a single operation of the rod withdraw switch. | |||
4.6.1.1.2.5.3 Scram | |||
During a scram the scram pilo t valves and scram valves ar e operated as previously described. With the scram valves open, accumulator pressure is admitted under the drive piston, and the area over the drive piston is vented to the scram discharge volume. | |||
The large differential pressure (initially approximately 1500 psi and always several hundred psi, depending on reactor vessel pressure) produces a large upward force on the index tube and control rod. This force gives the rod a high initial acceleration and provides a large margin of force to overcome any possible friction. After the initial acceleration is achieved, the drive continues at a nearly constant velocity. This characteristic provides a high initial rod insertion rate. As the drive piston nears the top of its stroke, the piston seals close off the large passage (buffer orifices) in the stop piston tube, and the drive slows. | |||
Prior to a scram signal the accumula tor in the hydraulic control unit has approximately 1450-1510 psig on the water side, and > | |||
980 and <1200 psig on the nitrogen side. As the inlet scram valve opens, the full water side pressure is | |||
available at the control rod drive acting on a 4.1 in 2 area. As CRD motion begins, this pressure drops to the gas-side pressure less line losses between the accumulator and the CRD. At low vessel pressures, the accumulator completely discharges with a resulting gas-side pres sure of approximately 575 psig. Reactor pressure provides the force necessary to scram the reactor when reactor pressure exceeds scram accumulator pressure. | |||
The control-rod-drive accumulators are required to scram the control rod when the reactor pressure is low. When the reactor pressure is low, the accumulator retains sufficient stored energy to ensure the complete insertion of the control rod in the required time. The accumulator is not required in order to scram the control rod in time when the reactor is close to or at fu ll operating pressure. In this instance, the reactor pressure alone will scram the control rod in the required time. However, the accumulator does provide an additional energy boost to the reactor pressure in providing scram action at vessel pressu res less than accumulator pressures. | |||
LSCS-UFSAR 4.6-17 REV. 14, APRIL 2002 The control rod drive system, with accumu lators, was designed to meet the scram time requirements specified in Technical Specification. | |||
4.6.1.1.2.6 Instrumentation | |||
The general functional requirements for the control rod drive are discussed in Subsection 4.6.1.1.2.4.1. | |||
4.6.1.2 Control Rod Drive Housing Supports | |||
4.6.1.2.1 Safety Objective | |||
The control rod drive (CRD) housing su pports prevent any significant nuclear transient in the event a drive housing brea ks or separates from the bottom of the reactor vessel. | |||
4.6.1.2.2 Safety Design Bases The CRD housing supports meet the following safety design bases: | |||
: a. Following a postulated CRD housing failure, control rod downward motion is limited so that any resulting nuclear transient cannot be sufficient to cause fuel damage. | |||
: b. The clearance between the CRD housings and the supports is sufficient to prevent vertical cont act stresses caused by thermal expansion during plant operation. | |||
4.6.1.2.3 Description The CRD housing supports are shown in Fi gure 4.6-8. Horizontal beams are installed immediately below the bottom of the reactor vessel, between the rows of CRD housings. The beams are supported by brackets welded to the steel form liner of the drive room in the reactor support pedestal. | |||
Hanger rods, approximately 10-feet long and 1-3/4-inches in diameter, are supported from the beams on stacks of disc springs. These springs compress approximately 2 inches under the design load. | |||
LSCS-UFSAR 4.6-18 REV. 13 The support bars are bolted between the bottom ends of the hanger rods. The spring pivots at the top, and the beveled, loose-fitting ends on the support bars prevent substantial bending moment in th e hanger rods if the support bars are overloaded. | |||
Individual grids rest on the support bars between adjacent beams. Because a single-piece grid would be difficult to handle in the limited work space and because it is necessary that control rod drives, position indicators, and incore instrumentation components be accessible for inspection and maintenance, each grid is designed for inplace assembly or disassembly. Each grid assembly is made from two grid plates, a clamp, and a bolt. The top part of the clamp guides the grid to its correct position directly below the respective CRD housing that it would support in the postulated accident. | |||
When the support bars and grids are installe d, a gap of approximately 1-1/2 inch at room temperature is provided between th e grid and the bottom contact surface of the control rod drive flange. During system heatup, this gap is reduced by a net downward expansion of the housings with respect to the supports. In the hot operating condition, the gap is reduced approximately 1/4 inch. | |||
In the postulated CRD housing failure, the CRD housing supports are loaded when the lower contact surface of the CRD flange contacts the grid. The resulting load is then carried by two grid plates, two support bars, four hanger rods, their disc springs, and two adjacent beams. | |||
The American Institute of Steel Construction (AISC) Manual of Steel Construction , "Specification for the Design, Fabricatio n and Erection of Structural Steel for Buildings," was used in designing the CRD housing support system. However, to | |||
provide a structure that absorbs as much energy as practical without yielding, the allowable shear, tension and bending stresses used 1.5 times the AISC allowable stresses. | |||
For purposes of mechanical design, the postulated failure resulting in the highest forces is an instantaneous circumferential separation of the CRD housing from the reactor vessel, with an internal pressu re of 1086 psig (react or vessel operating pressure) acting on the area of the separated housing. The weight of the separated housing, control rod drive, and blade, plus the pressure of 1086 psig acting on the area of the separated housing, gives a fo rce of approximately 32,000 pounds. This force is multiplied by an impact factor that conservatively assumes the housing travels through a 1-1/2 inch gap before it contacts the supports. The total force of approximately 120,000 pounds is then treated as a static load in design. | |||
LSCS-UFSAR 4.6-19 REV. 14, APRIL 2002 All CRD housing support subassemblies are fabricated of commonly available structural steel, except for the following items: | |||
Material a. grid bars ASTM-A-441, b. disc springs Schnorr, Type BS-125-71-8, and c. hex bolts and nuts ASTM-A-307. | |||
====4.6.2 Evaluations==== | |||
of the CRDS | |||
4.6.2.1 Failure Mode and Effects Analysis Engineering standards for electrical and physical separation, a design with high safety factors, and the unitary design approach for the CRD modules using ASME standards have each contributed toward an effective and proven CRDS for the control and safe shutdown of BWR's designed by GE. An analysis of failure modes and effects has not been completed for the LSCS units because the CRDS design has a proven history beginning with Dresde n-1. Further analytical evaluations are believed to be of less value than the accrual of real operating data and the incorporation of generic improvements based on actual experience. LSCS utilized this approach in lieu of FMEA. | |||
4.6.2.2 Protection from Common Mode Failures | |||
Based on NEDO-10189, NEDO-10349, and NEDO-20626, General Electric concludes that the complete failure of the BWR control rod scram system due to common mode failure is of such extremely low probability that no change in BWR design to account for the event is warranted. | |||
EGC does not believe the ATWS to be a credible event; nevertheless, the LSCS design includes three provisions to assist shutdown in this unlikely event: tripping of the recirculation pumps, scram discharge volume upgrades, and the addition of alternate rod insertion (ARI) and main steam isolation valve closure modifications. These modifications adequately prevent and, additionally, contribute to the | |||
mitigation of ATWS events. | |||
4.6.2.3 Safety Evaluation | |||
4.6.2.3.1 Control Rod Drives 4.6.2.3.1.1 Evaluation of Scram Time | |||
The rod scram function of the control rod drive system provides the negative reactivity insertion required by safety design basis in Subsec tion 4.6.1.1.1.1.1, LSCS-UFSAR 4.6-20 REV. 1 Item c, part 1. The scram time shown in the description is adequate as shown by the transient analyses of Chapter 15.0. | |||
4.6.2.3.1.2 Analysis of Malfunction Relating to Rod Withdrawal | |||
There are no known single malfunctions that cause the unplanned withdrawal of even a single control rod; providing initiating signal has not been given (Subsections 4.6.1.1.1.1.1, Item c, part 1, and 4.6.2.3.1.2.10). However, if multiple malfunctions are postulated, studies show that an unplanned rod withdrawal can occur at withdrawal speeds that vary with the combination of malfunctions postulated. In all cases the subsequent withdrawal speeds are less than that assumed in the rod drop accident analysis as discussed in Chapter 15.0. Therefore, the physical and radiological consequences of such rod withdrawals are less than those analyzed in the rod drop accident. | |||
4.6.2.3.1.2.1 Drive Housin g Fails at Attachment Weld | |||
The bottom head of the reactor vessel has a penetration for each control rod drive location. A drive housing is raised into position inside each penetration and fastened by welding. The drive is raised into the drive housing and bolted to a flange at the bottom of the housing. Th e housing material is seamless, Type 304 stainless steel pipe with a minimum tens ile strength of 75,000 psi. The basic failure considered here is a complete circumferential crack through the housing wall | |||
at an elevation just below the J-weld. | |||
Static loads on the housing wall include the weight of the drive and the control rod, the weight of the housing below the J-weld, and the reactor pressure acting on the | |||
6-inch diameter cross-sectional area of the housing and the drive. Dynamic loading results from the reaction force during drive operation. | |||
If the housing were to fail as described, the following sequence of events is foreseen. | |||
The housing would separate from the vessel. The control rod, drive, and housing would be blown downward against the support structure by reactor pressure acting on the cross-sectional area of the housing and the drive. The downward motion of the drive and associated parts would be determined by the gap between the bottom of the drive and the support structure and by the deflection of the support structure under load. In the current design, maximum deflection is limited to 3.65 inches. If the collet were to remain latched, no further control rod ejection would occur (Reference 4); the housing would not drop far enough to clear the vessel penetration. | |||
Reactor water would leak at a rate of approximately 220 gpm through the 0.03-inch diametral clearance between the housing and the vessel penetration. | |||
If the basic housing failure were to occur while the control rod is being withdrawn (this is a small fraction of the total drive operating time) and if the collet were to stay unlatched, the following sequence of events is foreseen. The housing would LSCS-UFSAR 4.6-21 REV. 1 separate from the vessel. The drive and housing would be blown downward against the control rod drive housing support. | |||
Calculations indicate that the steady-sta te rod withdrawal velocity would be 0.3 ft/sec. During withdrawal, pressure under the collet piston would be approximately 250 psi greater than the pressure over it. | |||
Therefore, the collet would be held in the unlatched position until driving pressure was removed from the pressure-over port. | |||
4.6.2.3.1.2.2 Rupture of Hydraulic Line(s) to Drive Housing Flange | |||
There are three types of possible rupture of hydraulic lines to the drive housing flange: (1) pressure-under line break; (2) pressure-over line break; and (3) coincident breakage of both of these lines. | |||
4.6.2.3.1.2.2.1 Pressu re-Under Line Break | |||
For the case of a pressure-under line break, a partial or complete circumferential opening is postulated at or near the point where the line enters the housing flange. Failure is more likely to occur after anot her basic failure wherein the drive housing or housing flange separates from the reactor vessel. Failure of the housing, however, does not necessarily lead directly to failure of the hydraulic lines. | |||
If the pressure-under line were to fail and if the collet were latched, no control rod withdrawal would occur. There would be no pressure differential across the collet piston and, therefore, no tendency to unlatch the collet. Consequently, the associated control rod could not be inserted or withdrawn. | |||
The ball check valve is designed to seal off a broken pressure-under line by using reactor pressure to shift the check ball to its upper seat. If the ball check valve were prevented from seating, reactor water would leak to the atmosphere. Because of the broken line, cooling water could not be supplied to the drive involved. Loss of cooling water would cause no immediate dama ge to the drive. However, prolonged exposure of the drive to temperatures at or near reactor temperature could lead to deterioration of material in the seals. High temperature would be indicated to the operator by the thermocouple in the position indicator probe. A second indication would be high cooling water flow. | |||
If the basic line failure were to occur while the control rod is being withdrawn, the hydraulic force would not be sufficient to hold the collet open, and spring force normally would cause the collet to latch and stop rod withdrawal. However, if the collet were to remain open, calculations indicate that the steady-state control rod withdrawal velocity would be 2 ft/sec. | |||
LSCS-UFSAR 4.6-22 REV. 15, APRIL 2004 4.6.2.3.1.2.2.2 Pressure-Over Line Break The case of the pressure-over line breakage considers the complete breakage of the line at or near the point where it enters the housing flange. If the line were to break, pressure over the drive piston would drop from reactor pressure to atmospheric pressure. Any significant re actor pressure (approximately 600 psig or greater) would act on the bottom of the drive piston and fully insert the drive. Insertion would occur regardless of the operational mode at the time of the failure. After full insertion, reactor water would leak past the stop piston seals. This leakage would exhaust to the atmosphere through the broken pressure-over line. | |||
The leakage rate of 1000 psi reactor pressu re is estimated to be 4 gpm nominal but not more than 10 gpm, based on experiment al measurements. If the reactor were hot, drive temperature would increase. This situation would be indicated to the reactor operator by the drift alarm, by th e fully inserted driv e, by a high drive temperature (indicated on a recorder in th e control room), and by operation of the drywell sump pump. | |||
4.6.2.3.1.2.2.3 Simultaneous Breakage of the Pressure-Over and Pressure-Under Lines | |||
For the simultaneous breakage of the pressure-over pressure-under lines, pressures above and below the drive piston would drop to zero, and the ball check valve would close the broken pressure-under line. Reactor water would flow from the annulus outside the drive, through the vessel ports, and to the space below the drive piston. | |||
As in the case of pressure-over line breakage, the drive would then insert at a speed dependent on reactor pressure. Full insertion would occur regardless of the operational mode at the time of failure. Reactor water would leak past the drive seals and out the broken pressure-over line to the atmosphere, as described previously. Drive temperature would incr ease. Indication in the control room would include the drift alarm, the fully-inserted drive, the high drive temperature on a recorder in the control room, and operation of the drywell sump pump. | |||
4.6.2.3.1.2.3 All Drive Flange Bolts Fail in Tension Each control rod drive is bolted to a flange at the bottom of a drive housing. The flange is welded to the drive housing. | |||
The CRD mechanism is bolted to the CRD housing flange by 8 bolts. Each bolt ha s significantly high load carrying capacity compared to the actual load. | |||
If a progressive or simultan eous failure of all bolts were to occur, the drive would separate from the housing. The control rod and the drive would be blown downward against the support structure. Impact velocity and support structure loading would be slightly less than that for drive housing failure because reactor pressure would LSCS-UFSAR 4.6-23 REV. 18, APRIL 2010 act on the drive cross-sectional area only and the housing would remain attached to the reactor vessel. The drive would be is olated from the cooling water supply. Reactor water would flow downward past the velocity limiter piston, through the large drive filter, and into the annular space between the thermal sleeve and the drive. For worst-case leakage calculations, the large filter is assumed to be deformed or swept out of the way so it would offer no significant flow restriction. At a point near the top of the annulus, wher e pressure would have dropped to 350 psi, the water would flash to steam and cause ch oke-flow conditions. Steam would flow down the annulus and out the space between the housing and the drive flanges to the atmosphere. Steam formation would limit the leakage rate to approximately 840 gpm. | |||
If the collet were latched, control rod ejection would be limited to the distance the drive can drop before coming to rest on the support structure. There would be no tendency for the collet to unlatch because pressure below the collet piston would drop to zero. Pressure fo rces, in fact, exert 1435 pounds to hold the collet in the latched position. | |||
If the bolts failed during control rod withdrawal, pressure below the collet piston would drop to zero. The collet, with 1650 pounds return force, wo uld latch and stop rod withdrawal. | |||
4.6.2.3.1.2.4 Weld Joining Flange to Housing Fails in Tension | |||
The failure considered is a crack in or near the weld that joins the flange to the housing. This weld extends through the wall and completely around the housing. The flange material is forged, Type 304 stainless steel, with a minimum tensile strength of 75,000 psi. The housing materi al is seamless, Type 304 stainless steel pipe, with a minimum tensile strength of 75,000 psi. The conventional, full-penetration weld of Type 308 stainless steel has a minimum tensile strength approximately the same as that for the parent metal. The design pressure and temperature are 1250 psig and 575 | |||
° F. Reactor pressure acting on the cross-sectional area of the drive, the weight of the control rod, drive, and flange, and the dynamic reaction force during drive operation result in a maximum tensile stress at the weld of approximately 6000 psi. | |||
If the basic flange-to-housing joint failure occurred, the flange and the attached drive would be blown downward against the support structure. The support structure loading would be slightly less than that for drive housing failure because reactor pressure would act only on the drive cross-sectional area. Lack of differential pressure across the collet piston would cause the collet to remain latched and limit control rod motion to ap proximately 3.65 inches. Downward drive movement would be small and, therefore, most of the drive would remain inside the housing. The pressure-under and pressu re-over lines are flexible enough to withstand the small displacement and rema in attached to the flange. Reactor LSCS-UFSAR 4.6-24 REV. 18, APRIL 2010 water would follow the same leakage path described above for the flange-bolt failure, except that exit to the atmosphere would be through the gap between the lower end of the housing and the top of the flange. Water would flash to steam in the annulus surrounding the drive. Th e leakage rate would be approximately 840 gpm. | |||
If the basic failure were to occur during co ntrol rod withdrawal (a small fraction of the total operating time) and if the colle t were held unlatched, the flange would separate from the housing. The drive an d flange would be blown downward against the support structure. The calculated steady-state rod withdrawal velocity would | |||
be 0.13 ft/sec. Because pressure-under and pressure-over lines remain intact, driving water pressure would continue to the drive, and the normal exhaust line restriction would exist. The pressure below the velocity limiter piston would drop below normal as a result of leakage fr om the gap between the housing and the flange. This differential pressure across the velocity limiter piston would result in a net downward force of approximately 70 po unds. Leakage out of the housing would greatly reduce the pressure in the annulus surrounding the drive. Thus, the net downward force on the drive piston would be less than normal. The overall effect of these events would be to reduce rod withdrawal to approximately one-half of normal | |||
speed. With a 560-psi differential across the collet piston, the collet would remain | |||
unlatched; however, it should relatch as soon as the drive signal is removed. | |||
4.6.2.3.1.2.5 Housing Wall Ruptures | |||
This failure is a vertical split in the drive housing wall just below the bottom head of the reactor vessel. The flow area of the hole is considered equivalent to the annular area between the drive and the thermal sleeve. Thus, flow through this annular area, rather than flow through the hole in the housing, would govern leakage flow. The housing is made of Ty pe 304 stainless steel seamless pipe, with a minimum tensile strength of 75,000 psi. The maximum hoop stress of 11,900 psi results primarily from the reactor design pr essure (1250 psig) acting on the inside of the housing. | |||
If such a rupture were to occur, reactor water would flash to steam and leak | |||
through the hole in the housing to the atmosphere at approximately 1030 gpm. | |||
Choke-flow conditions would exist as described previously for the flange-bolt failure. However, leakage flow would be greater because flow resistance would be less; that is, the leaking water and steam would not have to flow down the length of the housing to reach the atmosphere. A critica l pressure of 350 psi causes the water to flash to steam. | |||
No pressure differential across the collet piston would tend to unlatch the collet; but the drive would insert as a re sult of loss of pr essure in the drive housing causing a pressure drop in the space above the drive piston. | |||
LSCS-UFSAR 4.6-25 REV. 18, APRIL 2010 If this failure occurred during control rod withdrawal, drive withdrawal would stop, but the collet would remain unlatched. The drive would be stopped by a reduction of the net downward force action on the drive line. The net force reduction would occur when the leakage flow of 1030 gpm reduces the pressure in the annulus outside the drive to approximately 540 psig , thereby reducing the pressure acting on top of the drive piston to the same value. A pressure differential of approximately 710 psi would exist across the collet piston and hold the collet | |||
unlatched as long as the operator held the withdraw signal. | |||
4.6.2.3.1.2.6 Flange Plug Blows Out | |||
To connect the vessel ports with the bottom of the ball check valve, a hole of 3/4-inch diameter is drilled in the drive flange. The outer end of this hole is sealed with a plug of 0.812-inch diameter and 0.25-inch thickness. A full-penetration, Type 308 stainless steel weld holds the plug in place. The postulated failure is a full circumferential crack in this weld and subsequent blowout of the plug. | |||
If the weld were to fail, the plug were to blow out, and the collet remained latched, there would be no control rod motion. | |||
There would be no pressure differential across the collet piston acting to unlatch the collet. Reactor water would leak past the velocity limiter piston, down the a nnulus between the drive and the thermal sleeve, through the vessel ports and drilled passage, and out the open plug hole to the atmosphere at approximately 320 gpm. | |||
Leakage calculations assume only liquid flows from the flange. Actually, hot reactor water would flash to steam and choke-flow conditions would exist. Thus, the expected leakage rate would be lower than the calculated value. Drive temper ature would increase and initiate an alarm in the control room. | |||
If this failure were to occur during control rod withdrawal and if the collet were to stay unlatched, calculations indicate that control rod withdrawal speed would be | |||
approximately 0.24 ft/sec. Leakage from the open plug hole in the flange would cause reactor water to flow downward past the velocity limiter piston. A small differential pressure across the piston would result in an insignificant driving force of approximately 10 pounds, tending to increase withdraw velocity. | |||
A pressure differential of 295 psi across the collet piston would hold the collet unlatched as long as the driving signal was maintained. | |||
Flow resistance of the exhaust path from the drive would be normal because the ball check valve would be seated at the lower end of its travel by pressure under the drive piston. | |||
LSCS-UFSAR 4.6-26 REV. 18, APRIL 2010 4.6.2.3.1.2.7 Drive Pressure Control Va lve Closure (Reactor Pressure, 0 psig) | |||
The pressure to move a drive is generated by the pressure drop of practically the full system flow through the drive pressure control valve. This valve is a motor-operated valve with a normally closed, stan dby manually operated valve in parallel. The motor-operated valve is adjusted to a fixed opening, to develop a normal pressure (260 psig in excess of normal reactor pressure) on the upstream side of the motor-operated valve. In the event of mo tor-operated valve failure, this valve can be isolated (upstream and downstream gate valves) and its functi on replaced by the manually operated standby valve. | |||
If the flow through the drive pressure contro l valve were to be stopped, as by a valve closure or flow blockage, the drive pressure would increase to the shutoff pressure of the supply pump. The occurrence of this co ndition during withdrawal of a drive at zero vessel pressure will result in a drive pressure increase from 260 psig to no more than 1700 psig. Calculations indicate that the drive would accelerate from a nominal 3 in/sec to approximately 6 in/sec. A pressure differential of 1670 psi across the collet piston would hold the collet unlatched. Flow would be upward, past the velocity limiter piston, but retarding force would be negligible. Rod movement would stop as soon as the driving signal was removed. | |||
4.6.2.3.1.2.8 Ball Check Valve Fails to Close Passage to Vessel Ports | |||
Should the ball check valve sealing the passage to the vessel ports be dislodged and prevented from reseating following the insert portion of a drive withdrawal sequence, water below the drive piston would return to the reactor through the vessel ports and the annulus between the drive and the housing rather than through the speed control valve. Because the flow resistance of this return path would be lower than normal, the calculated withdrawal speed would be 2 ft/sec. During withdrawal, differential pressure across the collet piston would be approximately 40 psi. Therefore, the collet would tend to latch and would have to stick open before continuous withdrawal at 2 ft/sec, could occur. Water would flow upward past the velocity limiter piston, generating a small retarding force of approximately 120 pounds. | |||
4.6.2.3.1.2.9 Hydraulic Cont rol Unit (HCU) Valve Failures | |||
Various failures of the valves in the HCU can be postulated, but none could produce differential pressures approaching those described in the preceding paragraphs and | |||
none alone could produce a high velocity withdrawal. Leakage through either one or both of the scram valves produces a pressure that tends to insert the control rod rather than to withdraw it. If the pressure in the scram discharge volume should exceed reactor pressure following a scram, a check valve in the line to the scram discharge header prevents this pressure from operating the drive mechanisms. | |||
LSCS-UFSAR 4.6-27 REV. 18, APRIL 2010 4.6.2.3.1.2.10 Collet Fingers Fail to Latch When the drive withdraw signal is remove d, the drive continues to withdraw at a fraction of normal speed. Without some in itiating signal there is no known means for the collet fingers to become unlocked. If the drive withdrawal valve fails to close following a rod withdrawal, it would have the same effect as failure of the collet fingers to latch in the index tube. Beca use the collet fingers remain locked until they are unloaded, accidental opening of the drive withdrawal valve does not unlock them. 4.6.2.3.1.2.11 Withdrawal Speed Control Valve Failure | |||
Normal withdrawal speed is determined by differential pressures in the drive and is set for a nominal value of 3 in/sec. Withdr awal speed is maintain ed by the pressure regulating system and is independent of reactor vessel pressure. Tests have shown that accidental opening of the speed control valve to the full-open position produces a velocity of approximately 6 in/sec. | |||
The control rod drive system prevents rod withdrawal and it has been shown above that only multiple failures in a drive uni t and in its control unit could cause an unplanned rod withdrawal. | |||
4.6.2.3.2 Scram Reliability of CRDS | |||
High scram reliability is the result of a nu mber of features of the CRD system. For example: | |||
: a. Two sources of scram energy are used to insert each control rod when the reactor is operating: accumulator pressure and reactor vessel pressure. | |||
: b. Each drive mechanism has its own scram valves and scram pilot valves. Alternatively each drive mechanism may have a single pilot valve with dual solenoid operated pilot assemblies in place of two scram pilot valves. With either scram pilot valve configuration, only one drive can be affected if a scram valve fails to open. Two pilot solenoid s are provided for each drive. | |||
Both pilot solenoids must be de-e nergized to initiate a scram of that drive mechanism. | |||
: c. The reactor protection system and the HCU's are designed so that the scram signal and mode of operation override all others. | |||
: d. The alternate rod insertion (ARI) system provides an alternate means of exhausting the scram air header and closing the vent LSCS-UFSAR 4.6-27a REV. 18, APRIL 2010 and drain valves of the scram discharge volume, thereby providing an additional reactor scram mechanism which is diverse, redundant and independent of the reactor protection system. | |||
LSCS-UFSAR 4.6-28 REV. 13 | |||
: e. The collet assembly and index tu be are designed so they will not restrain or prevent control rod insertion during scram. | |||
: f. The scram discharge volume is monitored for accumulated water and will scram the reactor before the volume is reduced to a point that could interfere with a scram. | |||
4.6.2.3.2.1 Reliability Analysis | |||
A reliability analysis was performed to de monstrate that the ARI design meets the design failure rate criteria of 10 | |||
-6 failures to actuate per reactor-year (reference 5). | |||
The probability of spurious actuation was shown to be more than a factor of 10 less likely than the probability of failure to ac tuate. The basis for demonstrating the 10 6 criteria was the complete electrical in dependence of the ARI system from the electrical portion of the reactor protection system (RPS) including power supplies. When determining the overall electrical system failure probability (ARI and RPS), the independence results in an overall failure probability well beyond any practical | |||
means of engineering judgement (~10-11 failures to actuate per demand). Note that the mechanical portion of the CRD is unchanged by the ARI modification and now becomes the limiting factor in the overall scram system reliability. Hence, the ARI modification provides a conservative means of demonstrating adequate ATWS prevention for the expected ATWS initiators. | |||
The charging water header pressure is monitored with a low pressure alarm to provide warning to control room operators of an impending reactor scram due to low charging-water-header pressure. | |||
The scram assures that sufficient energy remains in the accumulators to shut down the reactor. | |||
4.6.2.3.2.2 Control Rod Support and Operation | |||
As described previously, each control rod is independently supported and controlled as required by safety design bases. | |||
4.6.2.3.3 Control Rod Drive Housing Supports | |||
4.6.2.3.3.1 Safety Evaluation | |||
Downward travel of the CRD housing and its control rod following the postulated housing failure equals the sum of these distances: (1) the compression of the disc springs under dynamic loading, and (2) the initial gap between the grid and the | |||
bottom contact surface of the CRD flange. If the reactor were cold and pressurized, the downward motion of the control rod would be limited to the spring compression LSCS-UFSAR 4.6-29 REV. 13 (approximately 2 inches) plus a gap of approx imately 1-1/2 inch. If the reactor were hot and pressurized, the gap would be reduced approximately 1/4 inch and the spring compression would be slightly less than in the cold condition. In either case, the control rod movement following a housing failure is substantially limited below one drive notch movement (6 inches). Sudden withdrawal of any control rod through a distance of one drive notch at an y position in the core does not produce a transient sufficient to damage any radioactive material barrier. | |||
The CRD housing supports are in place during power operation and when the nuclear system is pressurized. If a control rod is ejected during shutdown, the reactor remains subcritical because it is designed to remain subcritical with any one control rod fully withdrawn at any time. | |||
At plant operating temperature, a gap of approximately 1-1/4 inch exists between the CRD housing and the supports. At lower temperatures the gap is greater. Because the supports do not contact any of the CRD housing except during the postulated accident condition, vertical contact stresses are prevented. | |||
====4.6.3 Testing==== | |||
and Verification of the CRDS 4.6.3.1 Control Rods | |||
4.6.3.1.1 Testing and Inspection | |||
The tests performed on control rods plus their related surveillance program are covered in Subsection 4.6.3.2. | |||
4.6.3.2 Control Rod Drives | |||
4.6.3.2.1 Testing and Inspection | |||
4.6.3.2.1.1 Development Tests | |||
The development drive (one prototype) te sting to date included more than 5000 scrams and approximately 100,000 latching cycles. One prototype was exposed to simulated operating conditions for 5000 ho urs. These tests demonstrated the following: | |||
: a. The drive easily withstands the forces, pressures, and temperatures imposed. | |||
: b. Wear, abrasion, and corrosion of the nitrided Type 304 stainless parts are negligible. Mechanical performance of the nitrided surface is superior to that of materials used in earlier operating reactors. | |||
LSCS-UFSAR 4.6-30 REV. 13 | |||
: c. The basic scram speed of the drive has a satisfactory margin above minimum plant requirements at any reactor vessel pressure. | |||
: d. Usable seal lifetimes in ex cess of 1000 scram cycles can be expected. | |||
4.6.3.2.1.2 Factory Quality Control Tests | |||
Quality control of welding, heat treatment, dimensional tolerances, material verification, and similar factors is ma intained throughout the manufacturing process to ensure reliable performance of the mechanical reactivity control components. Some of the quality control tests performed on the control rods, control rod drive mechanisms, and hydraulic control units are listed as follows: | |||
: a. Control rod absorber tube tests: | |||
: 1. Material integrity of the tubing and end plug is verified by ultrasonic inspection. | |||
: 2. The boron-10 fraction of the boron content of each lot of boron-carbide is verified. | |||
: 3. Weld integrity of the finished absorber tubes is verified by helium leak-testing. | |||
: b. Control rod drive mechanism tests: | |||
: 1. Pressure welds on the drives are hydrostatically tested in accordance with ASME codes. | |||
: 2. Electrical components are ch ecked for electrical continuity and resistance to ground. | |||
: 3. Drive parts that cannot be visually inspected for dirt are flushed with filtered water at high velocity. No significant foreign material is permitted in effluent water. | |||
: 4. Seals are tested for leakage to demonstrate correct seal operation. | |||
: 5. Each drive is tested for sh im motion, latching, and control rod position indication. | |||
LSCS-UFSAR 4.6-31 REV. 13 6. Each drive is subjected to scram timing tests as required by Technical Specifications to verify correct scram performance. | |||
: c. Hydraulic control unit tests: | |||
: 1. Hydraulic systems are hydrostatically tested in accordance with the applicable code. | |||
: 2. Electrical components and systems are tested for electrical continuity and resistance to ground. | |||
: 3. Correct operation of the accumulator pressure and level switches is verified. | |||
: 4. The unit's ability to perfor m its part of a scram is demonstrated. | |||
: 5. Correct operation and adjustment of the insert and withdrawal valves is demonstrated. | |||
4.6.3.2.1.3 Operational Tests | |||
After installation, all rods and drive mech anisms can be tested through their full stroke for operability. | |||
During normal operation, each time a control rod is withdrawn, the operator can observe the incore monitor indications to verify that the control rod is following the drive mechanism. All control rods that are partially withdrawn from the core can be tested for rod-following by inserting or withdrawing the rod and returning it to its original position, while the operator observes the incore monitor indications. | |||
To make a positive test of control rod to control rod drive coupling integrity, the operator can withdraw a control rod to th e end of its travel and then attempt to withdraw the drive to the overtravel position. Failure of the drive to overtravel demonstrates rod-to-drive coupling integrity. | |||
Hydraulic supply subsystem pressures can be observed from instrumentation in the control room. Scram accumulator pressures can be observed on the nitrogen pressure gauges. | |||
4.6.3.2.1.4 Acceptance Tests | |||
The information in this subsection is bein g maintained for historical purposes only, as it is related to pre-startup testing. | |||
LSCS-UFSAR 4.6-32 REV. 18, APRIL 2010 Criteria for acceptance of the individual control rod drive mechanisms and the associated control and protection systems will be incorporated in specifications and test procedures covering three distinct phases: (1) preinstallation, (2) after installation prior to startup, and (3) during startup testing. | |||
The preinstallation specification will define criteria and acceptable ranges of such characteristics as seal leakage, friction, and scram performance under fixed test conditions which must be met before the component can be shipped. | |||
The after-installation, prestartup tests include normal and scram motion and are primarily intended to verify that piping, valves, electrical components, and instrumentation are properly installed. The test specifications will include criteria and acceptable ranges for drive speed, times settings, scram valve response times, and control pressures. These tests are intended more to document system condition | |||
than as tests of performance. | |||
As fuel is placed in the reactor, the startup test procedure will be followed. The tests in this procedure are intended to determine that the initial operational characteristics meet the limits of the specifications over the range of primary coolant temperatures and pressures from ambient to operating. The detailed specifications and procedures have not as yet been prepared but will follow the general pattern established for such specifications and procedures in BWR's presently under construction and in operation. | |||
4.6.3.2.1.5 Surveillance Tests | |||
The surveillance requirements (SR) for the control rod drive system are recommended as follows: | |||
: a. Sufficient control rods shall be withdrawn, following a refueling outage when core alterations are performed, to demonstrate with the technical specification design margin that the core can | |||
be made subcritical at any time in the subsequent fuel cycle with the strongest operable control rod fully withdrawn and all other operable rods fully inserted. | |||
: b. Each partially or fully withdrawn control rod shall be exercised as defined in the Technical Specifications. When any control rod is immovable as a result of excessive friction or mechanical interference, a determination must be made and appropriate action taken. | |||
The monthly control rod exercise test serves as a periodic check against deterioration of the control rod system and also verifies LSCS-UFSAR 4.6-33 REV. 13 the ability of the control rod drive to scram because if a rod can be moved with drive pressure, it will scram since higher pressure is applied during scram. | |||
The frequency of exercising the control rods under the conditions of three or more control rods valved out of service provides even further assurance of the reliability of the remaining control rods. | |||
: c. The coupling integrity shall be verified for each withdrawn control rod as follows: | |||
: 1. when the rod is first withdrawn, observe any indicated response of the nuclear instrumentation; and | |||
: 2. when the rod is fully withdrawn the first time, observe that the drive will not go to the overtravel position. | |||
Observation of a response from the nuclear instrumentation during an attempt to withdraw a control rod indicates indirectly that the rod and drive are coupled. The overtravel position feature provides a positive check on the coupling integrity, for | |||
only an uncoupled drive can reach the overtravel position. | |||
: d. During operation, accumulator pressure and level at the normal operating value are verified. | |||
Experience with control rod drive systems of the same type indicates that weekly verification of accumulator pressure and level is sufficient to assure operability of the accumulator portion of the control rod drive system. | |||
: e. After each major refueling outage, each operable control rod shall be subjected to scram time tests from the fully withdrawn | |||
position. | |||
Experience indicates that the scram times of the control rods do not significantly change over the time interval between refueling outages. A test of the scram times at each refueling outage is sufficient to identify any significant lengthening of the scram times. Routine accumulator surveillance is performed to authenticate the discharge pressure of the CRD pump and its associated hydraulic accumulator. Accu mulator hydraulic pressure retention above the analysis valu e of 1157 psig is observed after a CRD pump trip to assure scram action via charging-water-LSCS-UFSAR 4.6-34 REV. 14, ARPIL 2002 header pressure supplied from the accumulator. The 1157 psig value for this CRD-accumulator auto scram was selected because it exceeds the analytical point where the control rod maximum insertion times were defined. | |||
4.6.3.3 Control Rod Drive Housing Supports | |||
4.6.3.3.1 Testing and Inspection | |||
CRD housing supports are removed for inspection and maintenance of the control rod drives. The operational condition du ring which CRD housing supports can be removed is controlled by the Technical Specifications. When the support structure is reinstalled, it is inspected for correct assembly with particular attention to maintaining the correct gap between the CRD flange lower contact surface and the grid. 4.6.4 Information for Combined Performance of Reactivity Systems | |||
4.6.4.1 Vulnerability to Common Mode Failures Protection of the CRDS from common mode failures is described in Subsection 4.6.2.2, and in GE's "BWR Scram System Reliability Analysis," dated September 30, 1976 (Proprietary) which was pr ovided to Mr. D. F. Ross (NRC) by Mr. E. A. Hughes (GE) by letter of the same date. The evaluation of the ECCS and SLCS against common mode failures is presented in Section 6.3 and Subsection 9.3.5 respectively. In additi on, no balance-of-plant failure will prevent reactivity shutdown. Therefore, no commo n mode failures need be considered in Chapter 15.0. | |||
4.6.4.2 Accidents Taking Credit for Two or More Reactivity Control Systems | |||
There are no postulated accidents evaluated in Chapter 15.0 that take credit for two or more reactivity control systems preventing or mitigating the accident. | |||
====4.6.5 Evaluation==== | |||
of Combined Performance | |||
As indicated in Subsection 4.6.4.2, credit is not taken for multiple reactivity control systems for any postulated accidents in Chapter 15.0. | |||
LSCS-UFSAR 4.6-35 REV. 18, APRIL 2010 | |||
====4.6.6 References==== | |||
: 1. J. P. Fritz, "Testing of Cruciform Control Rods for BWR/6," NEDO-10565, GE APED, April 1972. | |||
: 2. R. J. Benche, "Visual and Photographic Examination of Dresden 1 High Exposure Control Rod B87," NEDO-10541, April 1972. | |||
: 3. R. G. Stirn et al., "Rod Drop Accident Analysis for Large Boiling Water Reactors," NEDO-10527, General Electric Co., Atomic Power Equipment Department, March 1972. | |||
: 4. J. E. Benecki, "Impact Testin g on Collet Assembly for Control Rod Drive Mechanism 7RD B144A," General Electric Company, Atomic Power Equipment Department, APED-5555, November 1967. | |||
: 5. "Reliability Evaluation Analysis - Unit 2 Alternate Rod Insertion System", COM-0249-R-003, February 1983. | |||
: 6. 51-9121141-000, "Licensing Impa cts of Control Rod Withdrawal Speeds for LaSalle Reactors," | |||
AREVA NP Inc., September 2009. | |||
: 7. SPC Document, "Exxon Nuclear Methodology for Boiling Water Reactors - Neutronic Methods for Design and Analysis," XN-NF-80-19(P)(A), Volume 1 and Supplements 1 and 2, Exxon Nuclear Company, Richland, WA, March, 1983. | |||
111------......1lO 59-----51 on,r II III I--ODOODO 000'DO ODD 0 I I NUMBER OF FUEL ASSEMBLIES 764 NUMBER OF CONTROL RODS 185 NUMBER OF LPRM STRINGS 43 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FI GURE 4.1-1 CORE ARRANGEMENT REV.a-APRIL 1984 111------......1lO 59-----51 on,r II III I--ODOODO 000'DO ODD 0 I I NUMBER OF FUEL ASSEMBLIES 764 NUMBER OF CONTROL RODS 185 NUMBER OF LPRM STRINGS 43 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FI GURE 4.1-1 CORE ARRANGEMENT REV.a-APRIL 1984 | |||
*0 DIM.IDENT. | |||
DIM.INCHES DIM.IDENT. | |||
DIM.INCHES DIM.IDE NT.DIM.INCHES LSCS-UFSAR 1:::1 r-tr=========-! | |||
II, 000000001 100000000 00000000 00000000 I oo6"OOOOI 00000000 ,.. | |||
00000000 00000000 00000000 00000000 I-00000000 J'-,-.------Jiil- | |||
............,.;;;;r I" rt.:e!!!/:'\./W**C.**__.. | |||
r 0(!)@0(i)000 lLACe Wt"IGI 0se@*>e@e I." (i;$@@)@)@@*; | |||
--,-- | |||
I-,-- | |||
I I eeeaeeee I*""1)' | |||
AOc.1,.1 e c8@8@l@@)e r".0".11 O,l,jtl."OOee@l e (!D@)ee,*This value is based on 100 mil channels.Channel thickness can be 80 or 100 mil.**This data is based on GE original equipment control blades.Different control blade design are also utilized.LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-2 CORE CELL GE 8X8R FUEL TYPE REV.13*0 DIM.IDENT. | |||
DIM.INCHES DIM.IDENT. | |||
DIM.INCHES DIM.IDE NT.DIM.INCHES LSCS-UFSAR 1:::1 r-tr=========-! | |||
II, 000000001 100000000 00000000 00000000 I oo6"OOOOI 00000000 ,.. | |||
00000000 00000000 00000000 00000000 I-00000000 J'-,-.------Jiil- | |||
............,.;;;;r I" rt.:e!!!/:'\./W**C.**__.. | |||
r 0(!)@0(i)000 lLACe Wt"IGI 0se@*>e@e I." (i;$@@)@)@@*; | |||
--,-- | |||
I-,-- | |||
I I eeeaeeee I*""1)' | |||
AOc.1,.1 e c8@8@l@@)e r".0".11 O,l,jtl."OOee@l e (!D@)ee,*This value is based on 100 mil channels.Channel thickness can be 80 or 100 mil.**This data is based on GE original equipment control blades.Different control blade design are also utilized.LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-2 CORE CELL GE 8X8R FUEL TYPE REV.13 LSCS-UFSAR ooooBooooT r 00000000 0 00000 000 w PO oua w'0000 000 EO 000IH 00 QoOO 0 000 p 00 000 c 00000000 ooodEL.L£T 0 0 00000000ooooooO-FUEL ROD gg88ggg I@-WATER ROOOOOOOOI-TIE ROO 00000000 I 5 | |||
DIM.I.D.DIM.INCHES DIM.LD.DIM.INCHES***This value is based on 100 mil channels.Channel thickness vary from 80 to 100 mil.This data is based on GE original equipment control blades.Different control blade designs are also utilized.LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYS1S REPORT FIGURE 4.1-2a CORE CELL GE 8X8NB FUEL TYPE REV.13 LSCS-UFSAR ooooBooooT r 00000000 0 00000 000 w PO oua w'0000 000 EO 000IH 00 QoOO 0 000 p 00 000 c 00000000 ooodEL.L£T 0 0 00000000ooooooO-FUEL ROD gg88ggg I@-WATER ROOOOOOOOI-TIE ROO 00000000 I 5 | |||
DIM.I.D.DIM.INCHES DIM.LD.DIM.INCHES***This value is based on 100 mil channels.Channel thickness vary from 80 to 100 mil.This data is based on GE original equipment control blades.Different control blade designs are also utilized.LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYS1S REPORT FIGURE 4.1-2a CORE CELL GE 8X8NB FUEL TYPE REV.13 LSCS-UFSAR ATRlUM-9B DESIGN 4-1 I!p o B I-"I-I)Ii...I I L W'o00000000 I rf'0 0 d bocoocY 000000000\,..;.. I 000000000 I 00 boo Mol 0008 000 1 0 00 000 000 we 000 I 000 we 000 000 000 I 000 000 000000000 I 000000000 I 0000 F 000000 E OO-d I J)I ,f" K1,1'00000000&r;;.l.o FUEL ROD I 000000000 A I 000000000 I 0008 000 we WATER I 000 we 000 CHANNEL I 000 000 I I 000000000 I..H I 000000000 c...I/I I l.l====:::J S | |||
.1 I CHANNEL FUEL ROO PELLET WATER CHANNEL I DIM.1.0.B I C D I E 1 F C H I I I DIM INCHES 0.080 I 5.278 I 0.J80 lit*1...1...lit...-T...I CONTROL BLADE'I BUNDLE LATTICe: CElL-.DIM.I.D.J I K I L l.t j N I 0 P 0>j..,>JA: | |||
S I DIM INCHES 1.58 1 4.875 I 0.250 I lit I lit I...*'0.281T 12.00*See fu,ference 25.**This value is bHS("d on 80 mil channels.Channel thickness can be 80 or 100 mil.lk-\SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-2b CORE CELL FANP ATRlUM-9B FU8L REV.15, APRIL 2004 LSCS-UFSAR ATRlUM-9B DESIGN 4-1 I!p o B I-"I-I)Ii...I I L W'o00000000 I rf'0 0 d bocoocY 000000000\,..;.. I 000000000 I 00 boo Mol 0008 000 1 0 00 000 000 we 000 I 000 we 000 000 000 I 000 000 000000000 I 000000000 I 0000 F 000000 E OO-d I J)I ,f" K1,1'00000000&r;;.l.o FUEL ROD I 000000000 A I 000000000 I 0008 000 we WATER I 000 we 000 CHANNEL I 000 000 I I 000000000 I..H I 000000000 c...I/I I l.l====:::J S | |||
.1 I CHANNEL FUEL ROO PELLET WATER CHANNEL I DIM.1.0.B I C D I E 1 F C H I I I DIM INCHES 0.080 I 5.278 I 0.J80 lit*1...1...lit...-T...I CONTROL BLADE'I BUNDLE LATTICe: CElL-.DIM.I.D.J I K I L l.t j N I 0 P 0>j..,>JA: | |||
S I DIM INCHES 1.58 1 4.875 I 0.250 I lit I lit I...*'0.281T 12.00*See fu,ference 25.**This value is bHS("d on 80 mil channels.Channel thickness can be 80 or 100 mil.lk-\SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-2b CORE CELL FANP ATRlUM-9B FU8L REV.15, APRIL 2004 LSCS*UFSAR,I B we CHANNEL o p<STC0160:.DGN CHANNEL FUEL ROD PELLET WATER CHANNEL DIM LD.A I B I C D I E I F G H I I DIM INCHES 0.100 I 5.278 I 0.:38*I*I***I*CONTROL BLADE BUNDLE LATTICE CELL DIM LD.J I K ILM I N I 0 P Q I S DIM INCHES 1.58 I 4.875 I 0.260*I*I**0.261 I 12.0*See Reference 26 LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-2c CORE CELL FANP ATRIUM-IO FUEL REV.15, APRIL 2004 LSCS*UFSAR,I B we CHANNEL o p<STC0160:.DGN CHANNEL FUEL ROD PELLET WATER CHANNEL DIM LD.A I B I C D I E I F G H I I DIM INCHES 0.100 I 5.278 I 0.:38*I*I***I*CONTROL BLADE BUNDLE LATTICE CELL DIM LD.J I K ILM I N I 0 P Q I S DIM INCHES 1.58 I 4.875 I 0.260*I*I**0.261 I 12.0*See Reference 26 LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-2c CORE CELL FANP ATRIUM-IO FUEL REV.15, APRIL 2004 LSCS-UFSAR C ar:rI:1DIR o:.l TE FuelRcxl o | |||
--1006.06.666 Dfi)OeOOeOeo 30000000000 500001)000 00°0-0000 000__0 0000000000 LatgeCmtralW amrRa:l 90410_00410(10 Notes: 1)View of bundle lattice looking down from top.2)Channel fastener is atA1 corner.LASAlLE COUNTY STATION UPDATED FINAL SAFETY AJ."lALYSIS REPORT FIGURE 4.1-2d GE14 LATTICE ARRANGEMENT REV.16, APRIL 2006 LSCS-UFSAR C ar:rI:1DIR o:.l TE FuelRcxl o | |||
--1006.06.666 Dfi)OeOOeOeo 30000000000 500001)000 00°0-0000 000__0 0000000000 LatgeCmtralW amrRa:l 90410_00410(10 Notes: 1)View of bundle lattice looking down from top.2)Channel fastener is atA1 corner.LASAlLE COUNTY STATION UPDATED FINAL SAFETY AJ."lALYSIS REPORT FIGURE 4.1-2d GE14 LATTICE ARRANGEMENT REV.16, APRIL 2006 LSCS - UFSAR | |||
Rev.20, APRIL2014 LSCS*UFSAR CHANNIL'ASTINI" ASlEMILY"UII..lUND""'A'1..HANOL'IIUIL"OO'NTI"'" I S'ACI"'UNUM flUIL CHANNIL"'IHQ GETTER\.OWl..Ttl'&.ATI II'HOI" 5"tI'ING'U., rTY!'.,.....'0'''',.I..LI'f HOSIPlICI lIu.L"OO LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1*3 (fYPICAL)FUEL ASSEMBLY (GE 8X8R SHOWN)REV.13 LSCS*UFSAR CHANNIL'ASTINI" ASlEMILY"UII..lUND""'A'1..HANOL'IIUIL"OO'NTI"'" I S'ACI"'UNUM flUIL CHANNIL"'IHQ GETTER\.OWl..Ttl'&.ATI II'HOI" 5"tI'ING'U., rTY!'.,.....'0'''',.I..LI'f HOSIPlICI lIu.L"OO LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1*3 (fYPICAL)FUEL ASSEMBLY (GE 8X8R SHOWN)REV.13 LAROE CENTRAL WATER ROD FUEL ROD LOWER TIE PLATE LSCS-UFSAR CHANNEL LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-3a FUEL ASSEMBLY GE 8x8NB FUEL TYPE REV.13 LAROE CENTRAL WATER ROD FUEL ROD LOWER TIE PLATE LSCS-UFSAR CHANNEL LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-3a FUEL ASSEMBLY GE 8x8NB FUEL TYPE REV.13 W<i...J a..\:::!I-et:: w a..a..::J a o a::: W<<....J a..\:::!I-<<LSCS-UFSAR l.l.J U s: w a 00 oz u\:::!ox I-..JCO l'*.+t*;0 f.1.........0 o..............*0 0-1......t....j.........u 3: o....JeD et::...J::J.<<LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1*8b FUEL ASSEMBLY FANP ATRIUM*9B FUEL REV.15,.('\PRIL 2004 W<i...J a..\:::!I-et:: w a..a..::J a o a::: W<<....J a..\:::!I-<<LSCS-UFSAR l.l.J U s: w a 00 oz u\:::!ox I-..JCO l'*.+t*;0 f.1.........0 o..............*0 0-1......t....j.........u 3: o....JeD et::...J::J.<<LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1*8b FUEL ASSEMBLY FANP ATRIUM*9B FUEL REV.15,.('\PRIL 2004 FRAMATOME-ANP LSCS-UFSAR | |||
-'WATER CHANNEL:I': ULTRAFLOW 111 , SPACER L:.IY_.*IY"-LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-3c FUEL ASSEMBLY FA,"J"P ATRIUM-9B FUEL REV.15, APRIL 2004 FRAMATOME-ANP LSCS-UFSAR | |||
-'WATER CHANNEL:I': ULTRAFLOW 111 , SPACER L:.IY_.*IY"-LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-3c FUEL ASSEMBLY FA,"J"P ATRIUM-9B FUEL REV.15, APRIL 2004 LSCS*UFSAR | |||
//;.7FRAMATOME ANPPART*LENGTH FUEL ROD/'lOWER FUEL/'A FlOO ADAPTER ,/A SPRING SLEEVE ATRIUM TII 10 WATER CHANNEL FUa ASSEMal.)' | |||
FOM IIOIUNG WATER REACTOR.LASALLE COUNTY STATION UPDATED FINAL SAFETY A.l"lALYSIS REPORT FIGURE 4.I-3d FUEL ASSEMBLY FilliP ATRIUM-I0 FUEL REV.15, APRIL 2004 LSCS*UFSAR | |||
//;.7FRAMATOME ANPPART*LENGTH FUEL ROD/'lOWER FUEL/'A FlOO ADAPTER ,/A SPRING SLEEVE ATRIUM TII 10 WATER CHANNEL FUa ASSEMal.)' | |||
FOM IIOIUNG WATER REACTOR.LASALLE COUNTY STATION UPDATED FINAL SAFETY A.l"lALYSIS REPORT FIGURE 4.I-3d FUEL ASSEMBLY FilliP ATRIUM-I0 FUEL REV.15, APRIL 2004 LSCS-UFSAR Spacer Fue/Rod Upper Tie Plate Debris Filter Lower Tieplate Spacer Two Large Central Water Rods Fourteen Part-Length Rods T lower Tleplate (it:..., 1-i I! | |||
v i (1 i LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-3e GE14 Fuel Bundle (Typical)REV.16, APRIL 2006 LSCS-UFSAR Spacer Fue/Rod Upper Tie Plate Debris Filter Lower Tieplate Spacer Two Large Central Water Rods Fourteen Part-Length Rods T lower Tleplate (it:..., 1-i I! | |||
v i (1 i LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-3e GE14 Fuel Bundle (Typical)REV.16, APRIL 2006 143 tn. | |||
HANDLE Io iJ SHEATH..JI....1-..."-----BLAD EVELOCITY LIMITER COUPLING SOCKET-------LA SALLE COU NT.Y STATION UPDATED FINAL SAFETY ANALYSIS REPOkT FIGURE 4.1-4 GENERAL ELECTRIC CONTROL ROD ASSEMBLY Rrr.5-APRIL 1989 143 tn. | |||
HANDLE Io iJ SHEATH..JI....1-..."-----BLAD EVELOCITY LIMITER COUPLING SOCKET-------LA SALLE COU NT.Y STATION UPDATED FINAL SAFETY ANALYSIS REPOkT FIGURE 4.1-4 GENERAL ELECTRIC CONTROL ROD ASSEMBLY Rrr.5-APRIL 1989 LSCS-UFSAR 6.5 ,n0t"''''I(I (;'1(1 0 , t, II)SHEATH----...-.IJ-..,..... | |||
143 In Coupling Release Handle COUPLING SOCKET--......_\...,--------SLAOE LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-4a GENERAL ELECTRIC ORIGINAL EQUIPMENT CONTROL ROD ASSEMBLY REV.14, APRIL 2002 I LSCS-UFSAR 6.5 ,n0t"''''I(I (;'1(1 0 , t, II)SHEATH----...-.IJ-..,..... | |||
143 In Coupling Release Handle COUPLING SOCKET--......_\...,--------SLAOE LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-4a GENERAL ELECTRIC ORIGINAL EQUIPMENT CONTROL ROD ASSEMBLY REV.14, APRIL 2002 I LSCS-UFSAR lIathium Plate SHEATH COUPLING RELEASE HANDLE (LOWER)COUPLING SOCKET----------,. | |||
()()()()()()Cl!J UPPER HANDLE NEUTRON ABSORBEF RODS-3 AT TIP ARE VELOCITY LIMITER LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1*4b GENERA.L ELECTRIC TYPlCAL DURALIFE 215 CONTROL ROD ASSEMBLY REV.14.APRIL 2002 I LSCS-UFSAR lIathium Plate SHEATH COUPLING RELEASE HANDLE (LOWER)COUPLING SOCKET----------,. | |||
()()()() | |||
()()Cl!J UPPER HANDLE NEUTRON ABSORBEF RODS-3 AT TIP ARE VELOCITY LIMITER LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1*4b GENERA.L ELECTRIC TYPlCAL DURALIFE 215 CONTROL ROD ASSEMBLY REV.14.APRIL 2002 I LSCS-UFSAR LASALLE COUNTY ST A nON UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.14c GENERAL ELECTRIC TYPICAL MARATHON CONTROL ROD ASSEMBLY REV.14.APRIL 2002 LSCS-UFSAR LASALLE COUNTY ST A nON UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.14c GENERAL ELECTRIC TYPICAL MARATHON CONTROL ROD ASSEMBLY REV.14.APRIL 2002 | |||
"-J_....WET sTEAM-RETURNING WATER STEAM WATER MIXTURE\i.WATER LEVEL RETURNING WATER TURNING VANES (INLET NOZZLE)STANDPIPE CORE DISCHARGE LA SALLE COUNTY STATION UPDATED FINAL"SAFETY ANALYSIS REPORT FIGURE 4.1-5 STEAM SEPARATOR REV.0-APRIL 1984"-J_....WET sTEAM-RETURNING WATER STEAM WATER MIXTURE\i.WATER LEVEL RETURNING WATER TURNING VANES (INLET NOZZLE)STANDPIPE CORE DISCHARGE LA SALLE COUNTY STATION UPDATED FINAL"SAFETY ANALYSIS REPORT FIGURE 4.1-5 STEAM SEPARATOR REV.0-APRIL 1984 STEAM DRYE R SKIRT VANES COLLECTING TROUGH LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-6 STEAM DRYER REV.a-APRIL 1984 STEAM DRYE R SKIRT VANES COLLECTING TROUGH LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-6 STEAM DRYER REV.a-APRIL 1984 | |||
***LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-7 STEAM DRYER PANEL REV.0-APRIL 1984***LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-7 STEAM DRYER PANEL REV.0-APRIL 1984 CEN11!RINQ SPRING COHTAOl ROD GM TO#'GWOE LA SALLE COUNTY STATION UPDATED FrNAL SAFETY ANALYSIS REPORT FIGURE 4.2-1 SCHEMATIC OF FOUR BUNDLE CELL ARRANGEMENT REV.a-APRIL 1984 SCHEMATIC OF REACTOR ASSEMBLY SHOWING THE LEAKM}E FLOW PATHS34 NOTE: PERIPHERAL FUEL SUPPORTS ARE WeLOl'iD INTO THE COAE SUPPORT PLATE.FOR THESE/BUNDLES.PATH NUMBERS 1.2, 5, AND 7 DO NOT EX 1ST.6 FUEL SUPPORT IN-cORE GUIDE ruBe...I SHROUD CONTROL GUiDE TUBE CHANNEL o LOWER TIE PLATE 7 I.CONTROL ROD GUIDE TUBe*FUEL SUPPORT 2.CONTROL ROO GUIDE TUSE*CORE SUPPORT PLATE 3.CORE SUP!"Ol'tT PLATE*INCOAe Tuae 4.COl'll!SUPPa"T PLATE--SHROUD 5.CONTROL ROO GUIOE TUBE*DRIVE HOUSING G.FuEL SUPPORT'l..OWEA He PLATE 7.CONTAOL ROD DRIVE COOl.ING WATER a CHANNEL-LOWER TIE PLATE 9.LOI'VER TIE PLATE HOl.ES{lWO/A$SEMBLYj CONTROL ROD DRIVE HOUSING LA SALLE COUNTY STATION VPDiHED f HUlL SAfETY ,l\NAl YSI S REPORT FIGURE 4.2-2 BYPASS FLOW PATHS0-APRIL 1984 , | |||
,.FIGURE 4.2-3 FUEL BUNDLE eXeR BP8X8R YJEL TY?ES LA SALLE COUNTY STATION UPDA7ED FiNAL SAFETY ANAL vsIS REPORT II I , ii" I I1---------------..",;1 | |||
[i i I X I-<:):2 eo\.\I i.J a.JC lU LoU II!.:.).....<Ii Il.*;: C III 0;>a:III....J U"" o:t:>LL...I.....<<u ii:>-...1-""r""-.-,....flI'......w (I Z II:...z i*I I<<x...l-J.....::--.::;: '"'""'" f""1"'"'4-APRIL 1988 EXPANSION SPR:NG SPACER LOCKING TAe WASHER LARGE'W,A.TER ROD LowER TIE PLATS j},...----------.,1 LA SALLE COU NTY STATION"!UPDATED FINAL SAFETY ANALYSIS REPORT 1.f J---------------...1, I l i FIGUR£4.2-3aIl!I GE I i4- 1988 LSCS-OFBil]!I Improved lie plote c:luigl"l. | |||
Ir"--Increosed plen...'" volumein Gooolinio fods-+Circular ferrylt high per10rmllnce 5poeer Laroe woter tube[1..3'4 inl:n O.D./0.040 ineh thiclenessl os spoeer ccptur.IfF I...-Rel:Juced tube 0.0.1I If thfu UTP tlq IJ I,K I 11,1 inch of std.aEeS v II 0.591 inch W/R tl.lbel rI | |||
...-:::;: | |||
Helium r-,..preprest",rizatiol"l level 15 A nu.Godolinio stu/ldown Zone' I r-O.64tO inch..."&I&1 V rod 0 red l:utcllio 0.411 inch pd.t o.o.*I 1,..--0.483 inch fuel rod 0.0.'0,410 ineh pellet length.---_....JLIlll. | |||
12 inch NA TU illcrn,k.t 01 top and Ij inch at bottom LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE: 4.2-3b\I ,!Fuel Bundle GE 8X8NB Fuel Type REV,8-APRJl 1992 o o a:: Io&J...<-I Q.Io&J i=<t LSCS-UFSAR l.U U:>w o 0<':>oz a::S2 u I--ICOIWl0----u::I...t t..**i..,_00*..............0 0--...t't**+..t..+*.<LASALLE COUNTY STATION UPDATED FIN.AL SAFETY ANALYSIS REPORT FIGURE 4.2-3c FUEL BUNDLE FANP ATRIUM*9B TYPE REV.15, APRIL 2004 LSCS-UFSAR v U i'III c Ll\.SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.2-:3d FUEL BUNDLE FANP ATRIUM*10 TYPE REV.15, APRIL 2004 I LSCS-UFSAR J I')x....T iIol.......)oIl;""";:- | |||
"'" I ('""'.....T I Spacer Lower Tieplate Upper Tie Plate Two Large Central Water Rods FoUrteen Part-Length Rods Spaoer Debris Filter Lower Tleplate LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.2-3e GE14 Fuel Bundle (Typical)REV.16, APRIL 2006 LSCS - UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT Figure 4.2-3f GNF2 Fuel Bundle (Typical) | |||
Rev. 20, APRIL 2014 LUCS-UFSAR REV. 14, APRIL 2002 | |||
Page Intentionally Left Blank LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.2-4 DELETED GUIDf TUBE'JE LOC In CONT R OL"'aD DR HOUSING GR1FICf./RflL.LER$/1 f,JEL$I.JPFORT CASTING COPE SUPPOFlT PLATE_--.I----+--.- | |||
CONTROL rwo 144,0, STROKE COLJPUNG LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.2-5 CONTROL ROD VELOCITY LIMITER REV.0-APRIL 1984 LSCS-UFSAR"i--", | |||
....,....-...--...;_0__OOQ.:v:0. | |||
... | |||
..:I...-..__._......."'<L....$__-.ell>to..o'h.__0ClUI_1_r.T lJ]..... | |||
I'D&.O.. | |||
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__ | |||
..-........... | |||
UlCoII_"j 24A5375N_. | |||
-1 1='';;)''''.' | |||
.\LASALLE COUNTY STATIONUPDATEDFINAL SAFETY ANALYSIS REPORT FIGURE 4.2-5a FABRICAST VELOCITY LHVllTER REV.14.APRIL 2002 IoSCS"UFS"'Ut 1l1Q iL 10)'4.....II: l::J""I.....*E i" z......is*g(,)!I.ASALJ.r: | |||
(:OUNTY ST,\TIOr-: | |||
UPDATED FI)J:\I, SAFETY ANAl.YSIS REPORT FIGURE'l.2-6 | |||
('1'il'ICAL) | |||
TEMI'ERATURE VB HEAT FLUX-DOL BX8R FUEL TY?E HEV.1.1 LSCS-U'SAJl 1llC1O LAS.".LLE COUNT'{STATTON UPDATED SAFEr):".'J"ALYSIS REPORT FIerRF.4,2*7 mL'ICAL) | |||
TF.MPF:RATl!Rf: | |||
VS.HEAT PLUX-Uf'£8X8R FeEL 1'\"PB to ,-.-----------------------------------,_l 1.101 A$$LWlJI'TIONS; IH ALL ADDS IN THE SI,IHOLf Ar'tHE SAME n:UPI!"AT1JItE I2l ACDS Al!A.CH IHOtCATEI:I 100 L:\SALLE COUNTY STATION UPDATED FfNN, S,\FE'ry ANALYSrS R!';f>ORT E:.'-lERGY RELEASE AS 1\ | |||
OF Trr,m r,TYPICAL) | |||
REV, 13 LSCS-UFSAR THIS PAGE INTENTIONALLY LEFT BLANK LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.3-1 a UNIT 1 CORE LOADING MAP FIGURE 4.3-1a REV. 10 -APRIL 1994 LSCS-UFSAR THIS PAGE INTENTIONALLY LEFT BLANK LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.3-1 b UNIT 2 CORE LOADING MAP FIGURE 4.3-1b REV. 10 - APRIL 1994 REV. hllhfidNI)l 4 w A: RTL =9~8 0 3 Y W a s X | |||
:L LA SALLE COUNTY STArIQN UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.3-2 L K- AS A FUNMON OF EXPOSURE AT VARIOUS VOID FRACTIONS, HIGH ENRICHMENT, DOMINANT FUEL TYPE ~=y ~~~ | |||
rv d NQf.J.~b~b~ ~'IO.tb+ C~Z~d tv c~ c c~ c~ v G NCYil~tf5dd Wt.}lt+' ~Z~d PZ'u . L - APRIL 1988 N0il:)VtfJ W01'15ezil FIGURE 4.3-3 tv 3 x W eti m a a c C ^. v NQh1~11tid Wd1V~ g~~ft LA SALLE C4UN'1'Y S'TJ~kT10N UPDATED FINAL SAFETY ANALYSIS R£PGR'f ATOM FRACTION AS A FUNCTION OF EXPOSURE, HIGH ENRICHMEIT, DOMINANT FUEL. TYPE. 4 , 014 10I DS { TYP=u:_: i Swoilotri~zl N ISSl~ 9-+ETd. 4 - APYIL 198,8 FIGURE 4.3-4 :u LA SALLE COUNTY STATION UPDATED FINAL SAF Y ANALYSIS REPORT FISSION FRACTION AS A FUNCTION OF EXPOSURE, HIGH ENRICHMENT, DOMINANT FUEL TYPE, 4 04 VOIDS (.7IPIC 11 f r.SCS-UFSAR 4 R 9 7 9 EXPOSUR9 (Gw4h) 9 I-ARALLE COUNTY STATION UPDATED FINAL SAFETY A 1ALYM REPORT vlc~fRrl 4.,3-5 (TYPICAL) | |||
NEL"i'RON G&NERATION TItiIE VS_ EXPMURE AT 40 PERCENT VOIDS RED'_ 13 1Q 1e LSCSXFSA R. 15 14 10 C?' = ENAICHMENT FU IL 1J3 &d"ICWtt3" RUEL 0.711 FUEL (NATilrt^L w 5 i FJtpo"! (aftht 10 I,AS4,L1X COU. N" N STATI ON L'PDMED FINAL SAFb'TY ANALYSIS REPORT FIGURE 4.3-6 (TI'PICAL) | |||
DELAYED NEUTRON FRACTION VS. EXPOSURE 4T 40 PERCENT VOIDS REV. 13 LSCS-UFSAR THIS PAGE INTENTIONALLY LEFT BLANK FIGURE 4.3-7 REV. 4 - APRIL 1988 LSCS-UFSAR THIS PAGE INTENTIONALLY LEFT BLANK FIGURE 4.3-8 REV. 4 - APRIL 1988 LSCS-UFSAR THIS PAGE INTENTIONALLY LEFT BLANK FIGURE 4.3-9 REV. 4 - APRIL 1988 LSCS-UFSAR THIS PAGE INTENTIONALLY LEFT BLANK FIGURE 4.3-10 REV. 4 - APRIL 1988 LSCS-UFSA-P 0 FIGURE 4.3-11 RE'. 13 LASA2.LR COUNTY STATION N N W J K K YJ Q a UPDATED FINAL SAFETY ANALYSIS REPORT BFGftvh-INC OF CYCLE AND END OF CYCLE CORE 81 ,'ERAGE XXIAL POWER - ?G4 CORE, BWIV4 AND $WEilsi x 4 d w-Z w 1 6 -t2 ,I SCS-ITFSAR 40 50 PERCENT volas lalml LASALLE COVINT1' STATION UPDATED FINAL SAFETY ANALYSTS REPORT FIOUkh 4_3-tl MODERATOR VOID RE!tCTMTY COEFFICIENT AT FOC-l INITIAL CYCLE RIM 13 LSC~J11{ p 1qQ i40 I20D Ww 2000 AVEAACE fUEL TEWA CRATURE PCi 2400 2804 e b Gwett LASMIE WUN'1Y hI'ATION UNP,1TEU I?INAL SAH'Fl'Y ANALYSIS RKI}r}lid' FIGURE 4.3-1.1 DOPPLER REACH VITY CQIT'FICI I~NT ,1S A FIJN(''C(UN OF VU EL FXPOSURR AN 1) AVF.RACl? Fi"KLTEMNRRJi`!'URE. A1' ,W AVER-AGE vUm caN'rvNT OF mm i iju l N'N li~(a (kII:N`i' (INITIAL CY(,I,i , ,) | |||
2.00 1.50 1.00 us 0.50 v 0.00 ISCS. U FS_AR Cold Shutdown Margin Cyde Exposure, +GWD/MT 0.0 - 5.0 10.0 15.0 20.0 LASALLE COUNTY STATION tTP1)!lTKf) | |||
PINAL SctPP.TY ANALYESIS | |||
?,PORT FIGURE 4.3-14 E?LWFIX OF A COLD SH1ITROWN 34ARGIN CURVE REV. 13 43 . 19 ..--..--- 25 07 02 t}d 10 14 1B 22 26 30 34 38 4? 46 50 54 58 - IA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.3- l S CONTROL ROD ASSIGNMENTS FOR GROUPS I THROUGH 4 (SEQUENCE A) RrV. 0 -- APRTL 1984 a~0~o~o~a~a~a~a 0 mammon iaiaioiaiaioioi 39 27 23 REV. 0 - APRIL 1984 42 06 10 14 15 22 26 30 34 35 42 46 50 54 S,a LA SALLE C"OUNTY STATION UPDATLD FTNAL SAFETY ANALYSIS REPORT FIGURE 4.3-16 CONTROL ROD ASSIGNMENTS FOR GROUPS a THROUGH 1 {7 (SEQUERC A ) ON o~a~no a~o~a on mmummmus =-SOUND REV. 0 - APRIL 1984 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT CONTROL_ ROIL ASSIGNMEN7S FOR GROUPS I THROUGH 4 (SEQUENCE B) sioioioioiaioio 59 r 3 .. $!) i '1 t . 51 3 4 3 41 r 2 1 2 43 --3 .. 3 4 39 -1 2 1 15 -.. 3 ... 3 3.1 -2 1 1 | |||
* 21 -.. 3 4 3 23 -t 2 1 19 -3 4 3 .. lS '-2 1 2 11 4 3 (J1 2 1 OJ L 3 .. 06 III l4 13 21 26 | |||
* 0 -A?R!L 1984 " 3 :J 2 1 2 i 3 .4. 3 , 2 1 2 M 4 l 4 J '1. 1 '1 1 3 " 3 c 1 2 1 2 l 4 3. 4-l 1 2 t .. 1 4 3 1 1-1 2 ..J l ... 3 2 1 2 .. 3 " 30 35 42 45 50 54 sa LA SALLE COuNTY ST ATION UPDATED FIHAL SAFETY ANALYSIS REPORT FIGURE 4.3 .. 17 CONTROL ROD ASSIGNf.1£N7S FOR GROUPS 1 THROUGH 4 (ScQUEi',CE B} | |||
43 39- 35 27 19 ---.-....- FIGURE 4.3-13 Q 06 10 14 18 22 2f, 30 34 38 42 46 50 54 sa LA SAILt-E COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT COUTROL ROD ASSIGNMENTS FOR GROUPS B THROUGH 10 (SEQUENCE B ) REV. 0 - A-PRIL 1984 oia Iwo mmmummmummonsam aoioiIm TID CONTROL FRACTION LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.3-1D HOT OPERATING EOO-1 SCRAM REACTIVITY REV. 0 -- APRIL 1984 i.SCS-ITFSAX X 4 t "xV '"AisonmNaNvx LA.SALLE COUNTY STATION UPDATED FFN.~L SAFETY ANALYSIS REPORT PIGUR8 4.3-20 ti t W XENON REAC11VITY BUILDUP AND BURNOUT AFTER SHL"TDOWN (TYPICAL} | |||
RED'. 13 REV. 19, APRIL 2012AT3323MWt LSCS-UFSAR | |||
!,J -' ".:: r--------_ | |||
.. tU't _____ "-____ """'"' _____ ..... ____ ...... _____ ...... ____ -.li.-.01 o l)t.' COA ft ""OliA j"..,... | |||
LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4,3-21 TYPICAL RADIAL POWER DISTRIBUTION LSCS-UFSA R REV. 19, APRIL 2012AZIMUTHALFASTFLUXDISTRIBUTIONFIGURE4.3-21aUPDATEDFINALSAFETYANALYSISREPORTLASALLECOUNTYSTATION REV. 19, APRIL 2012AT3323MWt LSCS-UFSAR ,K, | |||
__ -L __ __ | |||
4:1 .0 LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.3-22 TYPICAL AXIAL POWER DISTRIBUTION LSCS-UFSA R REV. 19, APRIL 2012AXIALFASTFLUXDISTRIBUTIONFIGURE4.3-22aUPDATEDFINALSAFETYANALYSISREPORTLASALLECOUNTYSTATION 1.0 0.02 001 D.D!0,02 iOS a IG Of;C:AV l';ATIO X2lXo uo10 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-1 DAMPING COEFFICIENT VS, DECAY RATIO (SECOND ORDER SYSTEMS)REV.0-APRIL 1984 REACT!VITY...PERTURBATI01>l REACTOR K!NETrCS t | |||
__....,.......NEUTRON flUX RESPONSE TOrAL REACTOR REACT IIIL T't FEEDBACK TOTAL[NDIVIDUAl CHANNEL TYPE REACTIVITY FEEDBACK FROM OTHER CHANNEL TYPES+REACTIVITY TO POWER TRANSFER FUNCTION AT CONSTANT INLET FLOW REACTIVITY TO flOW TRANSFER FUNCTION AT CONSTANT POWER ROIl TO POWER TRANSFER fUNCTION TO QTH[R CHANNELS LA COUNTY STATION UPDATID FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-2 HYDRODYNAMIC ANO CORE STABILITY MODEL REV.0-APRIL 1984 | |||
"1_VAt..'f:ii.J.....,._'---_._----------SALLE: COUNTY STAllON UPlI/\!Ell fUW'lAf[TY ft,IJAlkHORl ,...1111 18P1'.l;"'l'4I lJ.n Oel l:ld tntt.l1r-. | |||
....dp){If tll'I*'c.t!llflll&nu Ia.lilnicl.OUk3l1 lio.epS. | |||
ltf Nl.11l.\II'!." U''(llt'l,jl.ibnll | |||
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fllel!Il'Aptr.mu.e* | |||
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"-:0*,00'MI p..t:u....llt**:lJIII"¥"1hnfl-.lJI | |||
}TtWOifl, s.l.ar4lll11ll! | |||
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b UfuilllU' | |||
.:-c.lil rtCl)'lll).ll.n.2 | |||
...h"lIt th.eoW 10 llvClLJ..WllLJ ,e,r,M DjCQ¥lI'.,., hll'l.lpi' IN l-fl.fld"''t'tll-llltfl | |||
...,e:4ClI...bo'C>>(!tl(lllttd 1>II4Ch'lolTy bl\.... | |||
.!Qollffl\'lll | |||
\I ill**YCfO r.eilil tIlt."""1(11 Pl:.t1U'::tS h<<-,l(!A1 Itlt Iutl 6l::unnj OOo,a'lltjU; m3L""..m" mI.t¥i | |||
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'''(1;1:2'; | |||
.., f1GHRE 4.4-3 I,IDon 8liXK fH V;1l V£fLOW REV.() | |||
LSCS-UFSAR FIGURE 4.4-4 REV. 4 - APRIL 1988 | |||
THIS PAGE INTENTIONALLY LEFT BLANK | |||
LSCS-UFSAR FIGURE 4.4-5 REV. 4 - APRIL 1988 | |||
THIS PAGE INTENTIONALLY LEFT BLANK LSCSUFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE INITIAL CORE 10 PSI PRESSURE REGULATOR SETPOINT STEP AT 51.5%RATED POWER (NATURAL CIRCULATION) | |||
Rev.14.APRIL 2002 LSCS-UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-7 INITIAL CORE 10 CENT ROD REACTIVITY STEP AT 51.5%RATED POWER (NATURAL CIRCULATION) | |||
Rev.14, APRIL 2002 15CS UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-8 INITIAL CORE 6-INCH WATER LEVEL SETPOINT STEP AT 51.5%RATED POWER (NATURAL CIRCULATION) | |||
Rev.14, APRIL 2002 LSCS-UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4A9 INITIAL CORE 10 PSI PRESSURE REGULATORY SETPOINT AT 105%RATED POWER AND 100%RATED FLOW Rev.14, APRIL 2002 LSCS UFSAI<LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-10 INITIAL CORE 10 CENT ROD REACTIVITY STEP AT 105%RATED POWER AND 100%Ri\TED FLOW Rev.14, APRIL 2002 LSCS*UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-11 INITIAL CORE 10%LOAD DEMAND STEP AT 105%RATED POWER AND 100%RATED FLOW Rev.14, APRIL 2002 LSCS UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4,4-12 INITIAL CORE 6-INCH WATER SETPOlNT STEP AT 105%RATED POWER AND 100%RATED FLOW Rev.14, APRIL 2002 LSCS-UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-13 INITIAL CORE 10 PSI PRESSURE REGULATOR SETPOINT STEP AT 68%POWER AND 50%RATED FLOW Rev.14, APRIL 2002 LSCS-UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-14 INITIAL CORE 10 CENT ROD REACTIVITY STEP AT 68%POWER AND 50%RATED FLOW Rev.14, APRIL 2002 LSCS-UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-15 INITIAL CORE 10%LOAD DEMAND STEP AT 68%POWER AND 50%RATED FLOW Rev.14, APRIL 2002 LSCS-UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-16 INITIAL CORE 6-INCH WATER LEVEL SETPOINT STEP AT 68%POWER AND 50%RATED FLOW Rev.14.APRIL 2002 | |||
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Il:.l.UI TIC"ltl'lioIGt.,*LA SALLE COUNTY STATION UPDATED FINAL SAfEH ANAl.VSIS REPORT FIGURE 4.6-6 PROCESS OATA.CONTROL ROn OR1VE HYDP.AUUC SYSTEt1 (SHEH 2 of 2)T!Tel-'-'-r-io'--' | |||
:R.I NOTE 1. A SINGLE SCRAM PILOT VALVE WITH DUAL SOLENOID OPERATED PILOT ASSEMBLIES MAY BE INSTALLED IN PLACE OF TWO SCRAM PILOT VALVES. SCRAM PILOT VALVES (NOTE 1) OUTLET SCRAM VALVE OPERATOR DIRECTIONAL CONTROL VALVES NITROGEN PRESSURE INDICATOR LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.6-7 CONTROL ROD DRIVE HYDRAULIC CONTROL UNIT REV. 18 - APRIL 2010 FIGURE 4.6-7 CONTROL ROD DRIVE HYDRAULIC CONTROL UNIT LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT SPEED CONTROL VALVES SCRAM PILOT VALVES (NOTE 1)NITROGEN VOLUME WITHDRAWAL LINE;_L..-.ELECTRICAL JUNCTION BOX OUTLET SCRAM VALVE OPERATOR/I'L.ELECTRICAL | |||
/CONDUIT CONNECTION INSERT LINEn I I I i i!INLET SCRAM VALVE ACCUMULATOR FRAME ACCUMULATOR l:iQIE 1.A SINGLE SCRAM PILOT VALVE WITH DUAL SOLENOID OPERATED PILOT ASSEMBLIES MAY BE INSTALLED IN PLACE OF TWO SCRAM PILOT VALVES.REV.18-APRIL 2010 foll:ro Bel.i JPnAl'ED F:NAI:.U NTY STATION-1 | |||
*h.I;)S REPORT I LA SALLE CO,I REVI.- | |||
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.... PEiH.HIIl. . -----=ST ATION CO U NT Y LA SALLE (I. ry "-l't-'O SMFE.. __ ',JPflM I:. | |||
* _. ___ _ FIGURE 4 .. 6 .. 8 ROC. DR IVE HOUSING SUPPORT REV. 0 -,l .... PRIL 1984}} | |||
Latest revision as of 13:39, 17 March 2019
Text
LSCS-UFSAR 4.0-i REV. 20, APRIL 2014 CHAPTER 4.0 - REACTOR TABLE OF CONTENTS Page 4.0 REACTOR 4.1-1 4.1
SUMMARY
DESCRIPTION 4.1-1
4.1.1 Reactor
Vessel 4.1-1 4.1.2 Reactor Internal Components 4.1-1 4.1.2.1 Reactor Core 4.1-1 4.1.2.1.1 General 4.1-1 4.1.2.1.2 Core Configuration 4.1-3 4.1.2.1.3 Fuel Assembly Description 4.1-4 4.1.2.1.3.1 Fuel Rod 4.1-4 4.1.2.1.3.2 Fuel Bundle 4.1-4 4.1.2.1.4 Assembly Support and Control Rod Location 4.1-5 4.1.2.2 Shroud 4.1-5 4.1.2.3 Shroud Head and Steam Separators 4.1-6 4.1.2.4 Steam Dryer Assembly 4.1-6 4.1.3 Reactivity Control Systems 4.1-7 4.1.3.1 Operation 4.1-7 4.1.3.2 Description of Rods 4.1-7 4.1.3.3 Supplementary Reactivity Control 4.1-8 4.1.4 Analysis Techniques 4.1-8 4.1.4.1 Reactor Internal Components 4.1-8 4.1.4.1.1 MASS (Mechanical Analysis of Space Structure) 4.1-9 4.1.4.1.1.1 Program Description 4.1-9 4.1.4.1.1.2 Program Version and Computer 4.1-9 4.1.4.1.1.3 History of Use 4.1-9 4.1.4.1.1.4 Extent of Application 4.1-9 4.1.4.1.2 SNAP (MU LTISHELL) 4.1-9 4.1.4.1.2.1 Program Description 4.1-9 4.1.4.1.2.2 Program Version and Computer 4.1-10 4.1.4.1.2.3 History of Use 4.1-10 4.1.4.1.2.4 Extent of Application 4.1-10 4.1.4.1.3 GASP 4.1-10
4.1.4.1.3.1 Program Description 4.1-10 4.1.4.1.3.2 Program Version and Computer 4.1-10 4.1.4.1.3.3 History of Use 4.1-11 4.1.4.1.3.4 Extent of Application 4.1-11 4.1.4.1.4 NO HEAT 4.1-11 4.1.4.1.4.1 Program Description 4.1-11 4.1.4.1.4.2 Program Version and Computer 4.1-11 LSCS-UFSAR Table of Contents (Cont'd) 4.0-ii REV. 20, APRIL 2014 4.1.4.1.4.3 History of Use 4.1-11 4.1.4.1.4.4 Extent of Application 4.1-12 4.1.4.1.5 FI NITE 4.1-12 4.1.4.1.5.1 Program Description 4.1-12 4.1.4.1.5.2 Program Version and Computer 4.1-12 4.1.4.1.5.3 History of Use 4.1-12 4.1.4.1.5.4 Extent of Application 4.1-12 4.1.4.1.6 SAMIS 4.1-12 4.1.4.1.6.1 Program Description 4.1-12 4.1.4.1.6.2 Program Version and Computer 4.1-13 4.1.4.1.6.3 History of Use 4.1-13 4.1.4.1.6.4 Extent of Application 4.1-13 4.1.4.1.7 General Matrix Manipulation Program(GEMOP) 4.1-14 4.1.4.1.7.1 Program Description 4.1-14 4.1.4.1.7.2 Program Version and Computer 4.1-14 4.1.4.1.7.3 History of Use 4.1-14 4.1.4.1.7.4 Extent of Application 4.1-14 4.1.4.1.8 SHELL 5 4.1-14
4.1.4.1.8.1 Program Description 4.1-14 4.1.4.1.8.2 Program Version and Computer 4.1-15 4.1.4.1.8.3 History of Use 4.1-15 4.1.4.1.8.4 Extent of Application 4.1-15 4.1.4.1.9 HEATER 4.1-15 4.1.4.1.9.1 Program Description 4.1-15 4.1.4.1.9.2 Program Version and Computer 4.1-15 4.1.4.1.9.3 History of Use 4.1-16 4.1.4.1.9.4 Extent of Application 4.1-16 4.1.4.1.10 FAP-71 (Fatigue Analysis Program) 4.1-16 4.1.4.1.10.1 Program Description 4.1-16 4.1.4.1.10.2 Program Version and Computer 4.1-16 4.1.4.1.10.3 History of Use 4.1-16 4.1.4.1.10.4 Extent of Application 4.1-16 4.1.4.1.11 CREEP/P LASTICITY 4.1-17 4.1.4.1.11.1 Program Description 4.1-17 4.1.4.1.11.2 Program Version and Computer 4.1-17 4.1.4.1.11.3 History of Use 4.1-17 4.1.4.1.11.4 Extent of Application 4.1-17 4.1.4.1.12 SAP4G07 and ANSYS 4.1-17 4.1.4.2 Fuel Rod Thermal Analysis 4.1-17 4.1.4.3 Reactor Systems Dynamics 4.1-18 4.1.4.4 Nuclear Engineering Analysis 4.1-18 4.1.4.5 Neutron Fluence Calculations 4.1-18 4.1.4.6 Thermal Hydraulic Calculations 4.1-19 LSCS-UFSAR Table of Contents (Cont'd) 4.0-iii REV. 15, APRIL 2004 4.1.5 References 4.1-19 4.2 FUEL SYSTEM 4.2-1 4.2.1 Design Bases 4.2-1 4.2.1.1 Safety Design Bases 4.2-1 4.2.1.2 Power Generation Design Basis 4.2-3 4.2.1.2.1 Material Selection 4.2-3 4.2.1.2.2 Effects of Irradiat ion and Fuel Swelling 4.2-3 4.2.1.2.3 Fuel De nsification 4.2-4 4.2.1.2.3.1 GE Fuel 4.2-4 4.2.1.2.3.2 FANP Fuel 4.2-5 4.2.1.2.4 Incipient UO 2 Center Melting 4.2-5 4.2.1.2.4.1 GE Fuel 4.2-5 4.2.1.2.4.2 FANP Fuel 4.2-5 4.2.1.2.5 Maximum Allowable Stresses 4.2-6 4.2.1.2.5.1 GE Fuel 4.2-6 4.2.1.2.5.2 FANP Fuel 4.2-7 4.2.1.2.6 Capacity for Fission Gas Inventory 4.2-7 4.2.1.2.7 Maximum Internal Gas Pressure 4.2-7 4.2.1.2.7.1 GE Fuel 4.2-7 4.2.1.2.7.2 FANP Fuel 4.2-8 4.2.1.2.8 Internal Pressure and Cladding Stresses During Normal Conditions 4.2-8 4.2.1.2.9 Cycling and Fatigue Limits 4.2-8 4.2.1.2.9.1 GE Analysis 4.2-8 4.2.1.2.9.2 FANP Analysis 4.2-9 4.2.1.2.10 Deflections 4.2-9 4.2.1.2.10.1 GE Evaluation 4.2-9 4.2.1.2.10.2 FANP Evaluation 4.2-9 4.2.1.2.11 Flow-Induced Fuel Rod Vibrations 4.2-10 4.2.1.2.12 Fretting Corrosion 4.2-10 4.2.1.2.13 Seismic Loadings 4.2-10 4.2.1.2.14 Chemical Proper ties of Cladding and Fuel Material 4.2-11 4.2.1.2.15 Design Ratios 4.2-11 4.2.1.2.15.1 Limiting Parameter Values 4.2-11 4.2.1.2.15.1.1 Normal and Upset Design Conditions 4.2-11 4.2.1.2.15.1.2 Emergency and Faulted Design Conditions 4.2-12 4.2.1.2.15.2 Actual Parameter Values 4.2-12 4.2.1.2.16 Fuel Assembly Limits 4.2-13 4.2.1.2.16.1 Fuel Rods 4.2-13 4.2.1.2.16.2 Fuel Spacer 4.2-13 4.2.1.2.16.3 Water Rods 4.2-14 LSCS-UFSAR Table of Contents (Cont'd) 4.0-iv REV. 15, APRIL 2004 4.2.1.2.16.4 Channel 4.2-14 4.2.1.2.16.5 Tie Plates 4.2-15 4.2.1.2.17 Reactivity Control Assembly and Burnable Poison Rods 4.2-15 4.2.1.2.17.1 Safety Design Bases for Reactivity Control 4.2-15 4.2.1.2.17.1.1 Specific Design Characteristics 4.2-16 4.2.1.2.18 Surveillance Program 4.2-17
4.2.2 Description
and Design Drawings 4.2-18 4.2.2.1 Core Cell 4.2-18 4.2.2.2 Fuel Assembly 4.2-18 4.2.2.3 Fuel Bundle 4.2-18 4.2.2.4 Fuel Rod 4.2-19 4.2.2.5 Fuel Pellets 4.2-21 4.2.2.6 Fuel Channel 4.2-22 4.2.2.7 Reactivity Control Assembly and Burnable Poison Rods 4.2-23 4.2.2.7.1 Control Rods 4.2-23 4.2.2.7.1.1 General Electric Control Rods 4.2-22 4.2.2.7.1.2 ASEA-ATOM Control Rods 4.2-24 4.2.2.7.2 Velocity Limiter 4.2-25 4.2.2.7.3 Burnable Poison Rods 4.2-25 4.2.3 Design Limits and Evaluation 4.2-26 4.2.3.1 Fuel Damage Analysis 4.2-26 4.2.3.2 Fuel Damage Experience 4.2-27 4.2.3.3 Potential for a Water-Logging Rupture 4.2-28 4.2.3.4 Potential for Hydriding 4.2-29 4.2.3.5 Dimensional Stability 4.2-29 4.2.3.6 Fuel Densification 4.2-30 4.2.3.7 Fuel Cladding Temperatures 4.2-30 4.2.3.8 Peaking Factors 4.2-31 4.2.3.8.1 Local Peak ing Factors 4.2-31 4.2.3.8.2 Axial and Gross Peaking Factors 4.2-31 4.2.3.9 Temperature Transients with Water- logged Fuel Element 4.2-31 4.2.3.10 Potential Damaging Temperature Effects During Transients 4.2-32 4.2.3.11 Energy Release During Fuel Element Burnout 4.2-32 4.2.3.12 Energy Release fo r Rupture of Water- logged Fuel Elements 4.2-34 4.2.3.13 Fuel Rod Behavior Effects from Coolant Flow Blockage 4.2-34 4.2.3.14 Channel Evaluation 4.2-35 LSCS-UFSAR Table of Contents (Cont'd) 4.0-v REV. 15, APRIL 2004 4.2.3.15 Fuel Reliability 4.2-36 4.2.3.16 Fuel Operating and Developmental Experience 4.2-37 4.2.3.17 Fuel Assembly 4.2-37 4.2.3.17.1 Loads Assessment of Fuel Assembly Components 4.2-38 4.2.3.18 Spacer Grid and Channel Boxes 4.2-38 4.2.3.19 Burnable Poison Rods 4.2-38 4.2.3.20 Control Rods 4.2-38 4.2.3.20.1 Materials Adequacy Throughout Design Lifetime 4.2-38 4.2.3.20.2 Dimensional and Tolerance Analysis 4.2-38 4.2.3.20.3 Thermal Analys is of the Tendency to Warp 4.2-39 4.2.3.20.4 Forces for Expulsion 4.2-39 4.2.3.20.5 Functional Failure of Critical Components 4.2-39 4.2.3.20.6 Precluding Excessive Rates of Reactivity Addition 4.2-39 4.2.3.20.7 Effect of Fuel Rod Failure on Control Rod Channel Clearances 4.2-39 4.2.3.20.8 Mechanical Damage 4.2-39 4.2.3.20.8.1 First Mode of Failure 4.2-40 4.2.3.20.8.2 Second Mode of Failure 4.2-40 4.2.3.20.9 Analysis of Guide Tube Design 4.2-40 4.2.3.20.10 Evaluation of Control Rod Velocity Limiter 4.2-41 4.2.3.21 Rod Bowing 4.2-41 4.2.3.21.1 GE Evaluation 4.2-41 4.2.3.21.2 FANP Evaluation 4.2-42 4.2.3.22 Fission Gas Release 4.2-42 4.2.3.23 Ballooning and Rupture 4.2-43 4.2.3.23.1 GE Evaluation 4.2-43 4.2.3.23.2 FANP Evaluation 4.2-44 4.2.4 Testing and Inspection Plan 4.2-44 4.2.4.1 Testing and Inspection (Enrichment and Burnable Poison Concentrations) 4.2-45 4.2.4.1.1 Enrichment Control Program 4.2-45 4.2.4.1.2 Gadolinia Inspections 4.2-46 4.2.4.1.3 Reactor Control Rods 4.2-47 4.2.4.2 Surveillance Inspection and Testing of Irradiated Fuel Rods 4.2-47 LSCS-UFSAR Table of Contents (Cont'd) 4.0-vi REV. 15, APRIL 2004 4.2.4.3 Operating Experience with Gadolinia-Containing Fuel 4.2-48 4.2.5 References 4.2-49
4.3 NUCLEAR
DESIGN 4.3-1
4.3.1 Design
Bases 4.3-1 4.3.1.1 Safety Design Bases 4.3-1 4.3.1.2 Power Generation Design Bases 4.3-2 4.3.2 Description 4.3-2 4.3.2.1 Nuclear Design Description 4.3-2 4.3.2.1.1 Fuel Nuclear Properties 4.3-3 4.3.2.2 Power Distributions 4.3-4 4.3.2.2.1 Local Power Distribution 4.3-5 4.3.2.2.2 Radial Power Distribution 4.3-5 4.3.2.2.3 Axial Power Distribution 4.3-5 4.3.2.2.4 Power Distributi on Calculations 4.3-6 4.3.2.2.5 Power Distribution Measurements 4.3-6 4.3.2.2.6 Power Distribution Accuracy 4.3-6 4.3.2.2.7 Power Distribution Anomalies 4.3-6 4.3.2.3 Reactivity Coefficients 4.3-7 4.3.2.3.1 Void Reactivity Coefficients 4.3-7 4.3.2.3.2 Moderator Temperature Coefficient 4.3-7 4.3.2.3.3 Doppler Reactivity Coefficient 4.3-8 4.3.2.3.4 Power Coefficient 4.3-9 4.3.2.4 Control Requirements 4.3-9 4.3.2.4.1 Shutdown Reactivity 4.3-9 4.3.2.4.2 Reactivity Variations 4.3-10 4.3.2.5 Control Rod Patterns and Reactivity Worths 4.3-11 4.3.2.5.1 Control Rod Withdrawal Sequences 4.3-11 4.3.2.5.1.1 Control Rod Withdrawal Sequences in the Startup Range 4.3-12 4.3.2.5.1.2 Control Rod Withdrawal Sequences in the RWM Power Range 4.3-13 4.3.2.5.1.3 Maximum Control Rod Worth Pattern with a Single Error in the RWM Power Range 4.3-14 4.3.2.5.2 Control Rod Worth Calculations 4.3-14 4.3.2.5.2.1 Control Rod Worth in the Startup Range and RWM Power Range 4.3-14 4.3.2.5.2.2 Control Rod Worth in the Reactor Power Range > 10% Rated Power 4.3-15 4.3.2.5.3 Scram Reactivity 4.3-15 LSCS-UFSAR Table of Contents (Cont'd) 4.0-vii REV. 15, APRIL 2004 4.3.2.6 Criticality of Reactor During Refueling 4.3-16 4.3.2.6.1 Criticality of Reactor 4.3-16 4.3.2.6.2 Criticality of Fuel Assemblies 4.3-16 4.3.2.7 Stability 4.3-16 4.3.2.7.1 Xenon Transients 4.3-16 4.3.2.7.2 Thermal Hydraulic Stability 4.3-17 4.3.2.8 Vessel Irradiation 4.3-17 4.3.3 Analytical Methods 4.3-17
4.3.4 References
4.3-18
4.4 THERMAL
AND HYDRAULIC DESIGN 4.4-1
4.4.1 Design
Bases 4.4-1 4.4.1.1 Safety Design Bases 4.4-1 4.4.1.2 Power Generation Design Bases 4.4-1 4.4.1.3 Requirements for Steady-State Conditions 4.4-1 4.4.1.4 Requirements for Transient Conditions 4.4-2 4.4.1.5 Summary of Design Bases 4.4-2 4.4.1.5.1 Fuel Cladding Integrity 4.4-3 4.4.1.5.2 Fuel Assembly Integrity 4.4-3 4.4.1.5.3 Fuel-Cladding Gap Characteristics 4.4-3 4.4.2 Description of Thermal Hydraulic Design of Reactor Core 4.4-3 4.4.2.1 Summary Comparison 4.4-3 4.4.2.2 Critical Power Ratio 4.4-3 4.4.2.2.1 Boiling Correlations 4.4-4 4.4.2.2.1.1 GE Fuel 4.4-4 4.4.2.2.1.2 FANP Fuel 4.4-4 4.4.2.3 Maximum Average Planar Linear Heat Generation Rate (MAPLHGR) 4.4-5 4.4.2.3.1 Design Power Distribution 4.4-6 4.4.2.4 Void Fraction Distribution 4.4-7 4.4.2.5 Core Coolant Flow Distribution 4.4-7 4.4.2.6 Core Pressure Drop and Hydraulic Loads 4.4-8 4.4.2.6.1 Friction Pressure Drop 4.4-8 4.4.2.6.2 Local Pressure Drop 4.4-9 4.4.2.6.3 Elevation Pressure Drop 4.4-9 4.4.2.6.4 Acceleration Pressure Drop 4.4-10 LSCS-UFSAR Table of Contents (Cont'd) 4.0-viii REV. 17, APRIL 2008 4.4.2.7 Correlation and Physical Data 4.4-11 4.4.2.7.1 Pressure Drop Correlations 4.4-11 4.4.2.7.2 Void Fraction Correlation 4.4-11 4.4.2.7.3 Heat Transfer Correlation 4.4-12 4.4.2.8 Thermal Effects of Operational Transients 4.4-12 4.4.2.9 Uncertainties in Estimates 4.4-12 4.4.2.9.1 Transition Boiling Uncertainties 4.4-12 4.4.2.9.2 Variation of Fuel Damage Limit 4.4-13 4.4.2.9.3 Effects of Misoriented Fuel Bundle 4.4-13 4.4.2.10 Flux Tilt Considerations 4.4-13 4.4.3 Description of the Thermal and Hydraulic Design of the Reactor Coolant System 4.4-13 4.4.3.1 Plant Configuration Data 4.4-13 4.4.3.2 Operating Restrictions on Pumps 4.4-14 4.4.3.3 Power-Flow Operating Map 4.4-14
4.4.3.4 Temperature-Power Operating Map (PWR) 4.4-14 4.4.3.5 Load-Following Characteristics 4.4-14 4.4.3.6 Thermal and Hydraulic Characteristics Summary Table 4.4-14 4.4.4 Evaluation 4.4-14 4.4.4.1 Critical Heat Flux 4.4-14 4.4.4.2 Core Hydraulics 4.4-14 4.4.4.3 Influence of Power Distribution 4.4-14 4.4.4.4 Core Thermal Response 4.4-15 4.4.4.5 Analytical Methods 4.4-15 4.4.4.5.1 Reactor Model 4.4-15 4.4.4.5.2 System Flow Balances 4.4-16 4.4.4.5.3 System Heat Balances 4.4-17 4.4.4.5.4 Uncertainties in Design Analyses 4.4-18 4.4.4.6 Reactor Stability Analysis 4.4-18 4.4.4.6.1 Introduction 4.4-18 4.4.4.6.2 Description 4.4-19 4.4.4.6.3 Solution Description for Thermal-Hydraulic Stability 4.4-19 4.4.4.6.4 Stability Criteria 4.4-20 4.4.4.6.5 Expected Oscillation Modes 4.4-21 4.4.4.6.6 Analysis Approach 4.4-22 4.4.4.6.7 Mathematical Model 4.4-23 4.4.4.6.8 Initial Core Analysis Results 4.4-24 4.4.5 Testing and Verification 4.4-25 4.4.6 Instrumentation Requirements 4.4-25 4.4.6.1 Loose Parts Monitoring System (Deleted) 4.4-25
4.4.7 References
4.4-27 LSCS-UFSAR Table of Contents (Cont'd) 4.0-ix REV. 15, APRIL 2004 4.5 REACTOR MATERIALS 4.5-1 4.5.1 Control Rod System Structural Materials 4.5-1 4.5.1.1 Material Specifications 4.5-1 4.5.1.2 Special Materials 4.5-2 4.5.1.3 Processes, Inspections and Tests 4.5-2 4.5.1.4 Control of Delta Ferrite Content 4.5-3 4.5.1.5 Protection of Materials During Fabrication, Shipping and Storage 4.5-3 4.5.2 Reactor Internals Materials 4.5-4 4.5.2.1 Material Spec ifications 4.5-4 4.5.2.2 Controls on Welding 4.5-6 4.5.2.3 Nondestructive Examination of Wrought Seamless Tubular Products 4.5-6 4.5.2.4 Fabrication and Processing of Austenitic Stainless Steel 4.5-6 4.5.2.5 Regulatory Guide Conformance Assessment 4.5-6 4.6 FUNCTIONAL DESIGN OF REACTIVITY CONTROL SYSTEMS 4.6-1 4.6.1 Information for Control Rod Drive Systems (CRDS) 4.6-1 4.6.1.1 Control Rod Drive System Design 4.6-1 4.6.1.1.1 Design Bases 4.6-1 4.6.1.1.1.1 General Design Bases 4.6-1 4.6.1.1.1.1.1 Safety Design Bases 4.6-1 4.6.1.1.1.1.2 Power Generation Design Basis 4.6-2 4.6.1.1.2 Description 4.6-2 4.6.1.1.2.1 Control Rod Drive Mechanisms 4.6-2 4.6.1.1.2.2 Drive Components 4.6-3 4.6.1.1.2.2.1 Drive Piston 4.6-3 4.6.1.1.2.2.2 Index Tube 4.6-4 4.6.1.1.2.2.3 Collet Assembly 4.6-4 4.6.1.1.2.2.4 Piston Tube 4.6-4 4.6.1.1.2.2.5 Stop Piston 4.6-5 4.6.1.1.2.2.6 Flange and Cylinder Assembly 4.6-5 4.6.1.1.2.2.7 Lock Plug 4.6-6 4.6.1.1.2.3 Materials of Construction 4.6-6 4.6.1.1.2.3.1 Index Tube 4.6-6 4.6.1.1.2.3.2 Coupling Spud 4.6-7 4.6.1.1.2.3.3 Collet Fingers 4.6-7 4.6.1.1.2.3.4 Seals and Bushings 4.6-7 4.6.1.1.2.3.5 Summary 4.6-7 LSCS-UFSAR Table of Contents (Cont'd) 4.0-x REV. 15, APRIL 2004 4.6.1.1.2.4 Control Rod Drive Hydraulic System 4.6-8 4.6.1.1.2.4.1 Hydraulic Requirements 4.6-8 4.6.1.1.2.4.2 System Description 4.6-9 4.6.1.1.2.4.2.1 Supply Pump 4.6-9
4.6.1.1.2.4.2.2 Accumulator Charging Pressure 4.6-10 4.6.1.1.2.4.2.3 Drive Water Pressure 4.6-10
4.6.1.1.2.4.2.4 Cooling Water Header 4.6-11 4.6.1.1.2.4.2.5 Return Line 4.6-11 4.6.1.1.2.4.2.6 Scram Di scharge Volume 4.6-11 4.6.1.1.2.4.3 Hydraulic Control Units 4.6-12 4.6.1.1.2.4.3.1 Insert Drive Valve 4.6-12 4.6.1.1.2.4.3.2 Insert Exhaust Valve 4.6-13 4.6.1.1.2.4.3.3 Withdr aw Drive Valve 4.6-13 4.6.1.1.2.4.3.4 Withdraw Exhaust Valve 4.6-13 4.6.1.1.2.4.3.5 Speed Control Valves 4.6-13 4.6.1.1.2.4.3.6 Scram Pilot Valves 4.6-13 4.6.1.1.2.4.3.7 Scram Inlet Valve 4.6-13 4.6.1.1.2.4.3.8 Scram Exhaust Valve 4.6-14 4.6.1.1.2.4.3.9 Scram Accumulator 4.6-14 4.6.1.1.2.4.3.10 Alternate Rod Insertion Scram Valves 4.6-14 4.6.1.1.2.5 Control Rod Drive System Operation 4.6-15 4.6.1.1.2.5.1 Rod Insertion 4.6-15 4.6.1.1.2.5.2 Rod Withdrawal 4.6-15 4.6.1.1.2.5.3 Scram 4.6-16 4.6.1.1.2.6 Instru mentation 4.6-17 4.6.1.2 Control Rod Driv e Housing Supports 4.6-17 4.6.1.2.1 Safety Objective 4.6-17 4.6.1.2.2 Safety Design Bases 4.6-17 4.6.1.2.3 Description 4.6-17 4.6.2 Evaluations of the CRDS 4.6-19 4.6.2.1 Failure Mode and Effects Analysis 4.6-19 4.6.2.2 Protection from Co mmon Mode Failures 4.6-19 4.6.2.3 Safety Evaluation 4.6-19 4.6.2.3.1 Control Rod Drives 4.6-19 4.6.2.3.1.1 Evaluation of Scram Time 4.6-19 4.6.2.3.1.2 Analysis of Malfunction Relating to Rod Withdrawal 4.6-20 4.6.2.3.1.2.1 Drive Housing Fails at Attachment Weld 4.6-20 4.6.2.3.1.2.2 Rupture of Hydraulic Line(s) to Drive Housing Flange 4.6-21 4.6.2.3.1.2.2.1 Pressure-U nder Line Break 4.6-21 4.6.2.3.1.2.2.2 Pressure-Over Line Break 4.6-22 4.6.2.3.1.2.2.3 Simultaneous Breakage of Pressure- Over and Pressure-Under Lines 4.6-22 LSCS-UFSAR Table of Contents (Cont'd) 4.0-xi REV. 15, APRIL 2004 4.6.2.3.1.2.3 All Drive Flange Bolts Fail in Tension 4.6-22 4.6.2.3.1.2.4 Weld Joinin g Flange to Housings Fails in Tension 4.6-23 4.6.2.3.1.2.5 Housin g Wall Ruptures 4.6-24 4.6.2.3.1.2.6 Flange Plug Blows Out 4.6-25 4.6.2.3.1.2.7 Drive Pressure Control Valve Closure (Reactor Pressure, 0 psig) 4.6-26 4.6.2.3.1.2.8 Ball Check Valve Fails to Close Passage to Vessel Ports 4.6-26 4.6.2.3.1.2.9 Hydraulic Co ntrol Unit (HCU) Valve Failures 4.6-26 4.6.2.3.1.2.10 Collet Fing ers Fail to Latch 4.6-27 4.6.2.3.1.2.11 Withdrawal Speed Control Valve Failure 4.6-27 4.6.2.3.2 Scram Reliability of CRDS 4.6-27 4.6.2.3.2.1 Reliabilit y Analysis 4.6-28 4.6.2.3.2.2 Control Rod Su pport and Operation 4.6-28 4.6.2.3.3 Control Rod Driv e Housing Supports 4.6-28 4.6.2.3.3.1 Safety Evaluation 4.6-28 4.6.3 Testing and Verifica tion of the CRDS 4.6-29 4.6.3.1 Control Rods 4.6-29 4.6.3.1.1 Testing an d Inspection 4.6-29 4.6.3.2 Control Rod Drives 4.6-29 4.6.3.2.1 Testing an d Inspection 4.6-29 4.6.3.2.1.1 Develo pment Tests 4.6-29 4.6.3.2.1.2 Factory Quality Control Tests 4.6-30 4.6.3.2.1.3 Operat ional Tests 4.6-31 4.6.3.2.1.4 Accepta nce Tests 4.6-31 4.6.3.2.1.5 Surveillance Tests 4.6-32 4.6.3.3 Control Rod Driv e Housing Supports 4.6-34 4.6.3.3.1 Testing an d Inspection 4.6-34 4.6.4 Information for Combined Performance of Reactivity Systems 4.6-34 4.6.4.1 Vulnerability to Common Mode Failures 4.6-34 4.6.4.2 Accidents Taking Credit for Two or More Reactivity Control Systems 4.6-34 4.6.5 Evaluation of Combined Performance 4.6-34 4.6.6 References 4.6-35 LSCS-UFSAR 4.0-xii REV. 20, APRIL 2014 CHAPTER 4.0 - REACTOR LIST OF TABLES NUMBER TITLE 4.2-1 Typical Limiting LHGR's for Gadolinia-Urania Fuel Rods (kW/ft) 4.2-2a GE Stress Intensity Limits 4.2-2b FANP Stress Intensity Limits 4.2-3 Conditions of Design Resulting from In-Reactor Process Conditions Combined with Earthquake Loading 4.2-4(a) Data for the 8X8R Fuel Design 4.2-4(b) Data for the GE 8x 8NB(GE 9B) Fuel Design 4.2-4(c) Data for the FANP ATRIUM-9B Fuel Design 4.2-4(d) Data for the FANP ATRIUM-10 Fuel Design 4.2-4(e) Data for the GE14 Fuel Design 4.2-5 Site Fuel Inspection Fu el Inspection Objectives 4.3-1 Maximum Incremental Rod Worths Using BPWS for Each of the Given Rod Groups 4.3-2 Neutron Fluxes Related to Vessel Irradiation 4.3-2a Bounding Neutron Flux and Fluences Related to Reactor Vessel Irradiation 4.3-3 24 Group Multigroup Neutron Flux at the Core Equivalent Radius 4.4-1 Thermal and Hydraulic Design Charac- teristics of the Reactor Core 4.4-2 Void Distribution 4.4-2a Axial Power Distribution Used to Generate Void and Quality Distributions (Typical) 4.4-3 Flow Quality Distribution (Typical) 4.4-4 Core Flow Distribution (Typical) 4.4-5 Typical Range of Test Data 4.4-6 Description of Uncertainties (Deleted) 4.4-7 Reactor Coolant System Geometrical Data 4.4-8 Lengths and Sizes of Safety Injection Lines 4.4-9 Bypass Flow Paths LSCS-UFSAR 4.0-xiii REV. 20, APRIL 2014 CHAPTER 4.0 - REACTOR LIST OF FIGURES AND DRAWINGS FIGURES NUMBER TITLE 4.1-1 Core Arrangement 4.1-2 Core Cell - GE 8X8R Fuel Type 4.1-2a Core Cell - GE 8X8NB Fuel Type 4.1-2b Core Cell - FANP ATRIUM-9B Fuel 4.1-2c Core Cell FANP ATRIUM-10 Fuel 4.1-2d GE14 Lattice Arrangement 4.1-2e GNF2 Lattice Arrangement 4.1-3 Fuel Assembly - (GE 8X8R Shown) 4.1-3a Fuel Assembly GE 8X8NB Fuel 4.1-3b Fuel Assembly FANP ATRIUM 9B Fuel 4.1-3c Fuel Assembly FANP ATRIUM-9B Fuel 4.1-3d Fuel Assembly FANP ATRIUM-10 Fuel 4.1-3e GE14 Fuel Bundle (Typical) 4.1-4 General Electric Control Rod Assembly 4.1-4a General Electric Original Equipment Control Rod Assembly 4.1-4b General Electric Typical Duralife 215 Control Rod Assembly 4.1-4c General Electric Typical Marathon Control Rod Assembly 4.1-5 Steam Separator 4.1-6 Steam Dryer 4.1-7 Steam Dryer Panel 4.2-1 Schematic of Four Bundle Cell Arrangement 4.2-2 Bypass Flow Paths 4.2-3 Fuel Bundle - 8X8R and BP8X8R Fuel Types 4.2-3a Fuel Bundle - GE 8X8EB Fuel Type 4.2-3b Fuel Bundle - GE 8X8NB Fuel Type 4.2-3c Fuel Bundle FANP ATRIUM 9B Type 4.2-3d Fuel Bundle FANP ATRIUM-10 Type 4.2-3e GE14 Fuel Bundle (Typical) 4.2-3f GNF2 Fuel Bundle (Typical) 4.2-4 [Deleted]
4.2-5 Control Rod Velocity Limiter 4.2-5a Fabricast Velocity Limiter 4.2-6 Typical Cladding Temperature vs. Heat Flux - BOL - 8X8R Fuel Type 4.2-7 Typical Cladding Temperature vs. Heat Flux - Late Life- 8X8R Fuel Type 4.2-8 Typical Energy Release as a Function of Time 4.3-1 Initial Core Loading Map (Deleted) 4.3-1a Unit 1 Cycle 5 Core Loading Map (Deleted)
LSCS-UFSAR 4.0-xiv REV. 20, APRIL 2014 FIGURES (Cont'd)
NUMBER TITLE 4.3-1b Unit 2 Cycle 5 Core Loading Map (Deleted) 4.3-2 K as a Function of Exposure at Various Void Fractions, High Enrichment, Dominant Fuel Type (Typical) 4.3-3 Atom Fraction as a F unction of Exposure, High Enrichment, Dominant Fuel Type, 40% Voids (Typical) 4.3-4 Fission Fraction as a Function of Exposure, High Enrichment, Dominant Fuel Type, 40% Voids (Typical) 4.3-5 Neutron Generation Time vs. Exposure at 40% Voids (Typical) 4.3-6 Delayed Neutron Fraction vs. Exposure at 40% Voids (Typical) 4.3-7 Variation of Maximum Rod Power as a Function of Exposure for High Enrichment, 40% Voids (Deleted) 4.3-8 Variation of Maximum Rod Power as a Function of Exposure (Deleted) 4.3-9 Variation of Bundle Average Maximum R-Factor as a Function of Bundle Average Exposure for Uncontrolled
High Enriched Bundle (Deleted) 4.3-10 Radial Power Factors (Deleted) 4.3-11 Typical Beginning of Cycle and End of Cycle Core Average Axial Power - 764 Core, BWR/4 and BWR/5 4.3-12 Moderator Void Reactivity C oefficient at EOC-1 Initial Cycle (GE) 4.3-13 Doppler Reactivity Coefficient as a Function of Fuel Exposure and Average Fuel Temperature at an Average Void Content of 40% High Enrichment Initial Cycle (GE) 4.3-14 Cold Shutdown Example of a Margin Curve 4.3-15 Control Rod Assignments for Groups 1 through 4 (Sequence A) 4.3-16 Control Rod Assignments for Groups 5 Through 10 (Sequence A) 4.3-17 Control Rod Assignments for Groups 1 Through 4 (Sequence B) 4.3-18 Control Rod Assignments for Groups 5 Through 10 (Sequence B) 4.3-19 Hot Operating EOC-1 Scram Reactivity 4.3-20 Xenon Reactivity Buildup and Burnout After Shutdown 4.3-21 Radial Power Distribution at 3323 MWt 4.3-21a Azimuthal Fast Flux Distribution 4.3-22 Axial Power Distribution at 3323 MWt 4.3-22a Axial Fast Flux Distribution 4.4-1 Damping Coefficient vs. Deca y Ratio (Second Order Systems)
LSCS-UFSAR 4.0-xv REV. 20, APRIL 2014 FIGURES (Cont'd)
NUMBER TITLE 4.4-2 Hydrodynamic and Core Stability Model 4.4-3 Model Block Diagram with Valve Flow Control 4.4-4 Comparison of Tests Results with Reactor Core Analysis (Deleted) 4.4-5 Core Reactivity Stability (End of Cycle) (Deleted) 4.4-6 10 psi Pressure Regulator Setpoint Step at 51.5% Rated Power (Natural Circulation) 4.4-7 10 Cent Rod Reactivity Step at 51.5% Rated Power (Natural Circulation) 4.4-8 6-inch Water Level Setpoint Step at 51.5% Rated Power (Natural Circulation) 4.4-9 10 psi Pressure Regulator Setpoint Step at 105% Rated Power and 100% Rated Flow 4.4-10 10 Cent Rod Reactivity Step at 105% Rated Power and 100% Rated Flow 4.4-11 10% Load Demand Step at 105% Rated Power and 100%
Rated Flow 4.4-12 6-inch Water Setpoint Step at 105% Rated Power and 100% Rated Flow 4.4-13 10 psi Pressure Regulator Setpoint Step at 68% Power and 50% Rated Flow 4.4-14 10 Cent Rod Reactivity Step at 68% Rated Power and 50% Rated Flow 4.4-15 10% Load Demand Step at 68% Rated Power and 50% Rated Flow 4.4-16 6-inch Water Level Setpoint Step at 68% Rated Power and 50% Rated Flow 4.6-1 Control Rod to Control Rod Drive Coupling 4.6-2 Control Rod Drive Unit 4.6-3 Control Rod Drive Unit (Schematic) 4.6-4 Control Rod Drive Unit (Cutaway) 4.6-5 Control Rod Drive Hydraulic System Process Diagram 4.6-6 Process Data, Control Rod Drive Hydraulic System 4.6-7 Control Rod Drive Hydraulic Control Unit 4.6-8 Control Rod Drive Housing Support LSCS-UFSAR 4.0-xvi REV. 14, APRIL 2002 DRAWINGS CITED IN THIS CHAPTER*
DRAWING* SUBJECT M-97 P&ID Reactor Water Cleanup System, Unit 1 M-100 P&ID Control Rod Drive Hydraulic Piping, Unit 1 M-143 P&ID Reactor Water Cleanup System, Unit 2 M-146 P&ID Control Rod Drive Hydraulic Piping, Unit 2
- The listed drawings are included as "Gen eral References" only; i.e., refer to the drawings to obtain additional detail or to obtain background information. These drawings are not part of the UFSAR. They are controlled by the Controlled Documents Program.
LSCS-UFSAR 4.2-1 REV. 20, APRIL 2014 4.2 FUEL SYSTEM
4.2.1 Design
Bases
This section and its subsections were written to describe the design basis consideration used in the design of th e GE initial core and reload fuel.
Detailed descriptions of the design basis co nsiderations used in the design of the AREVA reload fuel can be found in Reference 46 and 49. If a AREVA reference contains the equivalent information as what is being presented for GE, that reference is provided.
4.2.1.1 Safety Design Bases The fuel assembly is designed to ensure, in conjunction with the core nuclear characteristics (Section 4.3), the core thermal and hydraulic characteristics (Section 4.4), the plant equipment characteristics and the instrumentation and protection system, that fuel damage does not result in the release of radioactive materials in excess of the guideline values of 10 CFR 20, 50, and 100.
The mechanical design process emphasizes that:
- a. the fuel assembly provides substantial fission product retention capability during all potential operational modes, and
- b. the fuel assembly provides sufficient structural integrity to prevent operational impairment of any reactor safety equipment.
Assurance of the design basis considerations is provid ed by the following fuel assembly capabilities:
- a. Pressure and temperature capabilities The fuel assembly and its components are capable of withstanding the predicted thermal, pressure, and mechanical interaction loadings occurring during startup testing, normal operation, and abnormal operation without impairment of operational capability.
LSCS-UFSAR 4.2-2 REV.
13 b. Handling capability The fuel assembly and each component thereof is capable of withstanding loading predicted to occur during handling without impairment of operational capability.
- c. Earthquake loading capability (OBE)
The fuel assembly and each component thereof is capable of sustaining incore loading predicted to occur from an operating basis earthquake (OBE), when occurring during normal operating conditions without impairment of operational
capability.
- d. Earthquake loading capability (SSE)
The fuel assembly and each component thereof is capable of sustaining incore loading predicted to occur from a safe
shutdown earthquake (SSE) when occurring during normal operation without:
- 1. exceeding deflection limits which allow control rod insertion, and
- 2. fragmentation or severance of any bundle component.
- e. Accident capability The capability of the fuel assembly to withstand the control rod drop accident, the pipe breaks inside and outside containment accidents, the fuel handling accident, and one recirculation pump seizure accident, is determined by analysis of the specific event. The ability of the fuel assemb ly to provide the preceding capabilities is evaluated by one or more of the following:
- a. design ratios developed by utilizing continually evolving, state-of-the-art numerical analysis techniques (Subsection 4.2.1.2.15);
- b. analytical procedures based on classical methods (Subsection 4.2.1.2.5); and
- c. experience and testing (Subsection 4.2.3.2).
LSCS-UFSAR 4.2-3 REV. 20, APRIL 2014 For the initial reloads of the AREVA ATRIUM-9B and ATRIUM-10 fuel, the control rod drop accident, the pipe breaks inside and outside containment accidents, and the fuel handling accident were all evaluated for the fuel assembly's capability to withstand their effects. The recirculation pump seizure event was not analyzed by AREVA for their ATRIUM-9B or ATRIUM-10 fuel; it was dispositioned as bounded by the LOCA accident. The control rod drop accident is evaluated each cycle for both AREVA and GE reloads.
4.2.1.2 Power Generation Design Basis
The fuel assembly is designed to ensure, in conjunction with the core nuclear characteristics, the core thermal and hydr aulic characteristics, the plant equipment characteristics and the instrumentation and protection system, that fuel damage limits will not be exceeded during either planned operation or abnormal operational transients caused by any single equipment malfunction or single operator error.
4.2.1.2.1 Material Selection
The basic materials used in fuel assemb lies are Zircaloy, natural zirconium, Type 304 stainless steel, Inconel-X, and ceramic uranium dioxide and gadolinia.
These materials have been shown from earlier reactor experience to be compatible with BWR conditions and to retain their design function capability during reactor operation. Additional information on material properties is referenced in Reference 41.
4.2.1.2.2 Effects of Irra diation and Fuel Swelling Irradiation affects both fuel and cladding material properties. The effects include increased cladding strength and reduced cladding ductility. In addition, irradiation in a thermal reactor environment results in the buildup of both gaseous and solid fission products within the UO 2 fuel pellet which tend to increase the pellet diameter, i.e., fuel irradiation swelling. Pellet internal porosity and pellet-to-cladding gap have been specifie d such that the thermal expansion and irradiation swelling are accommodated throughout life. The irradiation swelling model is based on the NRC approved methodology as described in Reference 41.
Observations and calculations based on this refined model for relative UO 2 fuel/cladding expansion indicate that the as-fabricated UO 2 pellet porosity is adequate (without pellet dishing) to accommodate the fission-product-induced UO 2 swelling out to expected exposures.
The primary purpose of the gap between the UO 2 fuel pellet and Zircaloy cladding is to accommodate differential diametral expansion of fuel pellet and cladding and, thus, preclude the occurrence of excessive gross diametral cladding strain. A short time after reactor startup, the fuel cracks radially and redistributes out to the LSCS-UFSAR 4.2-4 REV. 20, APRIL 2014 cladding. Experience has shown, however, the gap volume remains available in the form of radial cracks to accommodate gross diametral fuel expansion.
The value of thermal conductance used in BWR fuel design is derived from postirradiation data on exposed fuel with an initial pellet-to-cladding gap which is significantly larger than that employed in the General Electric fuel design.
Axial ratcheting of fuel cladding is no t considered in BWR fuel rod design.
Prototypical fuel rods have been operated in the Halden test reactor with axial elongation transducers. No significan t axial ratcheting has been observed (Reference 5).
Fission product buildup also tends to ca use a slight reduction in fuel melting temperature. The melting point of UO 2 is considered to decrease with irradiation based on data from Reference 6.
In the temperature range of interest (500 C), the fuel thermal conductivity is not considered to be significantly affected by irradiation as reported in Reference 7.
A small fraction of the gaseous fission products is released from the fuel pellets to produce an increase in fuel rod internal gas pressure as discussed further in Subsection 4.2.1.2.7. In general, such irradiation effects on fuel performance have been characterized by available data and are considered in determining the design features and performance. Thus, the irradiation effects on fuel performance are inherently considered when determining whether or not the stress intensity limits and temperature limits are satisfied.
In Reference 49, AREVA states that the BWR evaluation models for densification and swelling are included in the NRC approved fuel performance codes, References 50 and 51.
4.2.1.2.3 Fuel Densification
4.2.1.2.3.1 GE Fuel Fuel performance calculations that account for some specific effects of fuel densification have been performed with an approved version of the General Electric analytical model as described in Reference 41. The approved analytical model incorporates time-dependent fuel densification, time-dependent gap closure and
cladding creepdown for the calculation of gap conductance. Other fuel performance predictions, such as cladding response, ar e also calculated. Cladding collapse has not been observed in boiling water reactor fuel rods, but its theoretical occurrence is calculated with the NRC approved methodolo gy as described in Reference 41. All of the fuel cladding used at LSCS has been shown not to collapse during the life of the fuel.
LSCS-UFSAR 4.2-5 REV. 20, APRIL 2014 4.2.1.2.3.2 AREVA Fuel
Fuel densification and swelling are limited by design criteria specified for fuel temperature, cladding strain, cladding collapse, and internal pressure criteria (Reference 49).
Creep collapse of the cladding and the subsequent potential for fuel failure is avoided in the AREVA fuel system design by eliminating the formation of axial gaps. The maximum cladding circumferential creep and ovalization consistent with the time of maximum densification is computed during a creep collapse evaluation to demonstrate that no axial gaps are present. The evaluation must show that the pellet column is compact at the burnup of maximum densification (approximately 6000 MWd/MTU). The internal plenum spring provides an axial load on the fuel stack that is sufficient to assist in the closure of any gaps caused by handling, shipping, and densification. Evaluation of cladding creep stability in the unsupported condition is performed cons idering the compressive load on the cladding due to the difference between primary system pressure and the fuel rod internal pressure. AREVA fuel is designed to minimize the potential for the formation of axial gaps in the fuel and to minimize clad creepdown which would
prevent the closure of axial gaps or allow creep collapse (Reference 49).
4.2.1.2.4 Incipient UO 2 Center Melting
4.2.1.2.4.1 GE Fuel
The fuel rod is evaluated to ensure that fuel rod failure due to fuel melting is not expected to occur during normal steady-state operation. Incipient center melting is not expected to occur in fresh GE UO 2 fuel rods at the linear heat generation rate (LHGR) described in Reference 41 and Reference 59. The LHGR values for incipient center melt decrease slightly with burnup. The effect of gadolinia concentration and fuel exposure on the LHGR at calculated incipient center melting is also described in Reference 41.
4.2.1.2.4.2 AREVA Fuel
Fuel failure from the overheating of the fuel pellets is not allowed. The centerline temperature of the fuel pellets must rema in below melting during normal operation and anticipated operational occurrences.
The melting point of the fuel includes adjustments for burnup and gadolinia co ntent. AREVA establishes steady state and transient design LHGR limits for ea ch fuel type which protect against centerline melting. These LHGR limits ar e appropriate for normal operation and anticipated operational occurrences throug hout the design lifetime of the fuel (Reference 49).
LSCS-UFSAR 4.2-6 REV. 18, APRIL 2010 4.2.1.2.5 Maximum Allowable Stresses The strength theory, terminology, and stress categories presented in the ASME Boiler and Pressure Vessel Code,Section III, are used as a guide in the mechanical
design and stress analysis of the reactor fuel rods. The mechanical design is based on the maximum shear stress theory for combined stresses. The equivalent stress intensities used are defined as the difference between the most positive and least
positive principal stresses in a triaxial field. Thus, stress intensities are directly comparable to strength values found from tensile tests. Table 4.2-2a and b present a summary of the basic stress intensity limits that are applied for Zircaloy-2 cladding for both GE fuel and AREVA ATRIUM-9B and ATRIUM-10 fuel.
4.2.1.2.5.1 GE Fuel In this analysis of BWR Zircaloy-clad UO 2 pellet fuel, continuous functional variations of mechanical properties with exposure are not employed since the irradiation effects become saturated at very low exposure. At beginning of life, the cladding mechanical properties employ ed are the unirradiated values. At subsequent times in life, the cladding mechanical properties employed are the
saturated irradiated values. The only exception to this is that unirradiated mechanical properties are employed above the temperatures for which irradiation effects on cladding mechanical properties are assumed to be annealed out. It is significant that the values of clad yield strength and ultimate tensile strength employed represent the approximate lower bound of data on cladding fabricated by General Electric, i.e., approximately two standard deviations below the mean value.
In this analysis the calculated stress and the yield strength or ultimate strength are combined into a dimensionless quantity called the design ratio. This quantity is the ratio of calculated stress intensity to the design stress limit for a particular stress category. The design stress limit for a particular stress category is defined as a fraction of either the yield strength or ultimate strength, whichever is lower. Thus, the design ratio is a measure of the fracti on of the allowable stress represented by the calculated stress.
Analyses are performed to show that the stress intensity limits given in Table 4.2-2a and b are not exceeded during continuous operation with linear heat generation rates up to the design operating limit, or during transient operation above the design operating limit. Stresses due to external coolant pressure, internal gas pressure, thermal effects, spac er contact, flow-induced vibration, and manufacturing tolerances are considered. Cladding mechanical properties used in stress analyses are based on test data of fuel rod cladding for the applicable temperature.
Fuel rods are evaluated to assure that the fuel will not fail due to stresses or strains exceeding the fuel rod mechanical capability. The analysis performed is described in Reference 41.
LSCS-UFSAR 4.2-7 REV. 20, APRIL 2014 4.2.1.2.5.2 AREVA Fuel AREVA requires compliance with both Standard Review Plan criteria for pellet/cladding interaction for steady state and transient conditions over the lifetime of the fuel. The first one is that transient - induced deformations must be less than 1% uniform cladding strain. The second is that fuel melting cannot occur. Compliance with the fuel melting criteria is discussed in Section 4.2.1.2.4.2.
The design basis for the fuel cladding stress limits is that the fuel system will not be damaged due to fuel cladding stresses. Conservative limits are derived from the ASME Boiler Code,Section III, Article-2000; and the specified 0.2% offset yield strength and ultimate strength for Zircaloy (Reference 49).
4.2.1.2.6 Capacity for Fission Gas Inventory
The available fission gas retention volume is determined based upon the following assumptions:
- a. Nominal as-built plenum length and cladding inside diameter.
- b. Maximum expected fuel-cladding differential expansion.
- c. No credit for fuel-cladding annulus (gap).
- d. The "net" volume is corrected for the volume of the components contained within the fuel rod plenum.
4.2.1.2.7 Maximum Internal Gas Pressure
Fuel rod internal pressure is due to the helium which is backfilled during rod fabrication, the volatile content of the UO 2 , and the fraction of gaseous fission products which are released from the UO
- 2. Nominal tolerances are assumed in defining the hot plenum volume used to compute fuel rod internal gas pressure.
4.2.1.2.7.1 GE Fuel The fuel rod internal pressure is calculat ed using the perfect ga s law (P = NRT/V).
A quantity of 1.35 milligram-moles of fission gas is pr oduced per MWd of power production. In fuel rod pressure and stress calculations, fission gas release is calculated as per the NRC approved meth odology as described in Reference 41. This fission gas release model has been demonstrated by experiment to be conservative over the complete range of design temperature and exposure conditions (References 4 and 41). The calculated maximum fission gas release fraction in the highest design power density rod is less than 25%. This calculation is conservative because it assumes the worst peaking LSCS-UFSAR 4.2-8 REV. 18, APRIL 2010 factors applied constantly to this rod. The percentage of total fuel rod radioactivity released to the rod plenum is much less than 25% because of radioactive decay during diffusion from the UO
- 2.
Creepdown and creep collapse of the plenum are not considered because significant creep in the plenum region is not expected. The fuel rod is designed to be free-standing throughout its lifetime. The temperature and neutron flux in the plenum region are considerably lower than in the fueled region, thus the margin to creep collapse is substantially greater in the plenum. Direct measurements of irradiated fuel rods have given no indication of significant creepdown of the plenum.
The fuel rod is evaluated to assure that the effects of rod internal pressure during normal steady state operation will not result in fuel failure. The analysis is further described in Reference 41.
4.2.1.2.7.2 AREVA Fuel
To prevent unstable thermal behavior and to maintain the integrity of the cladding, AREVA limits the maximum internal rod pressure relative to system pressure to avoid significant hydride reorientation during cooldown conditions or depressurization conditions. When the fu el rod internal pressure exceeds system pressure, the pellet-cladding gap has to remain closed if it is already closed or it should not tend to open for steady or increasing power conditions. Outward circumferential creep which may cause an increase in pellet-to-cladding gap must be prevented since it would lead to higher fuel temperature and higher fission gas release. The maximum internal pressure is also limited to protect embrittlement of
the cladding caused by hydride reorientation during cooldown and depressurization conditions (Reference 49).
4.2.1.2.8 Internal Pressure and Cladding Stresses During Normal Conditions
The internal pressure is applied coincident with the applicable coolant pressure to compute the resulting cladding stresses, which, combined with cladding stresses from other sources, must satisfy the stress limits described in Subsection 4.2.1.2.5.
4.2.1.2.9 Cycling and Fatigue Limits
4.2.1.2.9.1 GE Analysis
The fatigue analysis utilizes the linear cumulative damage rule (Miner's hypothesis) as documented in "Fatigue Design Basis for Zircaloy Components" (Reference 12). The fatigue analysis is based on the estimated number of temperature, pressure, and power cycles. Th e fuel assembly and fuel rod cladding are evaluated to ensure that strains due to cyclic loadings will not exceed the fatigue capability.
LSCS-UFSAR 4.2-9 REV. 20, APRIL 2014 4.2.1.2.9.2 AREVA Analysis Cycle loading associated with relatively large changes in power can cause cumulative damage which may eventually lead to fatigue failure. Therefore, AREVA requires that the cladding not exceed a cumulative fatigue usage factor of 0.67. The O'Donnell and Langer fatigue curves are used in the analysis. These fatigue curves have been adjusted to incorporate the recommended '2 or 20' safety factor. This safety factor reduces the stress amplitude by factor of 2 or reduces the number of cycles by a factor of 20, whichever is more conservative. The fatigue curves provide the maximum allowed number of cyclic loading for each stress amplitude. The fatigue usage factor is the number of expected cycles divided by the number of allowed cycles. The total cladding usage factor is the sum of the individual usage factors for each duty cycle (Reference 49).
4.2.1.2.10 Deflections
4.2.1.2.10.1 GE Evaluation
The operational fuel rod deflections considered are the deflections due to:
- a. manufacturing tolerances, b. flow-induced vibration, c. thermal effects, and
- d. axial load.
There are two criteria that limit the magnitud e of these deflections.
One criterion is that the cladding stress limits must be satisfied; the other is that the fuel rod-to-rod and rod-to-channel clearances must be suffi cient to allow free passage of coolant water to all heat transfer surfaces. The fuel rod is evaluated to ensure that fuel rod bowing does not result in fuel failure due to boiling transition.
4.2.1.2.10.2 AREVA Evaluation
Differential expansion between the fuel rods, and lateral thermal and flux gradients can lead to lateral creep bow of the rods in the spans between sp acer grids. This lateral creep bow alters the pitch between the rods and may affect the peaking and
local heat transfer. The AREVA design basis for fuel rod bowing is that lateral displacement of the fuel rods shall not be of sufficient magnitude to impact thermal margins. Extensive post-irradiation exam inations have confirmed that such rod LSCS-UFSAR 4.2-10 REV. 18, APRIL 2010 bow has not reduced spacing between adjacent rods by more than 50%. The potential effect of this bow on thermal margins is negligible. Rod bow at extended burnup does not affect thermal margins du e to the lower power achieved at high exposure (Reference 49).
4.2.1.2.11 Flow-Induced Fuel Rod Vibrations Flow-induced fuel rod vibrations depend primarily on flow velocity and fuel rod geometry. The stress levels resulting from the vibrations are negligibly low and well below the endurance limit of all affected components. This phenomenon is further described in GE References 13 and 41.
Reference 47 discusses the AREVA calculatio ns for flow induced vibrations. Vibrational stresses due to flow induce d vibrations are calculated with the Paidoussis analysis which assumes:
- 1) The structural stiffness of the rod is due to cladding only.
- 2) The sections of the fuel rod be tween spacers and/or tie plate supports are modelled structurally as a simple beam with pinned ends.
- 3) Flow velocity, viscosity, and virtual mass for the amplitude calculations are evaluated as suggested by Paidoussis.
4.2.1.2.12 Fretting Corrosion
Fretting wear has been considered in establishing the fuel mechanical design basis.
Specific GE fuel designs described in Reference 41 have been incorporated to eliminate fretting wear. Tests of these designs have been conducted both
out-of-reactor as well as in-reactor prior to application in a complete reactor core basis. All tests and post-irradiation exam inations have indicated that fretting corrosion does not occur. Post-irradiation examination of many fuel rods indicates only minor fretting wear. Excessive wear at spacer contact points has never been observed with the current spacer configuration. Additional information on testing relative to fretting wear is contained in Reference 41. AREVA discusses fretting wear in Reference 49.
4.2.1.2.13 Seismic Loadings
The fuel is analyzed for loading in the reactor resulting from seismic accelerations. The fuel seismic design basis is the design basis presented in References 15, 17 and 41 for GE fuel. The fuel seismic design basis for AREVA fuel is presented in Reference 49. Reference 48 verifies that the AREVA seismic criterial were met for a particular reload.
LSCS-UFSAR 4.2-11 REV. 18, APRIL 2010 4.2.1.2.14 Chemical Properties of Cladding and Fuel Material
The fuel material, fuel rod, pellets, and cladding are discussed generally in Subsections 4.2.2.2 through 4.
2.2.5. Testing
and inspection of fuel is covered in Subsection 4.2.4. Reference 41 reports the specific fuel parameters of the fuel used for LSCS. Reference 19 presents the BWR fuel experience through September 1974.
Reference 42 represents later BWR fuel ex perience. References 46, 48, and 49 report specific fuel parameters for AREVA fuel.
4.2.1.2.15 Design Ratios
Design ratios are defined by the following relationship: D.R. = A/L where D.R. is the design ratio, L is the limiting parameter value, and A is the actual parameter value. Design ratios of less than one are demonstrated for component parameters influenced by loading conditions which may affect the structural or dimensional integrity of the fuel assemb ly or any component thereof.
4.2.1.2.15.1 Limiting Parameter Values The following information is based on GE methodology. For a discussion on AREVA methodologies see Reference 46, 47, 48, 49 and 55.
4.2.1.2.15.1.1 Normal and Upset Design Conditions
Limiting parameter values for each component are determined in the following manner as defined by Table 4.2-3:
- a. For stress resulting from mean value or steady-state loading, the limiting value is determined by consideration of the material
0.2% offset yield strength or the equivalent strain, as established at operating temperature.
- b. For stress resulting from load cycling, limiting parameter values are determined from fatigue limits.
- c. For stress resulting from loading of significant duration, the limiting parameter is determined from consideration of stress rupture as defined by the Larson-Miller parameter. If metal temperatures are below the level of applicability of stress rupture for the material or if the yield strength is more limiting then the limiting value of stress is determined from consideration of the material 0.2% offset yield strength or the equivalent strain, as establis hed at operating temperatures.
LSCS-UFSAR 4.2-12 REV. 15, APRIL 2004
- d. Where stress rupture and fatigue cycling are both significant, the following limiting condition is applied: I = 1 to n I = 1 to m
- e. Critical instability loads shall be derived from test data when available or from analytical methods when applicable test data are not available.
- f. Deflection limits are those values of component deformation which could cause an undesirable event such as impairment of control rod movement or an excessive leakage flow rate.
4.2.1.2.15.1.2 Emergency and Faulted Design Conditions Limiting parameter values are determined in the following manner as defined by Table 4.2-3:
- a. Stress limits are determined from consideration of the ultimate tensile strength or equivalent strain of the material, as
established at operating temperatures.
- b. Critical instability loads are determined from test data when available or from analytical methods when applicable test data is not available.
- c. Deflection limits are those values of deformation that if occurring could lead to a more serious consequence such as
prevention of control rod insertion.
4.2.1.2.15.2 Actual Parameter Values
The following information is based on GE methodology. For a discussion on FANP methodologies see References 46, 47, 48, 49 and 55.
Actual parameter values are determined from the following considerations:
- a. Effective stresses are determined at each point of interest using the theory of constant elastic strain energy of distortion:
1stressat cycles allowable cycles ofnumber actual stressat time allowable stressat timeactual LSCS-UFSAR 4.2-13 REV. 14, APRIL 2002 Stress concentration may be applied only to the alternating stress component.
- b. Design values of instability loads are scaled up to allow for uncertainty in manner of load application, variation in modulus of elasticity, and difference between the actual case and the
theoretical one.
- c. Calculated values of deflection for comparison with deflection limits may be based on the resulting permanent set after load removal if load removal occurs before damage may result.
4.2.1.2.16 Fuel Assembly Limits The design limits applicable to each component are discussed in the following paragraphs. In order to provide a fulle r understanding of how the limits will be applied, a functional description of each component and a discussion of the loadings on each component are provided.
The general configuration of the fuel asse mbly and the detailed configurations of the assembly components are the result of the evolutionary change in customer, performance, manufacturing and serviceability requirements and the experience obtained since the initial design conception. In general, the experience obtained in prior fuel designs is relied upon very heavily to qualify particular component configurations for production fuel applic ation. More sophisticated analytical techniques are continually being developed and applied to fuel design.
4.2.1.2.16.1 Fuel Rods
A discussion of the mechanical analysis of the fuel rod and the appropriate stress intensity limits was provided in Subsection 4.2.1.2.5. In additi on, a fuel rod fatigue analysis is performed as descri bed in Subsection 4.2.1.2.9.
As explained in Subsection 4.2.3.21, significant fuel rod bowing due to binding at the spacers is not expected to occur. Other contributors to rod bowing during normal operation and transients are manufacturing tolerances and thermal gradients. These factors are considered in the design.
4.2.1.2.16.2 Fuel Spacer
The primary function of the fuel spacer is to provide lateral support and spacing of the fuel rods, with consideration of thermal-hydraulic performance, fretting wear, strength, neutron economy, and producibility.
The mechanical loadings on the spacer structure during normal operation and transients result from the rod positioning spacer spring forces and from local LSCS-UFSAR 4.2-14 REV. 18, APRIL 2010 loadings at the water rod-spacer positioning device. During a seismic event, the spacer transmits the lateral acceleration loadings from the fuel rods into the channel, while maintaining the spatial relationship between the rods.
As noted above, the spacer represents an optimization of a number of considerations. Thermal-hydraulic development effort has gone into designing the particular configuration of the spacer parts. The resu ltant configurations give enhanced hydraulic performance. Extensive flow testing has been performed employing prototypical spacers to define single-phase and two-phase flow characteristics. Details of the mechanical design of the spacers used at LSCS can be found in Reference 41 for GE fuel and References 46, 47, 48 and 55 for AREVA fuel.
4.2.1.2.16.3 Water Rods or Water Channel The main mechanical function of the water ro d(s) (or water channel) is to maintain the axial position of the fuel spacers. For the ATRIUM-10 Fuel, the water channel also provides the structural connection between the upper and lower tie plates.
Differential thermal expansion between fuel ro ds and the water rods (or water channel) can introduce axial loadings into the water rod (or water channel) through the frictional forces between the fuel rods and the spacers. This differential growth is considered in the design process as discussed in Referenc e 41 for GE fuel and References 46 and 55
for AREVA fuel.
The water rods or water channel provide flow through the center portion of the fuel assembly, thereby providing additional modera tion within the bundle interior. This improves uranium utilization and operational flexibility.
4.2.1.2.16.4 Channel Assurance that the channels maintain their dimensional integrity, strength, and spatial position throughout their lifetime is provided in the following ways:
- a. Dimensional integrity, as related to relaxation of residual forming stresses, is provided through the channel specifications and by qualification of the manufacturing process to these specifications.
The operational experience with channels produced using the current process has demonstrated satisfactory relaxation characteristics (Reference 17).
- b. The performance of the channels currently in operation has shown no tendency for gross inservice deformations, although long-term creep deformation and channel bulg e have been identified as a potential life-limiting phenomenon (References 17 and 45).
- c. Channel material strength is assured through the material specification of yield and ultimate strength. Quality LSCS-UFSAR 4.2-15 REV. 18, APRIL 2010 measurements are made to show compliance with this specification. Irradiation subs tantially increases the material strength.
- d. Mechanical integrity of the channel (that is, assurance that the channel will maintain its spatial position and integrity) is provided by designing the channel to the limits stated in Subsection 4.2.2.6 and item e following. The design limits used are based on the unirradiated strength of the material, thereby providing substantial material strength margin throughout most
of the life of the channel.
- e. During normal and transient operation, the channel is subjected to differential pressure loadings. The pressure loadings are evaluated to ensure the channel will not experience excessive deflection and subsequent channel wear.
4.2.1.2.16.5 Tie Plates
The upper and lower tie plates serve the functions of supporting the weight of the fuel and positioning the rod ends during all phases of operation and handling. The loading on the lower tie plate during operation and transients comprise the fuel weight, the weight of the channel, and the forces from the expansion springs at the top of the fuel rods. The loading of the upper tie plate is the expansion springs' force. The expansion springs permit differential expansion between the fuel rods without introducing high axial forces into the rods.
Most of this loading arises from the weight of the fuel rods and the channel, which are not cyclic loadings. During accidents the tie plates are subjected to the normal operational loads plus the blowdown and seismic loadings. During handling, the tie plates are subjected to acceleration and impact loading. The stress design limit for the tie plates for all phases of operat ion and normal handling is discussed in Reference 41 for GE fuel. Reference 48 contains information regarding the upper and lower tie plate loads for AREVA ATRIUM-9B and ATRIUM-10 fuel.
The ATRIUM-10 fuel design includes AREVA's FUELGUARD debris resistant lower tie plate. This design was chosen for two reasons: 1) to address the main cause of BWR fuel failures over the past few years - de bris induced fuel rod fretting; and 2) to reduce overall assembly pressure drop to be tter assure adequate core flow is available for reactor power maneuvering.
The FUELGUARD lower tie plate design consis ts of a parallel array of blades with curved portions in the middle. The blades are arranged so that there is no line of sight through the grid thus preventing the passage of long narrow objects and objects larger than the pitch of the blades. The blades for the FUELGUARD on the
ATRIUM-10 are brazed in position.
LSCS-UFSAR 4.2-15a REV. 20, APRIL 2014The GE14 fuel design is assembled with a debris filter Lower Tie Plate (LTP) as standard equipment. The debris filter LTP increases the single phase pressure drop by approximately 0.3 psi over the non-debris filter LTP. The debris filter LTP has an underlying grid that screens out the debris and mitigates the debris related fuel rod failures by reducing the size of debris that can enter the fuel assembly. More detailed description of the debris filt er LTP can be found in Reference 58.
The GNF2 fuel design is assembled with the Defender LTP as standard equipment.
This LTP provides essentially the same si ngle-phase pressure drop as the GE14 debris filter LTP, while preventing smaller debris from entering the fuel assembly.
Additional description of the Defender LTP can be found in Reference 59.
4.2.1.2.17 Reactivity Control Assembly and Burnable Poison Rods 4.2.1.2.17.1 Safety Design Bases for Reactivity Control
The limiting criteria for shutdown reactivity margins are given in Subsection 4.3.1.1 as items a and f. The cold-clean shutdown margin is shown in Figure 4.3-14 for the initial cycle of Units 1 and 2. Th e presence of the burnable poison Gd 2 O 3 is apparent in the curve shape as keff rises concurrent with poison depletion. The negative reactivity worth of the gadolinia-containing fuel rods decreases in a nearly linear manner so that it closely matches the depletion of fissile material. The curve LSCS-UFSAR 4.2-16 REV. 14, APRIL 2002 shown in Figure 4.3-14 is typical for most cycles, although differences will exist from cycle to cycle.
The reactivity control mechanical design includes control rods and gadolinia burnable poison in selected fuel rods within fuel asse mblies and meets the following safety design bases.
- a. The control rods have sufficie nt mechanical strength to prevent displacement of their reactivity control material.
- b. The control rods have sufficient strength and are so designed as to prevent deformation that could inhibit their motion.
- c. Each control rod has a device to limit its free-fall velocity sufficiently to avoid damage to the nuclear system process barrier by the rapid reactivity increase resulting from a free-fall of one control rod from its fully inserted position to the position where the drive was withdrawn.
4.2.1.2.17.1.1 Specific Design Characteristics The acceptability of the control rod and control rod drive under scram loading condition is demonstrated by functional testing instead of analysis or adherence to formally defined stress limits. The results of such testing are given in Reference 10.
The basis of the mechanical design of the control rod blade clearances is that there is no interference which will restrict the passage of the control rod blade.
Mechanical insertion requirements during normal operation are selected to provide adequate operability and load following capability, and are able to control the reactivity addition resulting from burnou t of peak shutdown xenon at 100% power.
Scram insertion requirements are chosen to provide sufficient shutdown margin to meet all safety criteria for plant operational transients (Chapter 15.0).
The selection of materials for use in the control rod design is based upon their in-reactor properties. The irradiated properties of Type 304 austenitic stainless
steel, 316 stainless steel and CF3 which compr ise the major portion of the assembly, B 4 C powder, hafnium Inconel-X, and stellite are well known and are taken into account in establishing the mechanical design of the control rod components. The basic cruciform control rod design and materials have been operating successfully in all GE reactors. No problems associat ed with component materials have been observed.
The radiation effects on B 4C powder include the release of gaseous products, and the B 4C cladding is designed to sustain the resulting internal pressure buildup. The corrosion rate and the physical properties, e.g., density, modulus of elasticity, LSCS-UFSAR 4.2-17 REV. 14, APRIL 2002dimensional aspects, etc., of austenitic stainless steel, 316 stainless steel, CF3 and Inconel-X are essentially unaffected by the irradiation experienced in the BWR reactor core. The effects upon the mech anical properties, i.e., yield strength, ultimate tensile strength, percent elon gation, and ductility on the Type 304 stainless steel cladding also are well known and are considered in mechanical design. Visual examinations of control rods whic h have been subjected to high exposure rates have disclosed no significant material degradation (Reference 11).
Rod positioning increments (notch lengths) are selected to provide adequate power shaping capability. The combination of rod speed and notch length must also meet the limiting reactivity addition rate criteria.
For all LaSalle cores, supplementary reactivity control must be provided in such a way that the high initial keff can be compensated throughout the active core. Gadolinia containing fuel rods are used in normal fuel assemblies to attain this objective. Some assemblies contain mo re gadolinia than others to improve flattening both in the radial and axial directions.
The gadolinia is uniformly distributed in the UO 2 pellet and forms a solid solution.
The presence of the high cross section gadolinium isotopes results in a relatively low heat generation rate in those rods (this heat generation rate is also adjusted by the position of the gadolinia rods within the fu el assembly). During a fuel cycle, the gadolinia essentially burns out thus enab ling a progressive increase in rod power and a concurrent increase in net assembly power. At later stages of fuel exposure the power of the gadolinia-urania fuel rods decreases.
Precise quality control measures are utilized during the manufacture of gadolinia bearing UO 2 pellets and also during the assembly of these fuel pellets into fuel rods. Special procedures assure accurate placem ent and quantity control for placement of gadolinia rods.
4.2.1.2.18 Surveillance Program See Subsection 4.6.3.2 for information regarding the control rod surveillance program. The surveillance tests for the control rod drive system include an acceptance test, preinstallation test, operational test prior to startup, and tests during startup. Specific surveillance tests are performed following a refueling outage when core alterations are made, to demonstrate that the core can be made subcritical with a margin of 0.0038 k at any time in the subsequent fuel cycle with the strongest operable control rod fully withdrawn and all other operable control rods fully inserted.
LSCS-UFSAR 4.2-18 REV. 20, APRIL 20144.2.2 Description and Design Drawings 4.2.2.1 Core Cell
A core cell consists of a control rod and the four fuel assemblies which immediately surround it (Figure 4.2-1). Each core cell is associated with a four-lobed fuel support piece. Around the outer edge of the core, certain fuel assemblies are not immediately adjacent to a control rod and are supported by individual peripheral fuel support pieces.
The top guide is an "egg-crate" structur e of stainless steel bars which form a four-bundle cell. The four fuel assemblies are lowered into this cell and, when seated, springs mounted at the tops of th e channels force the channels into the corners of the cell such that the sides of the channels contact the grid beams (Figure 4.2-1).
4.2.2.2 Fuel Assembly
A fuel assembly consists of fuel bundle and the channel whic h surrounds it (Figure 4.1-3). The fuel assemblies are arranged in the reactor core to approximate a right circular cylinder inside the core shroud. Each fuel assembly is supported by a fuel support piece and the top guide. A summary of nuclear fuel data for the GNF (formerly GE) 8x8R, 8x8NB, GE14, and GNF2 fuel designs are presented in Tables 4.2-4a, 4.2-4b, 4.2-4e, and 4.2-4f, respectively. A summary of nuclear fuel data for the AREVA ATRIUM-9B and ATRIUM-10 fuel designs are presented in Tables 4.2-4c and 4.2-4d, respectively. Other pertinent data are presented in References 41, 44, 46, 48, 49, 58, and 59.
Beginning with LaSalle Unit 2 Cycle 13 and co ntinuing in subsequent cycles of Unit 2, AREVA Lead Fuel Assemblies (LFAs) ar e inserted into non-limiting core locations for demonstration purposes. The LFA program consists of eight Lead Fuel Assemblies of ATRIUM 10XM.
The ATRIUM 10XM fuel bundle shares a ge neral geometry with the ATRIUM 10, consisting of a 10x10 fuel rod lattice with a square internal water channel.
Variations from ATRIUM 10 include water channel crowns, a change in spacer material, addition of one spacer for a total of nine, larger fuel rod diameter, longer active fuel length, additional uranium mass, and in selected fuel rods doped fuel pellets and re-crystallized cladding. ATRIUM 10XM is described in Mechanical Design Report for LaSalle Unit 2 Cycle 13 Atrium 10XM Lead Test Assemblies , ANP-2756P, Revision 0, October 2008.
LSCS-UFSAR 4.2-18a REV. 20, APRIL 20144.2.2.3 Fuel Bundle The 8x8R and BP8x8R (Figure 4.2-3) fuel bundles contain 62 fuel rods and two water rods which are spaced and supported in a square (8 x 8) array by the lower and upper tie plates. The GE8x8EB fuel de sign (Figure 4.2-3a) provides for the use of up to four water rods. However, the GE 8X8EB fuel bundles at LaSalle have two water rods. The GE8X8NB fuel design (Fig ure 4.2-3b) contains 60 fuel rods and one large centrally located water rod.
The GE14 fuel design (Figure 4.2-3e) is bas ed on a 10x10 array that contains 78 full length rods, 14 part length rods and 2 large water rods that effectively replaced 8 fuel rods. The 14 part length rods terminate just past the top of the fifth spacer. Eight full length rods are used as tie rods. The rods are spaced and supported by the upper and debris filter lower tie plates and eight spacers over the length of the fuel rods. This assembly is encased in an interactive thick corner/thin wall fuel channel. Finger springs control the coolant leakage flow between the debris filter lower tie plate and the channel.
Additional assembly and component description for the GE14 fuel design are provided in Reference 58.
The GNF2 design consists of 92 fuel rods and two large central water rods contained in a 10x10 array. The two water rods encompass eight fuel rod positions. Eight of the fuel rods terminate at approximately two-thirds of the bundle length and are designated as long part length fuel rods.
Six fuels rods terminate at approximately one-third of the bundle length and are desi gnated as short part length fuel rods. Eight fuel rods are used as tie rods. The rods are spaced and supported by the upper and lower tie plates and eight spacers over the length of the fuel rods. For GNF2, the channel interacts with the Lower Tie Plate (LTP).
The fuel rods consist of high-density ceramic UO2 or (U, Gd)O2 fuel pellets stacked within Zircaloy-2 cladding. The cladding will generally have an inner zirconium liner. The fuel rod is evacuated and backfilled with helium.
Additional assembly and component description for the GE14 fuel design are provided in Reference 59.
The ATRIUM-9B reload fuel assembly design (Figure 4.2-3c) is a 9 x 9 array with 72 enriched uranium fuel rods. The interior is an inert water channel. The ATRIUM-10 reload fuel assembly design (Figure 4.2-3d) is a 10X10 array with 83 full-length fuel rods, 8 part-length fuel rods, and one centrally located water channel. The lower tie plate has a nosepiece which has the function of supporting the fuel assembly in the reactor. The upper tie plate has a handle for transferring the fuel bundle from one location to another. The identifying assembly number is engraved on the top of the handle and a boss projects from one side of the handle to LSCS-UFSAR 4.2-19 REV. 20, APRIL 2014 aid in assuring proper fuel assembly orientation. Both upper and lower tie plates position the rod ends for operation and handling. The tie plates also support the weight of the fuel during operation and handling in the 8x8R, 8x8NB, and ATRIUM-9B fuel designs. For the ATRIUM-10 fuel design, the weight of the fuel is supported by the water channel. Finger springs are also employed with the LSCS design. The finger springs are located between the lower tie plate and the channel for the purpose of controlling the bypass flow through that flowpath (Figure 4.2-2, Flow Path 8).
Additional details of the finger springs ar e provided in Section 9 of References 14 and 49. Zircaloy fuel rod spacers equipped with Inconel springs maintain rod-to-rod
spacing.
AREVA Fuel
For the AREVA ATRIUM-9B fuel, eight of the fu eled rods are tie rods. Some of the rods contain gadolinia as a burnable absorber. Fuel rod pitch is maintained by seven spacers. The spacers are a weld ed zircaloy-4 structure with Inconel 718 springs. The centrally located water channel captures the spacers to maintain the proper axial spacing.
The assembly contains one water cha nnel to improve uranium utilization and operational flexibility. It provides unv oided water to the inner portion of the assembly, thereby, providing additional mo deration. The relatively large amounts unvoided water in the interior of the assembly increases the hot-cold reactivity swing. This feature allows greater operat ional flexibility by allowing longer cycles while maintaining appropriate shutdown margin.
For fuel rod removal, the upper tie plate must be depressed against the compression springs a short distance in order to allow the locking sleeves to be rotated 90. After rotating the locking sleeves, the upper tie pl ate is then free to be removed for fuel rod extraction or replacement.
The lower tie plate consists of a machined stainless steel casting with a grid plate for lower end cap engagement and a lower nozzle to distribute coolant to the
assembly.
The upper tie plate is a cast and machined grid plate with attached bail handle to provide for fuel assembly handling and orie ntation. A unique serial identification number is engraved on the bail handle of each tie plate. This number can be read under water to allow identification of the assemblies in the core.
The identification of fuel type and enrichment may be marked on the end of each fuel rod upper end cap.
Additional assembly and component descriptions for the ATRIUM-9B fuel are provided in References 46 and 48.
LSCS-UFSAR 4.2-19a REV. 20, APRIL 2014The ATRIUM-10 fuel assembly consists of many of the same design features as the ATRIUM-9B presented above. Specifically, the ATRIUM-10 consists of a lower tie plate with a debris filter (F UELGUARD), an upper tie plate, 91 fuel rods, 8 spacer grids, a central water channel (or box) and miscellaneous assembly hardware. Of the 91 fuel rods, 8 are part-length fuel ro ds. The structural members of the fuel assembly include the tie plates, spacer grids, water channel, and connecting hardware. The structural connection between the lower tie plate and upper tie plate is provided by the water channel.
Seven spacers occupy the normal axial locations, while an eighth spacer is located a few inches above the lower tie plate. In a manner similar to an AREVA PWR desi gn, the lowermost spacer restrains the fuel rods just above the lower tie plate.
Additional assembly and component descriptions for the ATRIUM-10 fuel are provided in References 55 and 56.
4.2.2.4 Fuel Rod
Each fuel rod consists of high density (>
95% of theoretical density) UO 2 fuel pellets stacked in a Zircaloy cladding tube which is evacuated, backfilled with helium, and sealed by Zircaloy end plugs welded in each end. Beginning with the fresh fuel in LaSalle Unit 1 Cycle 2, all fuel rods are zirconium-barrier fuel with the exception of the fuel rods in 48 ATRIUM-10 fuel bundles first loaded in LaSalle Unit 2 Cycle 10.
The zirconium-barrier fuel has a zircaloy fuel cladding with a metallurgically bonded layer of zirconium on the inner surface. Adequate free volume is provided within each fuel rod in the form of pellet-to-cladding gap and a plenum region at the top of the fuel rod to accommodate thermal and irradiation expansion of the UO 2 and the internal pressures resulting from the helium fill gas, impurities, and gaseous fission products liberated over the design life of the fuel. A plenum spring, or retainer, is provided in the plenum space to prevent movement of the fuel column inside the fuel rod during fuel shipping and handling (Figure 4.1-3). For GE fuel bundles, a hydrogen getter is also prov ided in the plenum space as assurance against the inadvertent admission of moisture or hydrogenous impurities into a fuel rod. Additional information concerning the getter is provided in Section 8 of Reference 14 and in Reference 41.
Prior to the introduction of ATRIUM-10 fuel design at LaSalle, three types of rods were used in GE and ATRIUM-9B fuel bundles: standard rods, tie rods, and nonfueled water rods (Figures 4.2-3 throug h 4.2-3f). The eight tie rods in each bundle have upper end plugs which extend through the upper tie plate casting. The eight tie rods are structural members of th e fuel assembly. They serve to connect the upper and lower tie plates. The tie rods contain fuel and have upper and lower end caps designed for connection to the tie plates. These rods are threaded into the lower tie plate and latch into the upper tie plate to hold the assembly together. The tie rods carry the assembly weight duri ng handling and provide the coil spring reaction support. These tie rods support the weight of the assembly only during fuel handling operations when the assembly hangs by the handle; during operation, LSCS-UFSAR 4.2-20 REV. 20, APRIL 2014the fuel rods are supported by the lower tie plate. Fifty-four rods in the 8x8R and BP8x8R bundles are standard fuel rods. The GE8x8EB bundle has fifty-four standard fuel rods, eight tie rods, and two water rods. The GE8X8NB fuel design contains fifty-two standard fuel rods, eight tie rods, and one centrally located water rod.
The GE14 fuel design, inserted in LaSalle after the ATRIUM-10 design, contains 70 full length standard rods, 14 part length fu el rods, 8 tie rods and 2 large non fueled water rods.
The GNF2 fuel design (Figure 4.2-3f), in serted starting in LaSalle-1 Cycle 15, contains 70 full length standard rods, 8 long part length fuel rods, 6 short part length rods, 8 tie rods and 2 large non fueled water rods.
The ATRIUM-9B has 64 standard fuel rods, 8 tie rods, and one centrally located square water channel. The end plugs of the standard rods have shanks which fit into bosses in the tie plates. An expansion spring is located over the upper end plug shank of each rod in the assembly to keep the rods seated in the lower tie plate while allowing independent axial expansion by sliding within the holes of the upper tie plate. For AREVA 9X9 fuel, all fuel rods except for the tie rods have coil compression springs located between the top of the fuel rods and the bottom surface of the upper tie plate. These compression springs provide a force to aid in seating the fuel rods in the lower tie plate and react against the upper tie plate. The springs accommodate variations in rod lengths arising from manufacturing tolerances and permit axially non-uniform thermal and irradiation induced growth of the fuel rods (Reference 49).
Two rods in each 8x8R and BP8x8R fuel bundle are hollow water tubes, one of which (the spacer-positioning water rod) positions seven Zircaloy fuel rod spacers axially in the bundle. The GE8x8EB fuel bundle may have more water rods, and the GE8X8NB has one large centrally located water rod. The spacer rods are hollow Zircaloy tubes. The spacer-positioning water rod is equipped with the square bottom end plug. The spacer-positioning water rod is assembled to the spacers by sliding the rod through the spacer cells with the welded tabs oriented in the direction of the corner of the spacer cell. The rod is then rotated so that the tabs fit above and below the elements of the spac er structure, thereby positioning the spacer in the required axial position. The rod is prevented from rotating and unlocking the spacers by the engagement of its (square) lower end plug with the tie plate hole. Several holes are punched around the circumference of each of the water rods near each end to allow coolan t water to flow through the rod.
LSCS-UFSAR 4.2-21 REV. 20, APRIL 2014In the GE14 and GNF2 fuel design, two rods in the bundle are hollow water tubes, one of which positions eight high performance Zr-2 fuel rod spacers axially in the bundle. These two water rods are hollow Zircaloy tubes that encompass eight fuel rod positions. The spacer positioning wa ter rod has tabs welded on it above and below each spacer position. This water rod acts as the spacer capture rod for the fuel assembly. The tabs pr event excessive movement of the fuel spacers in either the upward or downward directions.
Several holes are punched around the circumference of each of the water rods near each end to allow coolant water to flow through the rod.
For the ATRIUM-9B fuel, one essentially square water channel is located in the central region of the fuel assembly replacing nine fuel rods in a 3 x 3 array. The water-filled channel has inlet and outlet holes located at the lower and upper end caps. These holes are dimensioned to maintain unvoided water during steady-state operation inside the water channel. The end fittings are made of zircaloy-4. The channel is made from two "U" shaped strips cut and formed from the same sheet of zircaloy-4.
The wall thickness is 0.0285 inches and prov ides adequate strength. The lower end cap of the water channel is threaded and it connects to the lower tie plate. The upper end cap penetrates the upper tie plate and provides a sliding joint to allow for differential growth.
The water channel has zircaloy stops welded on the outside of the channel at axial locations corresponding to the spacer locations. There is a small gap between the stops and each spacer to allow differential thermal expansion between the channel and the fuel rods (Reference 49).
The ATRIUM-10 fuel bundle design is similar in design to the ATRIUM-9B. The most significant difference is in the load-b earing member of the fuel bundle. The ATRIUM-10 does not utilize tie-rods. Instead the central water channel bears the load of the assembly. The attachment of the upper tie plate is accomplished using a simple locking mechanism. All moveable parts in the mechanism are captured such that no parts can come loose during tie plate removal or reactor operation. The reduced number of components results in pa rt from having a single upper tie plate locking mechanism. To keep the upper ti e plate in place, there is one, large compression spring on the water channel rather than the multitude of compression springs on individually fuel rods commonly associated with other designs. Also, no tie rod nuts or locking tabs are required as the water channel carries the weight of the fuel assembly during movement rather than tie rods as in most other BWR fuel designs. Additional component information for the ATRIUM-10 fuel design is provided in References 55 and 56.
LSCS-UFSAR 4.2-21a REV. 20, APRIL 20144.2.2.5 Fuel Pellets The fuel pellets consist of high density ceramic uranium dioxide manufactured by compacting and sintering uranium dioxide powder into right cylindrical pellets. The GE pellets have flat ends and chamfered edges while ATRIUM-9B and ATRIUM-10 pellets are dished and have an outward land taper. Ceramic uranium dioxide is chemically inert to the cladding at operating temperatures and is resistant to attack by water.
Several U-235 enrichments are used in the fuel assemblies. Fuel element design and manufacturing procedures have been developed to prevent errors in enrichment location within a fuel assembly. The LSCS fuel bundle incorporates the use of small amounts of gadolinium as a burnable poison in selected fuel rods.
The GE 8x8R, GE 8x8NB, GE14, GNF2, ATRIUM-9B and ATRIUM-10 fuel design features are summarized in Tables 4.2-4 through 4.2-4(f). Characteristics of other fuel types used at LSCS are given in References 41, 44, 46, 49, 58, and 59.
4.2.2.6 Fuel Channel Separate licensing topical reports (Ref erences 17, 41, 45, 58, and 59) provide complete descriptions and analytical results for channels supplied by General
Electric Company and used in conjunction with the fuel described herein. The use of the GE14 fuel design at LaSalle introd uces the first use of a non-uniform wall thickness channel. The GE14 channel is an interactive channel with a thick corner-thin wall design (120 mil corners and 75 mil wall thickness).
This channel is described in more detail in Reference 58.
Reference 57 contains the specific design details for the fuel channels supplied by AREVA. The GNF2 channel is essentially the same as the GE14 channel with the ex ception that the GNF2 channel interacts directly with the lower tie plate and does not require finger springs (Reference 59).
Beginning with LaSalle Unit 2 Cycle 13 and co ntinuing in subsequent cycles of Unit 2, AREVA Lead Fuel Assemblies (LFAs) ar e inserted into non-limiting core locations for demonstration purposes. The LFA program consists of eight Lead Fuel Assemblies of ATRIUM 10 fuel with advanced alloy channels.
The advanced alloy channel is dimensionally similar to channels described in Reference 57. The variation from previous ly approved channels is the zircaloy comprising the channel is Zircaloy-BWR vers es Zircaloy-2 or Zircaloy-4 described in Reference 57. Advanced alloy channels are included in Mechanical Design Report for LaSalle Unit 1 and 2 ATRIUM 10 Fuel Assemblies , ANP-2741P, Revision 0, August 2008. However, the following functional description is included in this report for completeness.
LSCS-UFSAR 4.2-22 REV. 13 The BWR Zircaloy fuel channel performs the following functions:
(1) Forms the fuel bundle flow path outer periphery for bundle coolant flow.
(2) Provides surfaces for control rod guidance in the reactor core.
(3) Provides structural stiffness to the fuel bundle during lateral loadings applied from fuel rods through the fuel
spacers.
(4) Minimizes, in conjunction with the finger springs and bundle lower tieplate, coolant bypass flow at the channel/lower tieplate interface.
(5) Transmits fuel assembly seismic loadings to the top guide and fuel support of the core internal structures.
(6) Provides a heat sink during loss-of-coolant accident (LOCA).
(7) Provides a stagnation envelope for in-core fuel sipping.
The channel is open at the bottom and makes a sliding seal fit on the lower tieplate surface. The upper end of the fuel assemblies in a four-bundle cell are positioned in the corners of the cell against the top guide beams by the channel fastener springs. At the top of the channel, two diagonally opposite corners have welded tabs, one of which supports the weight of the channel from a threaded raised post and the upper tieplate. One of these raised posts has a threaded hole. The channel is attached using the threaded channel fastener a ssembly, which also includes the fuel assembly positioning spring. Channel-to-c hannel spacing is provided for by means of spacer buttons located on the upper port ion of the channel adjacent to the control rod passage area.
In the mid 1970s, channel box wear and cracking was observed, first in a foreign plant and later in a few domestic boiling water reactors. The wear was located adjacent to incore neutron monitor and startup source locations. It was postulated and later confirmed by out-of-reactor testing, that the wear was caused by vibration of the incore tubes due primarily to a high-velocity jet of water flowing through the bypass flow holes in the lower core plate. To eliminate significant vibration of instrument and source tubes and the resultant wear on channel loop corners, LaSalle incorporated modifications similar to those described in Reference 36.
These modifications involve the elimination of the bypass holes in the lower core plate and addition of two holes in the lower tie plate of each assembly to provide an alternate flow path. This design modification has been determined to have LSCS-UFSAR 4.2-23 REV. 14, APRIL 2002 negligible adverse effects on the mechanical, thermal, and nuclear performance of the channel boxes. Channel box wear has been observed to have been significantly reduced in operating boiling water reactors following the design modification.
Proper orientation of fuel assemblies in the reactor core is readily verified by visual observation and is assured by verification procedures during core loading. Five separate visual indications of proper fuel assembly orientation exist:
- a. The channel fastener assemblies, including the spring and guard used to maintain clearances between channels, are located at one corner of each fuel assembly adjacent to the center of the control rod.
- b. The identification boss on the fuel assembly handle points toward the adjacent control rod.
- c. The channel spacing buttons are adjacent to the control rod passage area.
- d. The assembly identification numbers which are located on the fuel assembly handles are all readable from the direction of the center of the cell.
- e. There is cell-to-cell symmetry.
Experience has demonstrated that these desi gn features are clearly visible so that any misoriented fuel assembly would be re adily distinguished during core loading verification.
Appropriate description and design drawings of reactivity control assemblies are included in Subsection 4.6.1.1.2.
4.2.2.7 Reactivity Control Assembly and Burnable Poison Rods 4.2.2.7.1 Control Rods The control rods perform the dual function of power shaping and reactivity control. Four types of control rods are used at LSCS. Three designs are supplied by General Electric, and the fourth type supplied by ASEA-ATOM (ABB). Power distribution in the core is controlled during operation of the reactor by manipulating selected patterns of control rods. Control rod displacement tends to counterbalance steam void effects at the top of the core and re sults in significant axial power flattening.
4.2.2.7.1.1 General Electric Control Rods Figures 4.1-4(a,b,c) show drawings of the General Electric Control Rods.
LSCS-UFSAR 4.2-24 REV. 18, APRIL 2010 The General Electric original equipme nt and Duralife 215 control rod designs consist of a sheathed cruciform array of stainless steel tubes filled with boron-carbide powder. The control rods are 9.74 inches in total span and are separated uniformly throughout the core on a 12-inch pitch. Each control rod is surrounded by four fuel assemblies.
The main structural member of Original Equipment and Duralife 215 control rod designs is made of Type 304 stainless steel and consists of a top handle, a bottom casting with a velocity limiter and contro l rod drive coupling, a vertical cruciform center post, and four U-shaped absorber tube sheaths. The top handle, bottom casting, and center post are welded into a single skeletal structure. The U-shaped sheaths are resistance-welded to the center post, handle, and castings to form a
rigid housing to contain the boro n-carbide-filled absorber rods.
Rollers at the top and bottom of the control rod guide the control rod as it is inserted and withdrawn from the core. The control ro ds are cooled by the core bypass flow. The U-shaped sheaths are perforated to al low the coolant to circulate freely about the absorber tubes. Operating experience has shown that control rods constructed as described above are not susceptible to dimensional distortions.
The boron-carbide (B 4 C) powder in the absorber tubes is compacted to about 70% of its theoretical density. The boron-carbid e contains a minimum of 76.5% by weight natural boron. The boron-10 minimum content of the boron is 18% by weight.
Absorber tubes are made of Type 304 (or 304 rad resist) stai nless steel. Each absorber tube is 0.188 inch in outside diam eter and has a 0.025-inch wall thickness.
Absorber tubes are sealed by a plug welded into each end. The boron-carbide is longitudinally separated into individual compartments by stainless steel balls at approximately 16-inch intervals. The steel balls are held in place by a slight crimp of the tube. Should boron-carbide tend to compact in service, the steel balls distribute the resulting voids over the length of the adsorber tube.
The Marathon design consists of square outer tubes with round inner diameters welded together and filled with B 4C capsules and hafnium rods. The Marathon design utilizes a 316 stainless steel handle, tie rod, transition piece, fins and locking plug. As of 1999, velocity limiter utilized on General Electric designs (fabricast) is made of CF3 casting. The absorber tubes are made of Rad Resist 304S stainless steel and welded together for rigidity. Some Marathon control blade handles have rollers or buttons to provide guidance for control rod insertion and withdrawal. Some Marathon control blade handles have no rollers or pads.
4.2.2.7.1.2 ASEA-ATOM (ABB) Control Rods The second type of Control Rod utilized at LSCS is the ASEA-ATOM (ABB) CR82B. The ASEA-ATOM control rod functions the same as the General Electric control LSCS-UFSAR 4.2-24a REV. 18, APRIL 2010 rod, however the design of the ASEA-ATOM control rod is slightly different. Each of the four ASEA-ATOM control blade wi ngs has 520 horizontal holes (0.20 inch diameter) drilled directly into the blade wing (thus eliminating the perforated U-shaped absorber tube sheaths used in th e General Electric Control Rod design).
The first 6 inches of the blade (beneath the top handle) consist of 22 holes containing hafnium rodlets. The remaining 498 holes contain boron-carbide powder compacted to above 70% of its theoretical density. The boron-carbide contains between 76.5-81% by weight natural boro
- n. The boron-10 content in the ASEA-ATOM control rods is 19.9 +/- 0.3 atom %. The horizontal holes are covered with a stainless steel bar at the outer edge of the blade wing and are connected through a narrow slit. This allows gas pressure equalization between holes and prevents significant displacement of LSCS-UFSAR 4.2-25 REV. 17, APRIL 2008 the B 4 C powder.
4.2.2.7.2 Velocity Limiter
The control rod velocity limiter (Figures 4.2-5 and 4.2-5a) is an integral part of the bottom assembly of each control rod. This engineered safeguard protects against a high reactivity insertion rate by limiting the control rod velocity in the event of a control-rod-drop. It is a one-way device in that the control rod scram velocity is not significantly affected but the control rod dropout velocity is reduced to a permissible limit.
The velocity limiter is in the form of two nearly mated conical elements that act as a large clearance piston inside the control rod guide tube. The lower conical element is separated from the upper conical element by four radial spacers 90 degrees apart and is at a 15-degree angle relative to the upper conical element, with the peripheral separation less than the central separation.
The hydraulic drag forces on a control rod are proportional to approximately the square of the rod velocity and are negligible at normal rod withdrawal or rod insertion speeds. However, during the scram stroke, the rod reaches high velocity and the drag forces must be ove rcome by the drive mechanism.
To limit control rod velocity during dropout but not during scram, the velocity limiter is provided with a streamlined profile in the scram (upward) direction. Thus, when the control rod is scrammed, water flows over the smooth surface of the upper conical element into the annulus between the guide tube and the limiter. In the dropout direction, however, water is trapped by the lower conical element and discharged through the annulus between th e two conical sections. Because this water is jetted in a partially reversed direction into water flowing upward in the annulus, a severe turbulence is created, th ereby slowing the descent of the control rod assembly to less than 3.11 ft/sec for current control blade designs.
4.2.2.7.3 Burnable Poison Rods To meet the reactivity control requirements of any core load with excess reactivity, gadolinia-urania fuel rods are placed in each fuel assembly except for the natural uranium assemblies used in the initial cycle for both units and 48 low enriched ATRIUM-10 bundles first loaded in LaSalle Unit 2 Cycle 10. Some assemblies contain more gadolinia than others to improve transverse power flattening. Also, some assemblies contain axially distributed gadolinium to improve axial power flattening. GD 2 O 3 is uniformly distributed in the UO 2 pellet and forms a solid solution.
LSCS-UFSAR 4.2-26 REV. 20, APRIL 2014 4.2.3 Design Limits and Evaluation A discussion of the fuel thermal-mechanical design limits and evaluation results for the BP8x8R, GE8x8EB, GE8X8NB, GE14, and GNF2 fuel designs is given in Section 2 of Reference 41. A similar discussion of the limits and results for the 8x8R fuel design is given in Appendix C of this reference. A similar discussion of the thermal mechanical design limits and evaluation results for the AREVA fuel can be found in Reference 46 through 49, 55, and 56. The information contained in the following Subsections is provid ed as a historical reference.
4.2.3.1 Fuel Damage Analysis
Fuel damage is defined as a perforation of the fuel rod cladding which would permit the release of fission produc ts to the reactor coolant.
The mechanisms which could cause fuel da mage in reactor operational transients are: (a) severe overheating of the fuel ro d cladding caused by inadequate cooling, and (b) rupture of the fuel rod cladding due to strain caused by relative expansion of the UO 2 pellet. Cladding failure due to ov erpressure from vaporization of UO 2 following a rapid reactivity transient is not considered to be an operational transient.
A value of 1% plastic strain of the Zircaloy cladding has traditionally been defined as the limit below which fuel damage due to overstraining of the fuel cladding is not expected to occur. The 1% plastic strain value is based on General Electric data on the strain capability of irradiated Zircaloy cladding segments from fuel rods operated in several BWR's (Reference 4). None of the data obtained falls below the 1% plastic strain value. However, a statistical distribution fit to the available data indicates the 1% plastic strain value to be approximately the 95% point in the total population. This distribution implies, th erefore, a small (< 5%) probability that some cladding segments may have plastic elongation less than 1% at failure.
For fresh UO 2 fuel the calculated linear heat generation rate (LHGR) corresponding to 1% diametral plastic strain of the cla dding is approximately 25 kW/ft. Later in life, the calculated LHGR corresponding to 1% diametral plastic strain decreases to
approximately 24 kW/ft at 20,000 MWd/tU and approximately 22 kW/ft at 40,000 MWd/tU. However, due to a depletion of fissionable material, the high-exposure fuel has less nuclear capability and will operate at correspondingly lower powers, so that a wide margin is maintained throughout life between the operating LHGR and
the LHGR calculated to cause 1% cladding diametral strain.
The addition of small amounts of gadolinia to UO 2 results in a reduction in the fuel thermal conductivity and melting temperature. The result is a reduction in the LHGR's calculated to cause 1% plastic diametral strain for gadolinia-urania fuel rods. However, to compensate for this th e gadolinia-urania fuel rods are designed
to provide margins si milar to standard UO 2 rods.
LSCS-UFSAR 4.2-27 REV. 20, APRIL 2014 4.2.3.2 Fuel Damage Experience The early GE BWR fuel experience has been extensively described in previous reports. In general, the Zircaloy cladding performance in the very early plants was good; however, some fuel failure mechanis ms were encountered and corrected. They are not significantly affecting current fuel performance. Details of this experience are provided in References 4, 19, 20 and 40. Later BWR fuel experience is given in Reference 41.
One of the early causes of fuel failures was internal hydriding of the Zircaloy cladding due to internal attack by hydrogen. The source of hydrogen was primarily small amounts of moisture introduced into the fuel rod. A detailed analysis of the potential sources of hydrogen or moisture shows that the only source large enough to explain primary hydride failure was the UO 2 pellet itself. Major process steps such as increased fuel rod drying temperatures and dry grinding of pellets were incorporated in the manufacture of UO 2 pellets to ensure that no significant moisture could be present in the as-fabricat ed fuel rod. In addition, the fuel rod design was changed to incorporate a hydrogen gettering system to further assure that neither moisture nor any sporadic hydrogen is ever available to cause hydride failure of the cladding. Newer fuel designs, such as GNF2 fuel rods, are manufactured to a tighter hydrogen control limit as described in Reference 41; therefore, these newer fuel designs do not include the hydrogen getter.
Another fuel failure mechanism encountered in operating BWR fuel is crud induced localized corrosion (CILC). CILC, however, has not been experienced at LaSalle.
The one class of fuel failure mechanisms which has restricted operation on LaSalle Units 1 and 2 is known as "pellet-cladding interaction" (PCI). The failures are caused by the direct interaction between the irradiated urania fuel, including its inventory of fission products, and the zi rcaloy fuel sheath, or cladding. The incidence of such failures is closely linked to the power history of the fuel rod and to the severity and duration of power changes. Consequently, in order to reduce the probability of fuel failures due to the PCI phenomenon, operational constraints were placed on the reactors.
These constraints were placed on local nodal power increases (ramp rates). Although these constraints have been very successful in reducing the incidence of fuel failures, they were costly in terms of operational flexibility. Consequently, there was strong incentive to provide a type of fuel resi stant to PCI. There have been a number of fuel design improvements that were made to minimize PCI failures. These improvements include:
(a) the pellet geometry has been modified to include chamfered pellet ends and a shorter length in order to reduce the magnitude of inservice pellet distortions contributing to local cladding strains.
For AREVA Fuel, the pellet geometry includes a land taper, dish and short length for enriched and gadolinia pellets. These features have been shown to reduce PCI.
LSCS-UFSAR 4.2-28 REV. 18, APRIL 2010 (b) the cladding heat treatment temperature has been increased in order to reduce the statistical variability in cladding mechanical properties; (c) change from 7 x 7 to 8 x 8 to 9 x 9 to 10x10 lattice design to reduce fuel thermal duty; and (d) introduction of zirconium-barrier fuel.
Improvements (a), (b) and (c) were made prio r to 1975. These, however, did not totally eliminate the PCI problem and it was necessary for plants to continue operation within the ramp rate guidelines. Extensive testing at Quad Cities Unit 2 showed that the introduction of zirconium-barrier fuel eliminated the need for use of the ramp rate guidelines on those fuel assemblies.
The initial cycle fuel for LaSalle Units 1 and 2 did not incorporate the zirconium-barrier fuel. Consequently, operation was maintained within the PCIOMR guidelines for all fuel assemblies. However, reload fuel for subsequent cycles will be zirconium-barrier fuel with the exception of 48 low enriched ATRIUM-10 bundles first loaded in LaSalle Unit 2 Cycle 10. Operation of the zirconium-barrier fuel will be restrained only by the Technical Specifications. However, industry experience will continue to be utilized in order to implement appropriate administrative operating policies that may be more conservative than Technical Specifications. Operation of the non-barrier ATRIUM-10 fuel will be restrained by the guidelines provided by the fuel manufacturer (AREVA).
Operation with failed fuel rods has demonstr ated that the fission product release rate from defective fuel rods can be controlled by regulating power level. The rate of increase in released activity apparently associated with progressive deterioration of failed rods has been deduced from chrono logical plots of the off-gas activity measurements in operating plants. These data indicate that the activity release level can be lowered by lowering the local power dens ity in the vicinity of the fuel rod failure. This measured data also indicates that cata strophic failure of the fuel assembly does not occur upon continued operation and that the presence of a failed rod in a fuel assembly does not result in propagation of failure to neighboring rods. Shutdown can be scheduled, as required, to repair or repl ace fuel assemblies that have large defects.
Evaluation of the fission product release rate for failed fuel rods shows a wide variation in the activity release levels. Correlation of the release rates to defect type, size and specific power level indicates that fission product release rates are functions of power density and that progressive deterioration is a function of time. Available failure data are insufficient to quantify the detaile d correlation between these variables.
4.2.3.3 Potential For a Water-Logging Rupture
For water-logging to occur, the fuel cladding must have a small pinhole. Pinholes are eliminated during production by 100%
leak check of assemblies. The leak LSCS-UFSAR 4.2-29 REV. 18, APRIL 2010 detector system consists of a high vac uum system capable of attaining pressures less than 5 x 10
-3 torr, and a mass spectrometer capable of detecting leaks smaller than the design limit (1 x 10
-8 std. cc/sec). The fuel bundle or fuel rod is placed in the vacuum chamber and evacuated to less than 1 x 10
-4 torr. After the vacuum is attained, the mass spectrometer tuned to the helium mass range is switched into the system. The output meter of the mass spectrometer will indicate the presence of any helium gas in the chamber. The design basis for the fuel precludes the potential for a water-logging rupture throughout the fuel cycle.
4.2.3.4 Potential For Hydriding
The design basis for fuel in regard to the cladding hydriding mechanism is to assure, through a combination of engineering specifications and strict manufacturing controls, that production fuel will not contain excessive quantities of moisture or hydrogenous impurities. Anal ysis of BWR fuel performance on fuel manufactured since July 1972 has indicated that this failure mechanism has been eliminated in BWR fuel through the adoption of the changes in the fuel rod design and manufacturing processes described in Subsection 4.2.3.2 and in Reference 41.
AREVA addressed internal hydriding in Reference 49. The absorption of hydrogen by the cladding can result in cladding failure due to reduced ductility and formation of hydride platelets. Careful moisture control during fuel fabrication reduces the potential for hydrogen absorption on the inside of the cladding. The fabrication limit for total hydrogen in the fuel pellets is less than 2.0 ppm (References 46 and 49).
4.2.3.5 Dimensional Stability The fuel assembly and fuel components are designed to assure dimensional stability in service. The fuel cladding and channel specifications include provisions to preclude dimensional changes due to residual stresses. In addition, the fuel assembly has been designed to accommoda te dimensional changes that occur in service due to thermal differential expansio n and irradiation effects. For example, the fuel rods are free to expand lengthwise independent of each other, and the channel is free to expand relative to the fuel bundle.
The differential thermal expansion betwee n the tie plates and spacer grid is calculated to introduce a bending stress of less than 400 psi at the end of the fuel tube. Additional information regarding this calculation is presented in Section 4 of Reference 1.
LSCS-UFSAR 4.2-29a REV. 18, APRIL 2010 During shipment the fuel bundle is in a horizontal position with flexible packing separators installed between the fuel rod so that the weight of the fuel rods is supported by the shipping co ntainer rather than the spac er grids. AREVA fuel rods are supported by the fuel assembly spacers during shipment. AREVA has performed testing to verify that this is acceptable for the Atrium-9B fuel assembly.
Fuel bundle shipping procedures are qualified by a test performed on each new design, and each individual bundle is inspected relative to important dimensional characteristics following shipment to verify that no dimensional deviations have occurred.
LSCS-UFSAR 4.2-30 REV. 18, APRIL 2010 The two major handling loads of concern are (1) the loads due to maximum upward acceleration of the fuel assembly while grappled, and (2) the loads due to impact of the fuel assembly into the fuel support wh ile grappled. Analyses of these loading conditions have been performed and the resulting fuel assembly component stresses are within design limits. Additional in formation on fuel handling and shipping loads for GE fuel is presented in Sectio n 5 of Reference 1 and in Reference 41.
AREVA addresses fuel assembly handling loads in Reference 49. The AREVA assembly design must withstand all norm al axial loads from shipping and fuel handling operations without permanent de formation. AREVA uses either a stress analysis or testing to demonstrate compliance. The analysis or test uses an axial load of 2.5 times the static fuel assembly weight. At this load, the fuel assembly structural components must not show any yielding. Because of the design, failure from axial loads will occur at the tie rod end caps rather than in the cladding or tie plates. The fuel rod plenum has a design criteria associated with handling requirements. The spring must maintain a force against the stack weight to prevent column movement during handling (Reference 49).
4.2.3.6 Fuel Densification The amount of incore fuel densification in BWR Zircaloy clad UO 2 pellet fuel has been observed to be small and is not cons idered to have any significant effects on fuel performance. Detailed consideration of the occurrence and potential effects of incore fuel densification in General Electr ic BWR's is reported in Reference 5 and its supplements. See Section 4.2.1.2.3.
2 for a similar discussion for AREVA Fuel.
4.2.3.7 Fuel Cladding Temperatures
Fuel cladding temperatures for 8x8R type fuel are shown in Figure 4.2-6 as a function of surface heat flux for beginning of life conditions. A core distribution of segment powers is developed. The value of Zircaloy-2 thermal conductivity used in these calculations is approximately 9.0 Btu/hr-ft F. Calculated fuel cladding temperatures for 8x8R type fuel for late-in-life conditions are shown on Figure 4.2-7 as a function of heat flux. Th e solid lines on Figure 4.2-7 represent the expected fuel cladding temperatures. The temperatures employed in mechanical design evaluations are calculated using a conservative design allowance for the degradation in fuel rod surface heat transfer coefficient due to the accumulation of system corrosion products on the surface of the rod (crud) and cladding corrosion (zirconium oxide formation). The expected fuel cladding temperatures are calculated employing a more realistic allowance for the effects of crud and oxide on the fuel rod surface heat transfer coefficient. The calculated peak cladding temperatures are used in the th ermal and mechanical design analyses addressed in Reference 41. The fuel cladding temperatures for other fuel types can be found in Reference 41.
LSCS-UFSAR 4.2-31 REV. 18, APRIL 2010 AREVA also prevents the fuel rod cladding from overheating by minimizing the probability of exceeding thermal margin limits on limiting fuel rods during normal operation and anticipated operatio nal occurrences (Reference 49).
4.2.3.8 Peaking Factors The typical power distribution is divided into several components: the radial peaking factor, local peaking, and axial peaking. The maximum radial peaking factor is defined as the total power produced in the most limiting fuel assembly divided by the core average fuel assembly power. The maximum local peaking factor is defined as the maximum fuel rod heat flux in a fuel assembly divided by the fuel assembly average fuel rod heat flux. The maximum axial peaking factor is defined as the maximum heat flux along the length of a given fuel rod divided by
the average heat flux of that rod. The initial reactor core de sign employs typical power peaking factors shown in Table 4.4-1. Peaking factors for reload cores are such that margins to limits for LHGR, MCPR, and APLHGR remain within the COLR limits.
4.2.3.8.1 Local Peaking Factors The enrichment distribution in each fuel assembly is selected to reduce the relative local peak-to-average fuel rod power ratio within each assembly. The local peaking factor used for the initial design is provided in Table 4.4-1.
4.2.3.8.2 Axial and Gross Peaking Factors The axial and gross peaking factors used for the initial core design are provided in Table 4.41. Axial and gross peaking factors for reload cores are such that margins to limits for LHGR, MCPR, and APLHGR remain within the COLR limits.
4.2.3.9 Temperature Transients with Waterlogged Fuel Element As indicated in Subsection 4.2.3.3, the po tential for water-logging is considered in the fuel design. For waterlogging to occu r, the fuel cladding must have a small pinhole. Pinholes are eliminated during production by 100% leak check of assemblies. The leak detector system employ ed is described in Subsection 4.2.3.3. Since waterlogging is not expected and since it has not been observed in commercial power BWR fuel, no specific analysis of the consequences is performed.
In the unlikely event that a waterlogged fu el element does exist in a BWR core, it should not have a significant potential for cladding burst (due to internal pressure) during a transient power increase unless the transient started from a cold or very low power condition. Normal reactor heatup rates are sufficiently slow ( 100 F/hr increase in coolant temperature) such that water vapor formed inside a waterlogged fuel rod would be expected to evacuate the rod through the same passage it entered, LSCS-UFSAR 4.2-32 REV.
13 allowing internal and external pressures to equilibrate as the coolant temperature and pressure rise to the rated conditions.
Once the internal and external pressure s are at equilibrium, at rated coolant pressure and temperature, transient power increases should, in general, have the effect of only slightly reducing the internal fuel rod plenum volume due to differential thermal expansion between fuel and cladding, thus effecting a small, short-term increase in internal fuel rod pressure. The potential short-term increase in pressure due to this effect would, in general, be small, (e.g., a power increase from the cold condition to peak rated power would increase internal pressure less than 15% in the peak power fuel rod fuel rod). For the range of anticipated transients, the cladding primary membrane stress resulting from the temporary increase in internal pressure above the coolant pressure would not be expected to exceed the cladding stress design limits of Subsection 4.2.1.2.5.
4.2.3.10 Potential Damaging Temperature Effects During Transients
There are no predicted significant temperature effects during a power transient resulting from a single operator error or single equipment malfunction which result in fuel rod, control rod, or structural damage. The calculated fuel rod cladding strain for this class of transients is sign ificantly below the calculated damage limit. The predicted additional bowing deflection for this class of transients is small compared to the steady-state rod-to-channel clearance.
4.2.3.11 Energy Release During Fuel Element Burnout
The metal-water chemical reaction between zirconium and water is given by:
where H = 140 cal/g-mole. The reaction rate is conservatively given by the familiar Baker-Just rate equation:
where W is milligrams of zirconium reacted per cm 2 of surface area, is time (seconds), R is the gas constant, (cal/mol
- K), and T is the temp erature of zirconium
( K). This rate equation has been shown to be conservatively high by a factor of 2 (Reference 21). The above equation can be differentiated to give the rate at which the thickness of the cladding is oxidized. This becomes:
3-4.2
T A- exp X A th 2 1 where: th = rate at which the cladding thickness is oxidizing, 2-4.2 RT 45,500exp 10 x 33.3W 6 2 1)(4.2 H2HZrO02HZr22 2 LSCS-UFSAR 4.2-33 REV.
13 = oxidized cladding thickness, A 1 , A 2 = appropriate constants, and T = reaction temperature.
The reaction rate is inversely proportional to the oxide buildup; therefore, at a given cladding temperature the reaction rate is self-limiting as the oxide builds up. The total energy release from this chemical reaction over a time period is given by:
where: N rods = number of rods experiencing boiling transition (at temperature T), -H = heat of reaction, C = cladding circumferences, L = axial length rod experiencing boiling transition, and
= density of zirconium.
This equation can be integrated and compa red to the normal bundle energy release if the following conservative assumptions are made:
- a. At an axial plane all the rods experience boiling transition and are at the same temperature. This is highly conservative since, if boiling transition occurs, it will normally occur on the high power rod(s).
- b. Boiling transition is assumed to occur uniformly around the circumference of a rod. This generally occurs only at one spot.
- c. The rods are assumed to reach some temperature T instantaneously and stay at this temperature for an indefinite amount of time.
This integration has been performed per ax ial foot of bundle and the total energy release as a function of time has been comp ared to the total energy release of a high power bundle (6 MW) over an equal amo unt of time. The results are shown in Figure 4.2-8. For example, if the temperatu re of all rods along a 1-foot section of the bundle were instantly increased to 1500 F, the total amount of energy that has 4-4.2
Xdt CL H- N Q t rods T LSCS-UFSAR 4.2-34 REV. 20, APRIL 2014 been released at 0.1 seconds is 0.4% of the total energy that has been released by the bundle (6 MW x 0.1 second). Note th at the fractional energy release decreases rapidly with time even though a constant temperature is maintained. This is because the reaction is self-limiting as was discussed above with the Baker-Just equation.
The amount of energy released is dependent on the temperature transient, and the surface area that has experienced heatup. This, of course, is dependent on the initiating transient. For example, if boiling transition were to occur during steady-state operating conditions, the cladding surface temperature would range from 1000 F to 1500 F depending on the heat fluxes and heat transfer coefficient. Even assuming all rods experience boiling transi tion instantaneously, the magnitude of the energy release is insignificant. Significant boiling transition is not possible at normal operating conditions because of th e thermal margins at which the fuel is operated. This is also true for abnormal tr ansients. It can, therefore, be concluded that the energy release and potential for a chemical reaction is not an important consideration during normal op eration or abnormal transients.
4.2.3.12 Energy Release for Rupture of Waterlogged Fuel Elements Experiments have been performed to show that waterlogged fuel elements can fail at a lower damage threshold than nonwaterlogged fuel during rapid reactivity excursion from the cold condition (Referen ces 22 and 23), (i.e., 60 cal/g as compared to > 300 cal/g). No analysis of cladding stress has been performed by GE for such conditions. One can postulate that if such a failure occurred, the resultant energy release and pressure pulse would be much less than for a nonwaterlogged fuel rod which exceeded its damage threshold since the energy level required for damage is apparently much lower in the waterlogged fuel element. Any fuel dispersion that might result in such a case would further reduce the severity of such a transient.
4.2.3.13 Fuel Rod Behavior Effects from Coolant Flow Blockage In Reference 24, GE evaluated the conseque nces of a fuel bundle flow blockage incident. The percent of flow blocked to the bundle reduces the MCPR margin, and must be considered when evaluating the effects of a known lost part. A portion of reference 24 also discusses the consequences associated with 100% blockage of a fuel bundle; however, this event was never reviewed and approved by the NRC, nor has it ever been made a licensing requirement.
Reference 16 provides an updated discussion, applicable to GE9, GE14, AREVA ATRIUM-9B fuel, and AREVA ATRIUM-10 fuel, of the effects of flow blockages on MCPR margin. Reference 60 provides the flow blockage MCPR analysis for GNF2 fuel. This relationship is used to determine the impact of known lost parts. This document also discusses the potential for fuel fretting for parts small enough to migrate into the bundle. Fuel fretting may lead to fuel failures, which would be detected by the offgas system. If a blocked bundle becomes suddenly unblocked, the increase in reactivity is less than the delayed neutron fraction, and therefore a prompt critical excursion is avoided.
LSCS-UFSAR 4.2-35 REV. 18, APRIL 2010 4.2.3.14 Channel Evaluation
An evaluation of fuel channel loading due to internally applied pressure has been performed. Tests have been conducted to verify the applicability of the "fixed-fixed
beam" analytical model under uniform load.
To confirm the applicability of the analytical model, a channel section was pressurized and the resultant deflections were measured and compared with the deflections predicted by the analytical model. A 4-foot-long section of channel with welded end plates was used for the test. The channel section was pressurized at room temperature in steps up to a pressure which was equivalent to a calculated stress intensity of approximately three times the yield strength of the channel material. Measurements of channel deflection were made for each pressure step and at zero pressure following each step. The deflection of the channel walls was found to be linear with pressure in th e pressure range tested. The measured deflection was within approximately 5% to 10% of the deflection predicted by the analytical model. There was no measurable permanent deformation of the channel walls until the calculated stress in the wall had reached approximately 1.2 times the measured yield strength of the test channel.
The good performance of the channels have been demonstrated by both in-reactor experience and tests. The preponderance of the experience has been with channels that are 5.278 inches inside width with 0.080-inch wall thickness. Channel sizes ranging from 4.290 inches inside width with 0.060-inch walls to 6.543 inches inside width with 0.100-inch. walls, are included.
The LSCS channel is 5.278 inches inside width with either 0.100-inch or 0.080-inch walls, dependin g on the specific reload. Additional information regarding channel analyses is presented in Section 2 of
Reference 1 and in References 17, 45 and 57.
Channel Management
Channels are not being reused at LaSalle. This is one of the assumptions that is used for the MCPR safety lim it calculations by AREVA.
To preclude unacceptable fuel element channel box deflection, a channel verification program, as discussed below, is implemented at LaSalle.
The following general guidelines are followed to detect and control the potential of channel bowing.
- a. Records are kept of channel location and exposure for each operating cycle.
- b. Channels are not retained in the outer row of the core for more than two successive operating cycles.
LSCS-UFSAR 4.2-36 REV. 13 c. At the beginning of each fuel cycle, the combined outer row residence time for any two channels in any control rod cell should not exceed four peripheral cycles.
Prior to the beginning of a new operating cycle, control rod drive friction tests shall be performed for those core cells exc eeding the above general guidelines or containing fuel channels with exposures greater than 30,000 MWd/T (associated
fuel bundle exposures).
In lieu of friction testing, fuel channel me asurements may be used to justify use of fuel channels exceeding 30,000 MWd/T exposu re for a maximum of four additional operating cycles.
In the future, analytical channel lifetime prediction methods, benchmarked and backed by periodic measurements of a sample of the highest duty fuel channels, may be used to assure clearance between control rod blades and fuel channels without additional testing.
4.2.3.15 Fuel Reliability The information in this section is historical GE data on fuel reliability experience. The fuel component characteristics which ca n influence fuel reliability include: (a) the fuel pellet thermal and mechanical pr operties, dimensions, density, and U-235 enrichment; (b) the Zircaloy cladding thermal and mechanical properties, dimensions, and defects; (c) the fuel rod internal void volume and impurities; (d) the fuel rod-to-rod and rod-to-channel spacing; and (e) the spring constants of the fuel rod spacer springs which maintain contact between the spacer and the fuel rods.
Important fuel pellet, cladding, and associated hardware characteristics and dimensions for the 8x8R fuel design are provided in Table 4.2-4 and Figure 4.1-2. The characteristics of other fuel designs ma y be found in Reference 41, 47, 48, or 49.
The large volume of irradiation experience to date with GE BWR fuel indicates only a few mechanisms which have actually had a direct impact on fuel reliability; namely, cladding defects, excessive deposition of system corrosion products, cladding hydriding resulting from hydrogen impurity, and pellet-cladding interaction.
The cladding defects have been virtually eliminated through implementation of improved quality inspection equipment and more stringent quality control requirements during fuel fabrication. Ex cessive deposition of corrosion products has also been virtually eliminated throug h improved control of corrosion product impurities in the reactor feedwater and by manufacturing improvements. Cladding hydriding is the result of excessive amounts of hydrogenous impurities (moisture and/or hydrogenous material) inadvertently introduced into the rod LSCS-UFSAR 4.2-37 REV. 18, APRIL 2010 during the fuel fabrication process. An alysis of BWR fuel performance on fuel manufactured since July 1972 has indicated that this failure mechanism has been eliminated in BWR fuel through the adoption of the changes in the fuel rod design and manufacturing processes described in Subsection 4.2.3.2 and Reference 41.
Pellet-cladding interaction is the fuel failure mechanism which currently has the greatest effects on reactor operation at LaSalle. It has been identified as resulting from the combination of two basic effects:
(a) the observed variability in local cladding strains due to pellet-cladding interaction which can result in the random occurrence of higher-than-average local strain value; and (b) the statistical
variability in postirradiation ductility of the cladding which can result in the random occurrence of tubing segments with ductility lower than average. The fuel design improvements described in Subsecti on 4.2.3.2 have been shown to virtually eliminate PCI as a major cause of fuel failures. When zirconium-barrier fuel replaces all initial cycle fuel, the ramp rate guidelines may be virtually eliminated as a restraint on reactor operations. However, administrative restrictions may still be maintained.
The cladding liner material is an enhanced zirconium alloy. The purpose of the material enhancement to the liner is to reduce the potential for secondary hydriding following the intrusion of coolant into a fuel rod.
4.2.3.16 Fuel Operating and Developmental Experience
Production fuel rods employing gadolinia-urania fuel pellets have been in use since 1965. Fuel operating experience is docu mented in References 4, 19, 40 and 42.
4.2.3.17 Fuel Assembly
During shipment the fuel bundle is in a horizontal position with flexible packing separators installed between the fuel rods so that the weight of the fuel rods is supported by the shipping co ntainer rather than the spac er grids. AREVA fuel rods are supported by the fuel assembly spacers during shipment. AREVA has performed testing to verify that this is acceptable for the Atrium-9B and ATRIUM-10 fuel assemblies. Fuel bundle shipping procedures are qualified by a test performed on each new design, and each individual bundle is inspected relative to important dimensional characteristics following shipment to verify that no dimensional deviations have occurred.
The two major handling loads of concern are (1) the loads due to maximum upward acceleration of the fuel assembly while grappled, and (2) the loads due to impact of the fuel assembly into the fuel support wh ile grappled. Analyses of these loading conditions have been performed and the resulting fuel assembly component stresses are within design limits. Additional in formation of fuel handling and shipping loads is presented in Section 5 of Reference 1 and in Reference 41.
LSCS-UFSAR 4.2-38 REV. 18, APRIL 2010 AREVA has also evaluated their fuel for fu el handling and shipping concerns. The assembly design must withstand all norm al axial loads from shipping and fuel handling operations without permanent de formation. AREVA uses either a stress analysis or testing to demonstrate compliance (Reference 46 and 55).
The rod plenum spring also has design criteria associated with handling requirements. The spring must maintain a force against the stack weight to prevent column movement during handling.
4.2.3.17.1 Loads Assessment of Fuel Assembly Components
The analytical methods and acceptance criteria applied to determine the fuel assembly response to externally applied forc es are both deemed to be in accordance with the requirements of Appendix A to SRP 4.2. LaSalle County Station fuel assembly capability has been evaluated accordingly with acceptable results. Information on load assessment of fuel assembly components is provided in Table 3.9-4.
4.2.3.18 Spacer Grid and Channel Boxes Refer to Subsection 4.2.3.14.
4.2.3.19 Burnable Poison Rods
The failure rate of the gadolinia-urania fuel rods is negligible, from previous operating experience over the years.
4.2.3.20 Control Rods
4.2.3.20.1 Materials Adequacy Throughout Design Lifetime
The adequacy of the materials throughout the design life was evaluated in the mechanical design of the control rods. The primary materials, B 4 C powder, Hafnium, and Type 304 and Type 316L austen itic stainless steel, have been found suitable in meeting the demands of the BWR environment.
4.2.3.20.2 Dimensional and Tolerance Analysis
Layout studies are done to ensure that, given the worst combination of extreme detail part tolerances at assembly, no interference exists which will restrict the movement of control rods. In addition, preoperational verification is made on each control blade assembly to show that the acceptable levels of operational performance are met.
LSCS-UFSAR 4.2-39 REV. 13 4.2.3.20.3 Thermal Analysis of the Tendency to Warp
All parts of the control rod assembly remain at approximately the same temperature during reactor operation, ne gating the problem of distortion or warpage. Differential thermal growth is allowed for in the mechanical design. A minimum axial gap is maintained between absorber rod tubes and the control rod frame assembly for this purpose. In addition, dissimilar metals are avoided.
4.2.3.20.4 Forces for Expulsion
An analysis was made to evaluate the maximum pressure forces which could tend to eject a control rod from the core. The results of this analysis are given in
Subsection 4.6.2.3.1.2.2 under item "Rupture of Hydraulic Line(s) to Drive Housing Flange." In summary, if the collet were to remain open, which is unlikely, calculations indicate that the steady-state control rod withdrawal velocity would be 2 ft/sec for a pressure-under line break, the limiting case for rod withdrawal.
4.2.3.20.5 Functional Failu re of Critical Components The consequences of a functional failure of critical components have been evaluated and the results are covered in Subsection 4.6.2.3.2.
4.2.3.20.6 Precluding Excessive Rates of Reactivity Addition
In order to preclude excessive rates of reactivity addition, analysis has been performed both on the velocity limiter device and the effect of probable control rod failures (Subsection 4.6.2.3.2).
4.2.3.20.7 Effect of Fuel Rod Failure on Control Rod Channel Clearances
The control rod drive mechanical design ensures a sufficiently rapid and forceful insertion of control rods so that no channe l misalignments or distortion could hinder reactor shutdown by impeding a significant number of rods from full insertion.
4.2.3.20.8 Mechanical Damage
Analysis has been performed for all areas of the control system showing that system mechanical damage does not affect the cap ability to continuously provide reactivity control.
The following discussion summarizes the analysis performed on the control rod guide tube.
LSCS-UFSAR 4.2-40 REV. 13 The guide tube can be subjected to any or all of the following loads:
- a. inward load due to pressure differential,
- b. lateral loads due to flow across the guide tube, c. dead weight, and
- d. seismic.
In all cases analysis was performed considering both a recirculation line break and a steamline break, events which result in the largest hydraulic loadings on a control rod guide tube.
Two primary modes of failure were considered in the guide tube analysis: exceeding allowable stress and excessive elastic deformation. It was found that the allowable stress limit will not be exceeded and that the elastic deformations of the guide tube never are great enough to cause the free movement of the control rod to be jeopardized.
4.2.3.20.8.1 First Mode of Failure
The first mode of failure is evaluated by the addition of all the stresses resulting from the maximum loads for the faulted condition. This results in the maximum theoretical stress value for that conditio
- n. Making a linear supposition of all calculated stresses and comparing this value to the allowable limit defined by the ASME Boiler and Pressure Vessel Code yields a factor of safety of approximately 3. For faulted conditions the factor of safety is approximately 4.2.
4.2.3.20.8.2 Second Mode of Failure
Evaluation of the second mode of failure is based on clearance reduction between the guide tube and the control rod. The minimum allowable clearance is about 0.1 inch. This assumes maximum ovality and minimum diameter of the guide tube and the maximum control rod dimension. The analysis showed that if the approximate 6000 psi for the faulted condition were entire ly the result of differential pressure, the clearance between the control rod and th e guide tube would reduce by a value of approximately 0.01 inch. This gives a design margin of 10 between the theoretically calculated maximum displacement and the minimum allowable clearance.
4.2.3.20.9 Analysis of Guide Tube Design Two types of instability were considered in the analysis of guide tube design. The first was the classic instability associated with vertically loaded columns. The second was the diametral collapse when a circular tube experiences external to internal differential pressure.
LSCS-UFSAR 4.2-41 REV.
13 The limiting axially applied load is ap proximately 77,500 pounds resulting in a material compressive stress of 17,450 psi (code allowable stress). Comparing the actual load to the yield stress level gives a design margin greater than 20 to 1. From these values it can be concluded that the guide tube is not an unstable column.
When a circular tube experiences external to internal differential pressure, two modes of failure are possible depending on whether the tube is long or short. In the analysis here the guide tube is taken to be an infinitely long tube with the maximum allowable ovality and minimum wall thickness. The conditions will result in the lowest critical pressure calcul ation for the guide tube (i.e., if the tube were short, the critical pressure calculation would give a higher number). The critical pressure is approximately 140 psi.
However, if the maximum allowable stress is reached at a pressure lower than the critical pressure, then that pressure is limiting. This is the case for a BWR guide tube. The allowable stress of 17,450 psi will be reached at approximately 93 psi. Comparing the maximum possible pressure differential for a steamline break to the limiting pressure of 93 psi gives a design margin greater than 3 to 1. Ther efore, the guide tube is not unstable with respect to differential pressure.
4.2.3.20.10 Evaluation of Control Rod Velocity Limiter
The control rod velocity limiter limits the free fall velocity of the control rod to a value that cannot result in nuclear system process barrier damage. This velocity is evaluated by the rod-drop accident analysis in Chapter 15.0.
4.2.3.21 Rod Bowing
4.2.3.21.1 GE Evaluation
Irradiation-induced bowing in fuel rods and assemblies is a phenomenon which is not, in itself, a failure mechanism. However, rod bowing must be addressed in the design analysis so as to establish operational tolerances. General Electric has indicated that boiling water reactor fuel operating experience, testing, and analysis indicate that there is no significant problem with rod bowing even at small rod-to-rod and rod-to-channel clearances. Specifically, General Electric noted that: (1) no gross bowing has been observed (excluding the rod bowing-related failures in an early design); (2) a very low frequency of minor bowing has been observed; (3) mechanical analysis indicates deflections within design bases; and (4) thermal-hydraulic testing has shown that small rod-to-rod and rod-to-channel clearances pose no significant problem. Based on those report observations and Reference 37, that address: (1) updates the General Electric rod bowing experience; (2) verifies the accuracy with which General Electric measures rod bowing; and (3) documents the overall General Electric rod bowing sa fety analyses, there is no reason to anticipate a problem with fuel rod or asse mbly bowing during operation of LaSalle.
LSCS-UFSAR 4.2-42 REV. 18, APRIL 2010 4.2.3.21.2 AREVA Evaluation
Differential expansion between the fuel rods, and lateral thermal and flux gradients can lead to lateral creep bow of the rods in the spans between sp acer grids. This lateral creep bow alters the pitch between the rods and may affect the peaking and local heat transfer. The AREVA design basis for fuel rod bowing is that lateral displacement of the fuel rods shall not be of sufficient magnitude to impact thermal margins. Extensive post-irridation examin ations have confirmed that such rod bow has not reduced spacing between adjacent rods by more than 50%, The potential effect of this bow on thermal margins is negligible. Rod bow at extended burnup does not affect thermal margins due to the lower powers achieved at high exposure (Reference 49).
4.2.3.22 Fission Gas Release
The information in this section is historical GE data on fuel reliability experience.
In 1976, the NRC had questioned the validity of fission gas release calculations in most fuel performance codes, including GEGAP - III (Reference 34), for a burnup greater than 20,000 megawatt days per ton of uranium. The General Electric Company was informed of this concern (Ref. 28) and was provided with a method of correcting fission gas release calculat ions for burnups greater than 20,000 megawatt days per ton of uranium (Ref. 29). Subsequently, the General Electric Company provided (Ref. 30) a generic reanal ysis of fuel performance calculations using GEGAP - III with the NRC's fission correction factor for BWR 2/3/4 plants with 7x7 and 8x8 fuel assemblies. Although the reanalysis was not specifically performed for the LaSalle fuel, a referenc ed 8x8 reanalysis performed for early refloodings plants bounded the LaSalle case. In the generic reanalysis, fuel rod internal pressure was shown to remain below system pressure for rod peak burnups below 40,000 megawatt days per ton of ur anium. This conclusion remains unchanged for the prepressurized fuel design (Ref. 31). The generic reanalysis did, however, result in higher initial stored energy and rupture pressure in the loss-of-coolant accident conditions, the higher fission gas release results in a maximum increase of 85 degrees Fahrenheit in calculated peak cladding temperature at end-of-life (approximately 33,000 megawatt days per tons of uranium planar average exposure). This added temperature increment results in calculated peak cladding temperatures of less than 2100 degrees Fahrenheit for average burnups below 33,000 megawatt days per ton of uranium and thus would not violate the 2200 degrees Fahrenheit loss-of coolant accident peak cladding temperature limit required by 10 CFR 50.46.
A full reanalysis of the effects of fission gas release prior to exceeding a peak local burnup of 20,000 megawatt days per ton of uranium was required by the NRC for LaSalle. General Electric proposed that credit for approved emergency core cooling system evaluation model changes be used to offset any detrimental effects of fission gas release at high burnups (Ref. 32). The proposal was accepted by the NRC LSCS-UFSAR 4.2-43 REV.
13 provided the more recent generic analysis was applicable to LaSalle. Per reference 33 CECo stated the latter generic analysis is applicable to LaSalle. The issue of enhanced fission gas release at high burnup is satisfactorily resolved for LaSalle.
4.2.3.23 Ballooning and Rupture 4.2.3.23.1 GE Evaluation
The information in this section is historical GE data on fuel reliability experience. In another loss-of-coolant accident related area of concern, the NRC had been generically evaluating three fuel material models that are used in emergency core cooling system evaluations. These models predict cladding rupture temperature, cladding burst strain, and fuel assembly flow blockage.
In a letter from L. O. DelGeorge to A. Schwencer dated May 21, 1981, CECo endorsed the results of a generic sensitivity study performed by General Electric submitted to the NRC by letter dated May 15, 1981. As reported in this generic study, General Electric has assessed the boiling water reactor emergency core cooling system sensitivity to rupture temperature by using three rupture temperature models:
(1) the General Electric CHAS TE model, (2) the NUREG-0630 model, and (3) a proposed General Electric model termed the adjusted model. For the LaSalle type of 8 x 8 with 2 water rod fuel design (designated the "improved 8 x 8 design"), General Electric found that the use of the NUREG-0630 model resulted in an increased peak cladding temperature of up to 50 degrees Fahrenheit over that which was obtained with the CHASTE model. However, sensitivity studies performed on the adjusted model, which is a combination of the CHASTE and
NUREG-0630 models and may be the bett er of the three models, found the maximum impact on peak cladding temperature to be 10 degrees Fahrenheit.
With regard to the boiling water reactor emergency core cooling system sensitivity to burst strain, the General Electric submittal assessed the impact of using a burst strain model that bounds the burst strain model given in NUREG-0630.
It is estimated from the impact (i.e., < 5 degrees Fahrenheit) of the reduced versus the CHASTE model comparison that if the comparison had been made against the unaltered NUREG-0630 strain model, the impact would have been < 115 degrees Fahrenheit. In light of the calculated 2009 degrees Fahrenheit loss-of-coolant accident peak cladding temperature for La Salle, sufficient margin exists between the 2200 degrees Farenheit peak cladding te mperature limit as required by 10 CFR 50.46 and the calculated 2009 degrees Fahrenheit LaSalle peak cladding temperature to accommodate an uncertainty of 115 degrees Fahrenheit in the peak cladding temperature.
LSCS-UFSAR 4.2-44 REV. 18, APRIL 2010 4.2.3.23.2 AREVA Evaluation During a severe loss of coolant accident, the cladding swelling and burst strain can result in flow blockage. Therefore, the LOCA analysis must consider the cladding swelling and burst strain impacts on the flow. AREVA uses the models in NUREG 0630 for cladding rupture. There is no explicit limit on the deformation. However, the calculations with the deformation models must satisfy the event criteria given in 10CFR 50.46. This swelling and rupture model is an integral part of the LOCA evaluation and is not part of the mechanical design analysis (Reference 49).
4.2.4 Testing
and Inspection Plan Rigid quality control requirements are enforced at every stage of fuel manufacturing to ensure that the design specifications are met. Written manufacturing procedures and quality control plans define the steps in the manufacturing process. The quality control plan is provided in Reference 43. Each fuel tube is subjected to dimensional inspection and ultrasonic inspection to reveal defects in the cladding wall. Destructive tests are performed on representative samples from each lot of tubing, including chemical analysis, tensile, and burst tests. Integrity of end plug welds is assured by standardization of weld processes based on radiographic and metallographic inspection of welds. Completed fuel rods are helium leak tested to detect the escape of helium through the tubes and end plugs or welded regions. The UO 2 powder characteristics and pellet densities, composition, and surface finish are controlled by regular sampling inspections.
The UO 2 weights are recorded at every stage in manufacturing. Dimensional measurements and visual inspections of critical areas, such as fuel rod-to-rod clearances, are performed after assembly.
Each separate pellet enrichment group has at times been characterized by a single stamp. Such a co ntrol has varied over time and varied among fuel vendors. Fuel rods are individually numbered prior to fuel loading: (a) to aid in identifying which pellet type is to be loaded in each fuel rod; (b) to aid in identifying which position in the fuel assembly each fuel rod is to be loaded; and (c) to facilitate total fuel material accountability for a given project.
Prior to introduction of AREVA fuel, further iden tification of individu al fuel rod gadolinia concentrations and uranium enrichments is accomplished by symbolization on the upper end plug shank for each differing rod. Each upper end plug is ensured proper placement on a fuel rod by reference to the specific fuel rod type. Each fuel rod is ensured of proper placement within a fuel bundle by inspection of the fuel rod serial number on the lower end plug or clad bar code. For AREVA fuel beginning with AREVA ATRIUM-10 fuel loaded into LaSalle 2 Cycle 10, fuel rod identi ty was tracked by use of a bar code on the cladding. This facilitates proper tracking at the fuel fabrication factory including proper loading into the fuel bundle skeleton through automated controls. Computer software ensures that the correct rods are loaded into the proper locations in the fuel bundle.
Fuel rod inspection includes metallographic and radiographic (not applicable to upset shape welded fuel rods) examination of fuel rods on a sample basis. Sample tests are performed for qualification of weld stations, weld parameters, and weld operators prior to application. Production samples are tested as a check on th e process and process controls.
LSCS-UFSAR 4.2-45 REV. 18, APRIL 2010 Fuel assembly inspections consist of complete dimensional checks of channels and fuel bundles prior to shipment. A sample of fuel bundles is given another visual and dimensional inspection of significant dimensions at the reactor site prior to use. Comparable tests and inspections are used by AREVA.
Onsite receipt of fuel rods and other reactor internals is the responsibility of EGC. General Electric and AREVA do provide recommendations to the purchaser for receipt, inspection, and handling of these components. General Electric and AREVA also perform audits to ensure that these activities are performed in compliance with General Electric and AREVA requirements. Such audits, however, are performed solely to satisfy General Electric and AREVA interests relative to warranty fulfillment.
The sampling rate and method of the site fu el receiving inspection are outlined in Table 4.2-5. However, current LaSalle fuel inspection procedures meet the requirements outlined in the fuel vendor's Quality Plan, which may or may not be the same as the sampling rate shown in Table 4.2-5.
Verification of enrichment and burnable poison concentrations is described in Subsection 4.2.4.1.
4.2.4.1 Testing and Inspection (Enrichm ent and Burnable Poison Concentrations)
The shutdown reactivity requirement is veri fied during initial fuel loading and at any time that core loading is changed. Nuclear limitations for control rod drives and SLC are verified by periodically testing the individual system.
The following serves to identify the various test and inspections employed by the Fuel Vendor(s) in verifying the nuclear characteristics of the fuel and reactivity control systems. Comparable tests and inspections are used by AREVA.
4.2.4.1.1 Enrichment Control Program GE uses emission spectrometry for determining impurities and mass spectrometry for verifying the U-235 enrichment in samples of UO 2 powder. AREVA verifies that samples of incoming UF 6 and the resultant UO 2 powder are within limits for impurities by emission spectroscopy. Th e U-235 content of a st atistical sample of UF 6 is verified by gamma counting and by mass spectroscopy measurement.
A sample of the sintered pellets is also checked for impurities by emission spectroscopy. AREVA performs chemical verification of impurities and O/U measurements on sintered pellets by emission spectroscopy, wet chemistry and LSCS-UFSAR 4.2-45a REV. 14, APRIL 2002 inert gas fusion. GE verifies the O/U ratio of UO 2 pellets and gadolinia bearing pellets up to 6 w/o Gd203 concentration by gravimetric methods. The O/U ratio for gadolinia bearing pellets with concentrat ion above 6 w/o Gd203 is confirmed using a spectrophotometric method. (G E uses emission spectrometry)
Each rod is gamma scanned to screen out any possible but unlikely misplaced pellet or enrichment deviations.
LSCS-UFSAR 4.2-46 REV. 18, APRIL 2010 4.2.4.1.2 Gadolinia Inspections The same rigid quality control requirements observed for standard UO 2 fuel are employed in manufacturing gadolinia-urania fuel. Gadolinia-bearing UO 2 fuel pellets of a given enrichment and gadolinia concentration are maintained in separate groups throughout the manufacturi ng process. For General Electric, the percent enrichment and gadolinia concentration characterizing a pellet group are identified by a stamp on the pellet. ForAREVA, gadolinia pellets are uniquely identified with a symbol stamped on the pellet.
Fuel rods are individually numbered prior to loading of fuel pellets into the fuel rods: (1) to identify which pellet group is loaded in each fuel rod; (2) to identify which position in the fuel assembly each fuel rod is load ed; and (3) to facilitate total material accountability for a given project. The correct location of all fuel rods in the bundle is ensured through the use of a computer-controlled, automated bundle
assembly machine.
The following quality control inspections are made:
- a. Gadolinia concentration in the gadolinia-urania powder blend is verified.
- b. Sintered pellet UO 2-Gd 2 O 3 solid-solution homogeneity across a fuel pellet is verified by examination of ceramographic
specimens.
- c. Gadolinia-urania pellet identification is verified.
- d. Gadolinia-urania fuel rod identification is checked.
- e. Each gadolinia - urania fuel rod is scanned to assure proper assembly.
- f. Gadolinia content is verified by X-ray fluorescence measurements of each pellet or scanning the assembled rod.
All assemblies and rods of a given project are inspected to ensure overall accountability of fuel quantity and placement for the project.
AREVA uses similar practices and techniques for gadolinia inspection.
LSCS-UFSAR 4.2-47 REV. 13 4.2.4.1.3 Reactor Control Rods Inspections and tests are conducted at various points during the manufacture of control rod assemblies to ensure that design requirements are being met. All boron carbide lots are analyzed and certified by the supplier. Among the items tested are:
- d. particle size distribution.
Following receipt of the boron carbide and review of material certificates, additional samples from each lot are tested including those previously listed. Control is maintained on the B 4C powder through the remaining steps prior to loading into the absorber rod tubes.
Certified test results are obtained on other control rod components. The absorber rod tubing is subjected to extensive test ing by the tubing supplier, including 100% ultrasonic examination. Metallographic examinations are conducted on several tubes randomly selected from each lot to verify cleanliness and absence of conditions resulting from improper fabricat ion, cleaning or heat treatment. Other checks are made on the subassemblies and final control rod assembly, including weld joints inspected and B 4 C loading.
4.2.4.2 Surveillance Inspection an d Testing of Irradiated Fuel Rods
General Electric has a cooperative program of surveillance of BWR fuel, both production and developmental, which operates beyond current production fuel experience as it becomes available for inspection. The schedule of inspection is, of course, contingent on the availability of the fuel as influenced by plant operation.
This program is provided in Reference 41.
The lead experience fuel rods (with respect to exposure, linear heat generation rate, and the combination of both) are selectively inspected. Inspection techniques used include: a. leak detection tests, such as "sipping;"
- b. visual inspection with various aids such as binoculars, borescope, periscope, and/or underwater TV with a photographic record of observations as appropriate; LSCS-UFSAR 4.2-48 REV. 13 c. nondestructive testing of selected fuel rods by ultrasonic and eddy current test techniques; and
- d. dimensional measurements of selected fuel rods.
Unexpected conditions or abnormalities which may arise, such as distortions, cladding perforation, or surface disturbance s are analyzed. Resolution of specific technical questions indicated by site examinations may require examination of selected fuel rods in the Radioactive Material Laboratory (RML) facilities.
The results of the program are used to evaluate the boiling water reactor fuel design methods and criteria used by General Electric.
The results of the surveillanc e program are generally review ed with the Division of Reactor Licensing and documented in generic fuel experience licensing topical
reports. Historical fuel performance results prior to 1979 on highly precharacterized lead test assemblies are provided in several report s listed in Reference 38. The lead test assemblies are utilized as one means of providing some confirmation of design adequacy or early warning of negative features of the design. Details on lead test assembly programs are pr ovided in Reference 39.
In addition to fuel bundle inspection, the fuel channels are under surveillance in continuing programs. These surveillance programs are designed not only for the evaluation of present day products, but are also providing data in the areas of alternate materials and design modeling.
4.2.4.3 Operating Experience with Gadolinia-Containing Fuel
Production fuel rods employing gadolinia-urania fuel pellets have been in use since 1965. During this time, a substantial num ber of gadolina-urania rods have been successfully irradiated to appreciable exposures. Additional information on gadolinia-urania physical and irradiation characteristics, material properties, and operating experiences is provided in Reference 25.
Temperature coefficients are virtually unchanged because of gadolinia. The gadolinia-bearing pellets act as thermally gr ay or black adsorbers, and their effect on moderator coefficients in the lattice is not essentially different from that of the control which they replace. Doppler response is unaffected because the gadolinia has essentially no effect on the resonance group flux or on the U-238 content of the core. The concentration of gadolinia has been selected so that the initial concentration of the high cross section isotopes, Gd-155 and
-157, will be completely depleted by the end of the first cycle. The irradiation pr oducts of this process are other gadolinia LSCS-UFSAR 4.2-49 REV. 14, APRIL 2002 isotopes having low cross sections. Power in the gadolinium pins generally remains below 90% of the average bundle power. The control augmentation effect disappears on a predetermined schedule without changes in the chemical composition of the fuel or the physical makeup of the core.
The thermal margins described by the steady-state operating limits (LHGR, APLHGR and MCPR) are easily maintained in a gadolinia core because additional power shaping is possible through spatial va riation of the burnable poison loading. The damage limits on gadolinia-urania fuel rods are designed with similar margins as maintained for the UO 2 rods. 4.2.5 References
- 1. NEDO-20948-P, "BWR/6 Fuel Design," December 1975.
- 2. WAPD-TM-283, "Effects of High Burnup on Zircaloy-clad, Bulk UO 2 Plate Fuel Element Samples," September 1962.
- 3. WAPD-TM-629, "Irradiation Behavior of Zircaloy-Clad Fuel Rods Containing Dished End UO 2 Pellets," July 1967.
- 4. H. E. Williamson and D. C. Di tmore, "Experience with BWR Fuel Through September 1971," NEDO-10505, May 1972.
- 5. D. C. Ditmore and R. B. Elkins, "Densification Considerations in BWR Fuel Design and Performance," NEDM-10735, December 1972.
- 6. J. A. Christensen, "Melting Point of Irradiated Uranium Dioxide," WACP-6065, February 1965.
- 7. "Thermal Conductivity of Uranium Dioxide," Technical Report series No. 59, IACA, Vienna, 1966.
- 8. Supplement 1 to the Technical Re port on Densification of General Electric Reactor Fuels, December 1973.
- 9. This reference has been deleted.
- 10. J. P. Fritz, "Testing of Cruciform Control Rods for BWR/6" NEDO-10565, April 1972.
- 11. R. J. Benche, "Visual and Photographic Examination of Dresden-1 High Exposure Control Rod B-87" NEDO-10541, April 1972.
LSCS-UFSAR 4.2-50 REV. 19, APRIL 2012 12. W. F. O'Donnel and B. F. Langer, "Fatigue Design Basis for Zircaloy Components," Nuclear Science and Engineering, Vol. 20, pp. 1-12, 1964. 13. E. P. Quinn, "Vibration of Fu el Rods in Parallel Flow," GEAP-4059, July 1962.
- 14. NEDO-20360-1P, Revision 4, "General Electric Boiling Water Reactor Generic Reload Application for 8x8 Fuel," March 1976.
- 15. "Design Safety Standards for Boiling Water Reactors," complied by Safety and Standards Unit, NEDE-10370, June 1971.
- 16. Design Analysis No. L-003508, Rev.
0, "LaSalle Lost Parts Analysis," July 2010.
- 17. "BWR Fuel Channel Mechanical Design and Deflection," NEDO-21354, September 1976.
- 18. This reference has been deleted.
- 19. R. B. Elkins, "Experience with BWR Fuel Through September 1974," NEDO-20922, June 1975.
- 20. H. E. Williamson and D. C. Ditmore, "Current State of Knowledge High Performance BWR Zircaloy-Clad UO 2 Fuel," NEDO-10173, May 1970.
- 21. "Thermal Response and Cladding Performance of an Internally Pressurized, Zircaloy Clad, Simulated BWR Fuel Bundle Cooled by
Spray Under Loss-of-Coolant Conditions," GEAP-13112, April 1971.
- 22. L. A. Stephan, "The Response of Waterlogged UO 2 Fuel Rods to Power Burst," IDO-ITR-105, April 1969.
- 23. L. A. Stephan, "The Effects of Cladding Material and Heat Treatment on the Response of Waterlogged UO 2 Fuel Rods to Power Bursts," IN-ITR-111, January 1970.
- 24. "Consequences of a Postulated Flow Blockage Incident in a Boiling Water Reactor," NEDO-10174, Re vision 1, October 1977.
- 25. G. A. Potts, "Urania-Gadolinia Nuclear Fuel Physical and Irradiation Characteristics and Material Proper ties," NEDE-20943 (proprietary),
NEDO-20943 (nonproprietary), January 1977.
LSCS-UFSAR 4.2-51 REV. 13 26. "Creep Collapse Analysis of BWR Fuel Using Safe-Collapse Model," NEDE-20606 (Proprietary), NEDO-20606A (Non-Proprietary) August 1976. 27. This reference has been deleted.
- 28. D. F. Ross (NRC) Letter to G.G. Sherwood (GE) dated November 23, 1976.
- 29. NUREG-0418, "Fission Gas Release from Fuel at High Burnup," March, 1978.
- 30. G.G. Sherwood (GE) Letter to D.F. Ross (NRC) dated December 22, 1976.
- 31. General Electric report, NEDO - 23786-1, "Fuel Rod Pressurization -
Amendment 1," May 1978.
- 32. Letters from R. E. Engel to T.A. Ippolito dated May 6, 1981 and May 23, 1981.
- 33. L. O Del George letter to A. Schwencer (NRC) dated September 21, 1981.
- 34. General Electric topical report NEDO-20181, "GEGAP-III: A Model for the Predictions of Pellet - Cladding Thermal Conductance in BWR Fuel Rods," November 1973
- 35. Letter from W.R. Butler (NRC) To I. Stuart (GE), dated April 4, 1975.
- 36. General Electric report, NEDO - 21156, "Supplemental Information for Plant Modification's to Eliminate Significant In-core Vibration,"
January, 1976.
- 37. General Electric topical report, NEDE - 24284, "Fuel Rod Bowing in General Electric Boiling Water Reactors," Dated August 1980.
- 38. NUREG 0633, "Fuel Performance Annual Report," December 1979.
- 39. General Electric report, NEDC 24609, "Boiling Water Reactor Fuel Rod Performance Evaluation Program," February 1979.
- 40. General Electric report NEDE , 21660-P, "Experience with BWR Fuel through December 1976," July 1977.
LSCS-UFSAR 4.2-52 REV. 18, APRIL 2010 41. General Electric report NEDE-2 4011-P-A, "General Electric Standard Application For Reactor Fuel (GESTAR II)", (latest approved revision).
- 42. Letter Number MFN-078-086/JSC-067-08 6, J. S. Channley (GE) to C. H. Berlinger (NRC), "1985 Fuel Experi ence Report," August 13, 1986.
- 43. "Nuclear Energy Business Group BWR Quality Assurance Program Description," NEDO-11209-04A, March 1978.
- 44. NEDE-31152, "GE Fuel Bundle Designs," (latest revision).
- 45. General Electric Report GE-NE-523-A191-1294, "Final Report of the Impact of Using GE9 80-Mil Fuel Channel for the LaSalle Units 1 and 2", latest revision and supplements.
- 46. Advanced Nuclear Fuels Corporation Generic Mechanical Design for Advanced Nuclear Fuels 9x9-IX and 9x9-9X Reload Fuel, ANF-89-014(A), Revision 1 and Suppl ements 1-2, Advanced Nuclear Fuels Corporation, Richland, WA, October 1991.
- 47. Generic Mechanical Design for Exxon Nuclear Jet Pump BWR Reload Fuel, XN-NF-85-67(A), Exxon Nuclear Fuels Corporation, September 1986.
- 48. AREVA document, Cycle Specific Fuel Design Report.
- 49. Generic Mechanical Design Criteria for BWR Fuel Designs, ANF-89-098(P)(A), Revision 1, Suppl ement 1, Advanced Nuclear Fuels Corporation, May 1995.
- 50. RODEX2A (BWR) Fuel Rod Thermal-Mechanical Evaluation Model, XN-NF-85-74(P)(A), Exxon Nuclear Company, August 1986.
- 51. RODEX2 Fuel Rod Thermal-Mechanical Response Evaluation Model, XN-NF-81-58(P)(A) Supplements 1 and 2 Revision 2, Exxon Nuclear Company, May 1986.
- 52. General Electric Report NEDE-22290 Su pplement 2, Production of Advanced Longer Life Control Rod in BWR 2-6 Plants.
- 53. General Electric Report NEDE-3 1578PA, Class III, GE Marathon Control Rod Assembly, October 1991.
- 54. General Electric Report NLM-CR-4505, Revision C, Class III, Tubricast Velocity Limits for Dura life and Marathon Control Blades, September 1998.
LSCS-UFSAR 4.2-52a REV. 20, APRIL 2014
- 55. Fuel Design Evaluation for Siemens Power Corporation ATRIUM-10 BWR Reload Fuel, EMF-95-52(P), Revision 2, December 1998.
- 56. Mechanical Design Evaluation for ATRIUM-10 BWR Reload Fuel to 54 MWd/KgU Assembly Exposure, EMF-98-006(P), January 1998.
- 57. Mechanical Design for BWR Fuel Channels, EMF-93-177(P)(A) Revision 1 and Supplement 1, August 2005.
- 58. GE14 Compliance with Amendment 22 of NEDE-24011-P-A (GESTAR II), NEDC - 32868P, latest approved revision.
- 59. Design Analysis L-003664, Revision 1, "GNF2 Advantage Generic Compliance with NEDE-24011-P-A (GESTAR II), GEH Report NEDC-33270P" 60. Design Analysis L-003697, Revision 0, "GNF S-0000-0142-0455, Analysis of Flow Blockage Consequences for LaSalle Units 1 and 2 GNF2 New Fuel Introduction" LSCS-UFSAR TABLE 4.2-1 TABLE 4.2-1 REV. 13 TYPICAL LIMITING LHGR'S FOR GADOLINIA-URANIA FUEL RODS (kW/ft)
EXPOSURE (MWd/tU)
INCIPIENT CENTER MELTING 1% PLASTIC STRAIN
OF CLADDING EXPECTED OPERATING MAXIMUM (4 wt% Gd 2 0 3) 0 18.4 23.0 ~ 4 20,000 17.8 21.4 ~ 11 40,000 16.7 18.2 ~ 8
LSCS-UFSAR TABLE 4.2-2a. TABLE 4.2-2 REV. 13 General Electric STRESS INTENSITY LIMITS YIELD STRENGTH (S y) ULTIMATE TENSILE STRENGTH (S u) Primary membrane stress 2/3 1/2 Primary membrane plus bending stress intensity 1 1/2 to 3/4 Primary plug secondary stress
intensity 2 1.0 to 1.5
LSCS-UFSAR TABLE 4.2-2b TABLE 4.2-2b REV. 15, APRIL 2004 FANP STRESS INTENSITY LIMITS*
Stress Intensity Limits**
YIELD STRENGTH (y) ULTIMATE TENSILE STRENGTH (u) Primary membrane stress 2/3 y 1/2 y Primary membrane plus
bending stress intensity 1.0 y 1/2 u Primary plug secondary stress
intensity 2.0 y 1.0 u
- Characteristics of the stress categories are defines as follows:
a) Primary stress is a stress developed by the imposed loading which is necessary to satisfy the laws of equilibrium between external and internal forces and moments. The basic characteristics of a primary stress is that it is not self-limiting. If a primary stress exceeds the yield strength of the material through the entire thickness, the pr evention of failure is entirely dependent on the strain-hardening properties of the material.
b) Secondary stress is a stress developed by the self-constraint of a structure. It must satisfy an imposed strain pattern rather than being in equilibrium with an external load. The basic characteristic of a secondary stress is that it is self-limiting. Local yielding and minor distortions can satisfy the discontinuity conditions due to thermal expansions which cause the stress to occur.
- The stress intensity is defined as twice the maximum shear stress and is equal to the largest algebraic difference between any two of the three principal stresses.
LSCS-UFSAR TABLE 4.2-3 TABLE 4.2-3 REV. 0 - APRIL 1984 CONDITIONS OF DESIGN RESULTING FROM IN-REACTOR PROCESS CONDITIONS COMBINED WITH EARTHQUAKE LOADING
CONDITIONS OF DESIGN REACTOR INITIAL CONDITIONS PERCENT OF SAFE SHUTDOWN EARTHQUAKE IMPOSED 0% 50% 100% Startup Testing Upset -- -- Normal Normal Upset Faulted Abnormal Upset -- --
LSCS-UFSAR TABLE 4.2-4(a)
TABLE 4.2-4(a) REV. 15, APRIL 2004 DATA FOR THE 8x8R FUEL DESIGN Core (Full Core Data) Fuel cell spacing (control rod pitch), in. 12Number of fuel assemblies 764 Total number of fueled rods 47368Core power density (rated power), kW/l 50.0 Total core heat transfer area, ft 2 74872 Fuel Assembly Data Overall length, in.
176Nominal active fuel length, in.
150Fuel rod pitch, in.
0.640 Space between fuel rods, in.
0.157Fuel channel wall thickness, in.
0.100 Fuel bundle heat transfer area, ft 2 98.0Channel width (inside), in.
5.278 Fuel Rod Data Outside diameter, in.
0.483 Cladding inside diameter, in.
0.419Cladding thickness, in.
0.032Fission gas plenum length, in.
10.0Pellet immersion density, % T.D.
95 Pellet outside diameter, in.
0.410 Pellet length, in.
0.410 Water Rod Data Outside diameter, in.
0.591 Inside diameter, in.
0.531 LSCS-UFSAR TABLE 4.2-4(b)
TABLE 4.2-4(b) REV. 15, APRIL 2004 DATA FOR THE GE8x8NB (GE9B) FUEL DESIGN Core (Full Core Data) Fuel cell spacing (control rod pitch), in.
Number of fuel assemblies Total number of fueled rods Core power density (rated power), kW/l Total core heat transfer area, ft 2 12 764 45840 50.0 approximately 71816 Fuel Assembly Data Nominal active fuel length, in. Fuel rod pitch, in.
Space between fuel rods, in. Fuel channel wall thickness, in.
Fuel bundle heat transfer area, ft 2 Channel width (inside), in.
150 0.640 0.157 0.100* approximately 94 5.278 Fuel Rod Data Outside diameter, in. Cladding inside diameter, in. Cladding thickness, in. Pellet immersion density, % T.D.
Pellet outside diameter, in.
Pellet length, in.
0.486 0.419 0.032 96.5 0.411 0.410 Water Rod Data Outside diameter, in.
Inside diameter, in.
1.340 1.260
- Either 100 or 80 mil channels are used, depending on the reload.
LSCS-UFSAR TABLE 4.2-4(c) TABLE 4.2-4(c) REV. 15, APRIL 2004 DATA FOR THE FANP ATRIUM-9B FUEL DESIGN Core (Full Core Data) Fuel cell spacing (control rod pitch), in. 12Number of fuel assemblies 764Total number of fueled rods 55008Core power density (rated power), kW/l 50.0
Total core heat transfer area, ft 2 77426 Fuel Assembly Data Overall length, in.
176Nominal active fuel length, in.
149Fuel rod pitch, in.
0.569 Space between fuel rods, in.
0.136Fuel channel wall thickness, in.
0.08*Fuel bundle heat transfer area, ft 2 101.343Channel width (inside), in.
5.278 Fuel Rod Data Outside diameter, in.
0.433 Cladding inside diameter, in.
0.3807Cladding thickness, in.
0.026Fission gas plenum length, in.
10.578Pellet immersion density, % T.D.
96 Pellet outside diameter, in.
0.3737 Pellet length, in.
Enriched, in. Natural, in. 0.393 0.545 Water Box Data Outside dimension, in.
1.516 Water box wall thickness, 0.0285
- Either 100 or 80 mil channels ar e used, depending on the reload.
LSCS-UFSAR TABLE 4.2-4(d) TABLE 4.2-4(d) REV. 17, APRIL 2008 DATA FOR THE AREVA ATRIUM-10 FUEL DESIGN Core (Full Core Data) Fuel cell spacing (control rod pitch), in. 12Number of fuel assemblies 764 Total number of fueled rods 69524Core power density (rated power), kW/l 52.0 Total core heat transfer area, ft 2 86315 Fuel Assembly Data Overall length, in.
176.386Nominal active fuel length, in.
- Full length fuel rods 149* Part length fuel rods 90Fuel rod pitch, in.
0.510 Space between fuel rods, in.
0.114Fuel channel wall thickness, in.
0.100 Fuel bundle heat transfer area, ft 2 113.0Channel width (inside), in.
5.278 Fuel Rod Data Outside diameter, in.
0.3957 Cladding inside diameter, in.
0.3480Cladding thickness, in.
0.024Fission gas plenum length, in.
- Full length fuel rod 11.52(TIG)/ 11.53 (USW)
- Part length fuel rod 5.26(TIG)/ 5.42 (USW) Pellet immersion density, % T.D. (typical, pellet enrichment dependent) 96.26Pellet outside diameter, in. 0.3413 Pellet length, in. Enriched, in. Natural, in.
0.413 0.551 Water Box Data Outside dimension, in. 1.378Water box wall thickness, 0.0285 LSCS-UFSAR TABLE 4.2-4(e)
TABLE 4.2-4(e) REV. 16, APRIL 2006 DATA FOR THE GE14 FUEL DESIGN Core (Full Core Data)
Fuel cell spacing (control rod pitch), in.
12 Number of fuel assemblies 764 Total number of fueled rods 70288 92*764Core power density (rated power), kw/l 53.01 NEDE 31152P Rev8 Total core heat transfer area, ft 2 86332 113*764 Fuel Assembly Data Nominal active fuel length, in.
- Full length fuel rods 150 GE Dwg 217C1442
- Part length fuel rods 84 GE Dwg 217C1444Fuel rod pitch, in. 0.510 NEDE 31152P Rev 8 Space between fuel rods, in. 0.106 NFM DIR-00-081, Nov 30, 2000, GE14 Design ReviewFuel channel wall thickness (corner/median), in. 0.120/0.075 NEDE 31152P Rev 8 Fuel bundle heat transfer area, ft 2 113 NEDE 31152P Rev 8Channel width (inside), in. 5.278 NEDE 31152P Rev 8 Fuel Rod Data Outside diameter, in. 0.404 NEDE 31152P Rev 8Cladding inside diameter, in. 0.352 NEDE 31152P Rev 8Cladding thickness, in. 0.026 NEDE 31152P Rev 8Fission gas plenum length, in.
- Full length fuel rod 9.64 GE Dwg 217C1442
- Part length fuel rod 10.94 GE Dwg 217C1444Pellet immersion density, %T.D. (typical, pellet enrichment dependent) 97.0 NEDE 31152P Rev 8 Pellet outside diameter (cold), in. 0.345 GE Dwg 137C9061 Pellet length, in. 0.370 GE Dwg 137C9061 Water Rod Data Outside diameter, in. 0.980 NEDE 31152P Rev 8Inside diameter, in. 0.920 NEDE 31152P Rev 8 LSCS-UFSAR TABLE 4.2-4(f)
TABLE 4.2-4(f) REV. 20, APRIL 2014 DATA FOR THE GE14 FUEL DESIGN Core (Full Core Data)
Fuel cell spacing (control rod pitch), in.
12 Number of fuel assemblies 764 Total number of fueled rods 70288 92*764 Core power density (rated power), kw/l 52.5 Total core heat transfer area, ft2 86332 113*764 Fuel Assembly Data Nominal active fuel length, in.
- Full length fuel rods 150 GE Dwg 217C1442
- Part length fuel rods 84 GE Dwg 217C1444 Fuel rod pitch, in.
0.510 NEDE 31152P Rev 8 Space between fuel rods, in.
0.106 NFM DIR-00-081, Nov 30, 2000, GE14 Design Review Fuel channel wall thickness (corner/median), in.
0.120/0.075 NEDE 31152P Rev 8 Fuel bundle heat transfer area, ft2 113 NEDE 31152P Rev 8 Channel width (inside), in.
5.278 NEDE 31152P Rev 8 Fuel Rod Data Outside diameter, in.
0.404 NEDE 31152P Rev 8 Cladding inside diameter, in.
0.352 NEDE 31152P Rev 8 Cladding thickness, in.
0.026 NEDE 31152P Rev 8 Fission gas plenum length, in.
- Full length fuel rod 9.64 GE Dwg 217C1442
- Part length fuel rod 10.94 GE Dwg 217C1444 Pellet immersion density, %T.D. (typical, pellet enrichment dependent) 97.0 NEDE 31152P Rev 8 Pellet outside diameter (cold), in.
0.345 GE Dwg 137C9061 Pellet length, in.
0.370 GE Dwg 137C9061 Water Rod Data Outside diameter, in.
0.980 NEDE 31152P Rev 8 Inside diameter, in.
0.920 NEDE 31152P Rev 8
LSCS-UFSAR TABLE 4.2-5 (SHEET 1 OF2) TABLE 4.2-5 REV. 13 SITE FUEL RECEIVING INSPECTION
- , ** FUEL INSPECTION OBJECTIVES CHARACTERISTIC INTENDED METHOD E XPECTED FREQUENC Y Container Damage and Leak Visual 100% Bundle Damage Visual 100% Shipping Separators Visual 100%
Removed Cleanliness Visual 100%
Rod Integrity Visual, gauge when required 100% Lock Tab Washers Visual 100% Channel Integrity Visual 100%
Channel Cleanliness Visual 100%
Guard Integrity and Installation Visual and Torque Wrench 100% Spacer Damage Visual 100% for first 5
bundles and every
20th thereafter, otherwise the middle 3 spacers. Rod to Rod Feeler gauge 100% of first 5 bundles
and every 20th thereafter, otherwise two sections, all spaces, alternate the sections. Rod-to-Simulated
Channel Simulated Channel and Feeler Gauge 100% of first 5 bundles
and every 20th thereafter, otherwise 2 sections, 4 sides per section, alternate sections excluding end sections.
LSCS-UFSAR TABLE 4.2-5 (SHEET 2 OF 2) TABLE 4.2-5 REV. 15, APRIL 2004 CHARACTERISTIC INTENDED METHOD EXPECTED FREQUE NSpring Length Visual 100% for all bundles Gauge 100% for first 5
bundles and every fourth thereafter, otherwise visual inspection.
Finger Spring Seated in Pocket Visual 100% for all bundles Gauge 100% for first 5
bundles and every
fourth thereafter, otherwise visual inspection.
NOTE: Deviations require 100% inspection of the next 5 bundles for that characteristic. Two deviations for a characteristic within 6 consecutive bundles require revision of the AQL (acceptable quality level) with the General El ectric, Wilmington, North Carolina, U.S.A., facility.
Where a reduced inspection was performed, a ll inspection steps shall be designated S OK (stamped OK).
- Current LaSalle fuel inspection procedures meet the requirements outlined in the fuel vendor's Quality plan, which may or may not be the same as the sampling rate in Table 4.2-5.
- These inspection objectives are specific to GE fuel. FANP fuel has similar inspection objectives for the FANP designs.
LSCS-UFSAR 4.4-1 REV. 14, APRIL 2002
4.4 THERMAL
AND HYDRAULIC DESIGN
4.4.1 Design
Bases
4.4.1.1 Safety Design Bases
Thermal hydraulic design of the LaSalle County Station (LSCS) core is established
and based upon the following design bases:
- a. Actuation limits for the devices of the nuclear safety systems are employed such that no fuel damage occurs as a result of
abnormal transients (Chapter 15.0). Specifically, the minimum
critical power ratio (MCPR) operating limit is specified such that
at least 99.9% of the fuel rods in the core will not experience
boiling transition during the most severe abnormal operational
transient. A 1% plastic strain limit is specified to ensure that
clad overstraining does not occur. b. Thermal hydraulic safety limits are used in setting safety margins and the consequences of fuel barrier failure to public
safety. c. The nuclear system must meet the requirements in 10CFR50, Appendix A, General Design Criterion 12 - Suppression of
Reactor Power Oscillations.
4.4.1.2 Power Generation Design Bases
The thermal-hydraulic design of the core provides the following operational
characteristics:
- a. ability to achieve rated core power output throughout the design life of the fuel without sustaining premature fuel failure, and b. flexibility to adjust core output over the range of plant load and load maneuvering requirements in a stable, predictable manner
without sustaining fuel damage.
4.4.1.3 Requirements for Steady-State Conditions
Steady-State Limits
For purposes of maintaining adequate thermal-hydraulic margin during normal steady-state operation, the minimum critical power ratio must not be less than the
required MCPR operating limit, the operational linear heat generation rate (LHGR)
is maintained below the LHGR limit for the fuel type, and the maximum average LSCS-UFSAR 4.4-2 REV. 14, APRIL 2002 planar linear heat generation rate (M APLHGR) must be maintained below the limits for the plant. This does not specify the operating power nor does it specify peaking factors. These parameters are determined subject to a number of
constraints, including the thermal limits given previously. The core and fuel
thermal-hydraulic design basis for steady-state operation, has been defined to provide margin between the steady-state operating condition and any fuel damage
condition to accommodate uncertainties and to ensure that no fuel damage results, even during the worst anticipated transient conditions at any time in life. For
LSCS, the operating limits for all three fuel thermal design limits are contained in the Core Operating Limits Report.
4.4.1.4 Requirements for Transient Conditions
Transient Limits
The transient thermal-hydraulic limits are established such that no fuel damage is
expected to occur during the most severe abnormal operating transient. Fuel
damage is defined as perforation of the cladding that permits release of fission
products (Section 4.2). Mechanisms that cause fuel damage in reactor transients
are:
- a. severe overheating of fuel cladding caused by inadequate cooling, and b. fracture of the fuel cladding caused by relative expansion of the uranium dioxide pellet inside the fuel cladding.
For design purposes, the transient thermal-hydraulic limit requirement is met if at
least 99.9% of the fuel rods in the core do not experience boiling transition during
any abnormal operating transient. No fuel damage is expected to occur even if a
fuel rod actually experiences a boiling transition.
A value of 1% plastic strain of Zircaloy cladding has been established as the limit
below which fuel damage from overstrainin g the fuel cladding is not expected to occur. Available data indicate that the threshold for damage is in excess of this value. The linear heat generation rate required to cause this amount of cladding
strain decreases with burnup.
4.4.1.5 Summary of Design Bases
In summary, the steady-state thermal-hydraulic operating limits have been
established to ensure that the design basis is satisfied for the most severe abnormal
operational situation, whether a transient or an accident. Transient analyses are
performed that demonstrate compliance with overpower transient limits assuming
steady-state operation has been in compliance with steady state operating limits.
An overpower which occurs during an abnormal operational transient must not LSCS-UFSAR 4.4-3 REV. 14, APRIL 2002 result in violation of the MCPR safety limit for the plant. Demonstration that the transient limits are not exceeded is sufficient to conclude that the thermal hydraulic
design basis is satisfied.
The MCPR, LHGR and MAPLHGR limits are su fficiently general so that no other limits need to be stated. The cladding and fuel bundle integrity criterion is assured as
long as MCPR, LHGR and MAPLHGR limits are met. There are no additional design criteria on coolant void fraction, core coolant flow-velocities, or flow distribution, nor are they needed. Core design and target rod patterns ensure CPRs remain above the
MCPR limits, thereby ensuring bundle para meters (e.g., flow, power, void fraction) remain within prescribed ranges. The coolant flow velocities and void fraction become
constraints upon the mechanical and physics design of reactor components and are
partially constrained by stability and control requirements.
4.4.1.5.1 Fuel Cladding Integrity The fuel cladding integrity is defined in Subsection 4.2.1. The fuel cladding integrity
from a thermal hydraulic viewpoint is assured by the operating and transient MCPR
requirements.
4.4.1.5.2 Fuel Assembly Integrity The fuel channel provides adequate lateral structural support for the fuel bundle and protects the fuel rods and spacers from impact and abrasion. The upper tie-plate
handle is capable of supporting the weight of the fuel assembly. Specific design characteristics are given in Section 4.2.
4.4.1.5.3 Fuel-Cladding Gap Characteristics
The subject of fuel to cladding gap characteristics is covered in Section 4.2.
4.4.2 Description
of Thermal Hydraulic Design of Reactor Core 4.4.2.1 Summary Comparison An evaluation of plant performance from a thermal and hydraulic standpoint is
provided in Subsection 4.4.4.
Transient evaluations are given in Chapter 15. A tabulation of thermal and hydraulic
parameters of the LSCS reactor initial core, along with a comparison to the initial core
of other reactors of a similar design, are given in Table 4.4-1.
4.4.2.2 Critical Power Ratio There are three different types of boiling heat transfer in water forced convection
systems: nucleate boiling, transition boiling, and film boiling. Nucleate boiling, at
lower heat transfer rate, is an extremely efficient mode of heat transfer, allowing LSCS-UFSAR 4.4-4 REV. 18, APRIL 2010 large quantities of heat to be transferred with a very small temperature rise at the heated wall. As heat transfer rate is increased the boiling heat transfer surface
alternates between film and nucleate boiling, leading to fluctuations in heated wall
temperatures. The point of departure from the nucleate boiling region into the
transition boiling region is called the boiling transition. Transition boiling begins at
the critical power, and is characterized by fluctuations in cladding surface
temperature. Film boiling occurs at the highest heat transfer rates; it begins as
transition boiling comes to an end. Film boiling heat transfer is characterized by
stable wall temperatures which are higher than those experienced during nucleate
boiling.
4.4.2.2.1 Boiling Correlations
4.4.2.2.1.1 GE Fuel
The occurrence of boiling transition is a function of the local steam quality, boiling
length, mass flow rate, pressure, flow geom etry, and local peaking pattern. General Electric has conducted extensive experimental investigations of these parameters.
These parametric studies encompass the enti re design range of these variables. In the experimental investigations, a boiling transition event was associated with a
25º F rise in rod surface temperature. The (critical) quality at which boiling
transition occurs as a function of the distance from the equilibrium boiling
boundary is predicted by the GEXL (G eneral E lectric Critical Quality X - Boiling Length) correlation. This correlation is based on accurate test data of full-prototype simulations of reactor fuel assemblies operating under conditions duplicating those in actual reactor designs. The GEXL correlation is a best fit to the data and is used
together with a statistical analysis to assure adequate reactor thermal margins (References 1 and 11).
The figure of merit used for reactor design and operation is the critical power ratio (CPR). This is defined as the ratio of the bundle power which would produce
equilibrium quality equal to but not exceeding the correlation value (critical
quality), to the bundle power at the reactor condition of interest (i.e., the ratio of
critical bundle power to operating bundle power). In this definition, the critical
power is determined at the same mass flux, inlet temperature, and pressure which
exist at the specified reactor condition.
The core is sized with sufficient coolant flow to assure that the MCPR is maintained
greater than the operating limit at rated conditions.
4.4.2.2.1.2 AREVA Fuel
In the AREVA methodology, the fuel assembly critical power corresponding to a particular reactor operating state is determined from the SPCB (References 25
and 26) or ANF-B (Reference 17) critical power correlations.
LSCS-UFSAR 4.4-5 REV. 18, APRIL 2010 The ANF-B and SPCB correlations provides a generic tool for evaluating critical power and to assess thermal margin for all current domestic AREVA BWR fuel designs. It is based on a data base characteristic of AREVA product designs. The database contains data for AREVA fuel designs with both axially uniform and nonuniform power profiles.
The ANF-B and SPCB critical power correlations are an empirical representation of
planar average thermal-hydraulic fluid conditions at which boiling transition has
been experimentally determined. The minimum heat flux required to produce
boiling transition is predicted from fluid conditions of pressure, mass velocity, and
enthalpy averaged over the plane of interest. The correlation contains correction
factors for the effects of boiling transition due to a nonuniform axial heat flux
profile and the grouping of relatively high-powered rods.
The test assemblies include full-length rods; typical BWR grid spacers; 4x4, 5x5, and 9x9 rod configurations; and a variety of rod diameters, assembly hydraulic
diameters, rod-to-wall spacings, and rod-to-rod spacings. The database was
compiled from data taken at two test l aboratories: Columbia University and the ATLAS facility. The uniform axial data was used to develop the correlation, while
the nonuniform axial data was used to validat e the correlation with the Tong factor.
Therefore, the correlation has been checked against independent test data.
The correlations address the effects of op erating pressure, mass velocity, enthalpy, axial power profile, local power peaking and distribution, rod diameter, and fuel
assembly hydraulic diameter and heated length on boiling transition.
The ANF-B and SPCB correlations have also been used to predict the number of
rods experiencing boiling transition (predi ct multiple indications) for the test database. The probability of boiling transition for each rod in a test section was
determined from the critical power prediction based on that rod. The probabilities
for all the rods in the test assembly, as predicted by ANF-B and SPCB, were then
summed to yield the prediction of the total number of rods experiencing boiling
transition. The ANF-B and SPCB correlations were found to conservatively
overpredict the expected number of rods that experience boiling transition (References 16, 17, 25 and 26).
4.4.2.3 Maximum Average Planar Linear Heat Generation Rate (MAPLHGR)
The MAPLHGR limit for fuel assures that the peak cladding temperature of fuel
following a postulated design basis loss-of-coolant accident (LOCA) will not exceed
the peak cladding temperature (PCT) and maximum oxidation limits specified in
10CFR50.46. The calculational procedure used to establish the MAPLHGR limits is
based on a LOCA analysis. The analysis is performed using calculational models LSCS-UFSAR 4.4-6 REV. 18, APRIL 2010 which are consistent with the requirements of Appendix K to 10CFR50. The models are described in Reference 20 for AREVA and Reference 12 for GE.
The PCT following a postulated LOCA is primarily a function of the average heat
generation rate of all the rods of a fuel assembly at any axial location and not strongly influenced by the rod-to-rod power distribution within the assembly.
The MAPLHGR limits for two-loop operation for a particular cycle are specified in
the COLR.
For single-loop operation, an APLHGR limit corresponding to the product of the
two-loop limit and a reduction factor specified in the COLR can be conservatively
used to ensure that the PCT for single-loop operation is bound by the PCT for two-
loop operation.
4.4.2.3.1 Design Power Distribution Thermal-hydraulic design of the reactor -- including the selection of the core size and effective heat transfer area, the design steam quality, the total recirculation
flow, the inlet subcooling, and the specification of internal flow distribution -- is
based on the concept and application of a design power distribution. The design
power distribution represents a conservative thermal operating state at rated
conditions and includes design allowances for the combined effects (on the fuel rod, and the fuel assembly heat flux and temper ature) of the gross and local steady-state power density distributions and adjustments of the control rods.
The design power distribution is used in conjunction with flow and pressure drop
distribution computations to determine the thermal conditions of the fuel and the
enthalpy conditions of the coolant throughout the core.
The design power distribution is based on detailed calculations of the neutron flux
distribution.
LSCS-UFSAR 4.4-7 REV. 18, APRIL 2010 The core average and maximum void fractions are dependent on the reactor operating state and power distributions. Typical average and maximum void
fraction results for AREVA fuel can be found in Reference 13.
4.4.2.4 Void Fraction Distribution
The core average and maximum void fractions for the initial core at rated condition
are given in Table 4.4-1. The typical axial distribution of core void fractions for the
average radial channel and the maximum radial channel (end of node value) is
given in Table 4.4-2. The core average and maximum exit value are also provided.
Similar distributions for steam quality are provided in Table 4.4-3. The core
average axial power distributions used to produce these tables are given in
Table 4.4-2a.
4.4.2.5 Core Coolant Flow Distribution
Correct distribution of core coolant flow among the fuel assemblies is accomplished
by the use of an accurately calibrated fixed or ifice at the inlet of each fuel assembly.
The orifices are located in the fuel support piece. They serve to control the flow
distribution and, hence, the coolant conditions within prescribed bounds throughout
the design range of core operation.
The core is divided into two orificed flow zones. The outer zone is a narrow, reduced-power region around the periphery of the core. The inner zone consists of
the core center region. No other control of flow and steam distribution other than
that incidentally supplied by adjusting the power distribution with the control rods, is used or needed. The orifices can be changed during refueling, if necessary.
The sizing and design of the orifices ensure stable flow in each fuel assembly during all phases of operation at normal operating conditions.
Design core flow distribution calculations are made using the design power
distribution which consists of a hot and average powered assembly in each of the two orifice zones. Typical design bundle powers and resulting relative flow
distributions are given in Table 4.4-4.
The flow distribution to the fuel assemblies is calculated on the assumption that the
pressure drop from lower plenum to upper plenum (across all fuel assemblies) is the same. This assumption has been confirmed by measuring the flow distribution in a modern boiling water reactor as reported in Reference 2.
LSCS-UFSAR 4.4-8 REV. 18, APRIL 2010 There is reasonable assurance, therefore, that the calculated flow distribution throughout the core is in close agreement with the actual flow distribution of an operating reactor. The use of the design power distribution discussed previously
ensures that the chosen orificing covers the range of normal operation. The expected
shifts in power production during core life are less severe and are bounded by the design
power distribution.
4.4.2.6 Core Pressure Drop and Hydraulic Loads
The pressure drop across various core compon ents under steady-state design conditions is included in Table 4.4-1 for the initial core. Initial Cycle analyses for the most limiting
conditions, the recirculation line break and the steamline break, are reported in Chapter
- 15. For SAFER/GESTR information, see Reference 12. For core pressure drop
information for AREVA fuel, see Reference 13.
The components of bundle pressure drop considered are friction, local, elevation, and
acceleration pressure drops. Pressure drop measurements made in operating reactors
confirm that the total measured core pressure drop and calculated core pressure drop
are in good agreement.
Subsections 4.4.2.6.1 through 4.4.2.6.4 describe the pressure drop models that were used
by GE for the initial core. AREVA utilizes similar pressure drop correlations and methodology. For more detail on these correlations and methodologies see References 14
and 15.
4.4.2.6.1 Friction Pressure Drop
Friction pressure drop is calculated using the relationship:
where: p f = friction pressure drop, psi, w = mass flow rate, g = acceleration of gravity, = water density, D H = channel hydraulic diameter, A ch = channel flow area, L = length, ()1-4.4 AD fL 2g w P2 TPF 2chH 2 f=
LSCS-UFSAR 4.4-9 REV. 14, APRIL 2002 f = friction factor, and 2 TPF = two phase friction multiplier.
This formulation is similar to that used throughout the nuclear power industry. The
formation for the two-phase multiplier is based on data which compares closely to that
found in the open literature (Reference 3).
4.4.2.6.2 Local Pressure Drop
The local pressure drop is defined as the i rreversible pressure loss associated with an area change such as the orifice, lower tie-plate, and spacers of a fuel assembly.
The general local pressure drop model is similar to the friction pressure drop and is
given by:
where:
P L = local pressure drop, psi; K = local pressure drop loss coefficient; A = reference area for local loss coefficient; and 2 TPL = two-phase local multiplier and w, g, and are defined the same as for friction. This basic calculation is similar to that used throughout the nuclear power industry. The formulation for the two-phase multiplier is similar to that reported in the open literature (Reference 4) with the addition of empirical constants to adjust the results to fit data taken at General
Electric Company for the specific designs of the BWR fuel assembly.
4.4.2.6.3 Elevation Pressure Drop
() 2-4.4 A K 2g w P 2 TPL 2 2 L= ()()3-4.4 -1 L; P g f E+==
LSCS-UFSAR 4.4-10 REV. 14, APRIL 2002 The elevation pressure drop is based on the well-known relationship where:
P E = elevation pressure drop, psi; L = incremental length:
= average coolant density:
= average void fraction over length -L; and f , g = saturated water and vapor density, respectively.
4.4.2.6.4 Acceleration Pressure Drop
A reversible pressure change occurs when an area change is encountered, and an
irreversible loss occurs when the fluid is accelerated through the boiling process.
The basic formulation for the reversible pressure change resulting from a flow area
change is given by:
where: P ACC = acceleration pressure drop, A 2 = final flow area, and A 1 = initial flow area
and other terms are as previously defined. The basic formulation for the
acceleration pressure change due to density change is:
where: ()()4-4.4
A A ;A 2 w 1P 1 2 2 2 2 2 ACC==g ()5-4.4
1 1A g w P IN M OUT M 2 ch 2 ACC= ()(), -1 x1g x1 f 2 2 M+=
LSCS-UFSAR 4.4-11 REV. 18, APRIL 2010 M = momentum density, and x = steam quality and other terms are as previously defined. The total acceleration pressure drop in
boiling water reactors is on the order of a few percent of the total pressure drop.
4.4.2.7 Correlation and Physical Data
The General Electric Company has obtained substantial amounts of physical data
in support of the pressure drop and thermal hydraulic loads discussed in Subsection
4.4.2.6. Correlations have been developed to fit this data to the formulations
discussed.
Subsection 4.4.2.7.1 through 4.4.2.7.3 de scribe the thermal hydraulic correlations
used by GE for the initial core. AREVA has also qualified their thermal hydraulic correlations for use in calculating pressure drop, void fraction, and heat transfer in
References 14 and 15.
4.4.2.7.1 Pressure Drop Correlations
The General Electric Company has taken significant amounts of friction pressure
drop data in multirod geometries representative of modern BWR plant fuel bundles
and correlated both the friction factor and two-phase multipliers on a best fit basis
using the pressure drop formulations reported in Subsections 4.4.2.6.1 and
4.4.2.6.2. Tests are performed in single-phase water to calibrate the orifice and the
lower tie-plate, and in both single- and two-phase flow to arrive at best fit design
values for spacer and upper tie-plate pressure drop.
The range of test variables is specified to include the range of interest to boiling
water reactors. New data are taken whenever there is a significant design change
to ensure the most applicable methods are in use at all times.
Applicability of the single-phase and two-phase hydraulic models discussed in
Subsections 4.4.2.6.1 and 4.4.2.6.2 is conf irmed by prototype (64-rod bundle) flow tests. The typical range of the test data is summarized in Table 4.4-5.
4.4.2.7.2 Void Fraction Correlation
The void fraction correlation is similar to models used throughout the nuclear
power industry and includes effects of pressure, flow direction, mass velocity, quality, and subcooled boiling.
LSCS-UFSAR 4.4-12 REV. 18, APRIL 2010 4.4.2.7.3 Heat Transfer Correlation The Jens-Lottes (Reference 5) heat transfer correlation is used in fuel design to
determine the cladding-to-coolant heat transfer coefficient for nucleate boiling.
4.4.2.8 Thermal Effects of Operational Transients
The evaluation of the core's capability to withstand the thermal effects resulting
from anticipated operational transients is covered in Chapter 15 and Appendix G.
In summary, all transients due to normal operation and to single operator error or
equipment malfunction result in MCPR greater than the transient MCPR limit.
4.4.2.9 Uncertainties in Estimates
Uncertainties in thermal-hydraulic parameters are considered in the statistical
analysis which is the basis for setting the transient MCPR limit such that at least
99.9% of the fuel rods in the core are expected not to experience boiling transition
during any abnormal operating transient. The statistical model and analytical
procedure are described in detail in References 1 and 11. The uncertainties
considered and their input values for the analysis are given in References 1 and 11.
For AREVA fuel, the statistical models and the methodology for calculating the MCPR safety limit are described in References 16, 17, 21, 25 and 26.
4.4.2.9.1 Transition Boiling Uncertainties
The fuel cladding employed for the nuclear fuel is Zircaloy. This material is
selected primarily for its nuclear properties. Zircaloy also has good corrosion and
strength properties at normal operating conditions. However, continued operation
at the elevated temperatures possible in the transition and film boiling regimes
could cause gradual reduction in strength and accelerated corrosion, resulting in
damage to the cladding.
The boiling transition does not necessarily correspond to the fuel damage threshold, especially in the high steam-quality range. Boiling transition is identified as the
heat transfer rate below which cladding overheating does not occur. Damage would
not actually occur until well into the film boiling regime. For example, during inpile
tests (Reference 6), Zircaloy-clad uranium dioxide fuel was purposely operated at
heat transfer rates well into film boiling for a total time exceeding 5 minutes, then
operated at typical boiling water reactor conditions for 10 days. Post-irradiation
examination showed evidence of overheating but no cladding failure. To ensure
good performance and long life of the cladding, conservative limits have been
established to ensure that normal operations remain well below the transition
boiling regime.
LSCS-UFSAR 4.4-13 REV. 14, APRIL 2002 4.4.2.9.2 Variation of Fuel Damage Limit Incipient center melting of the uranium dioxide pellet occurs at a higher kW/ft than
the peak LHGR during any abnormal operating transient. If UO 2 center melting occurs and the molten uranium dioxide is redistributed and densified, the damage
limit for strain can reduce to a lower value. The redistribution and densification
phenomena are functions of time and temperature. Plant transients of short
duration in the molten range do not result in appreciable redistribution or
densification. For the plant events that meet the transient MCPR limit, there is no
appreciable change in the kW/ft damage limit.
4.4.2.9.3 Effects of Misoriented Fuel Bundle
The concern with a misoriented assembly is primarily that the redistribution of
power among the fuel pins could lead to higher local powers than indicated by the
core monitoring system. In addition, a miso rientation could lead to slightly higher assembly powers as well. A detailed descri ption of this evaluation may be found in section 15.4.7.
4.4.2.10 Flux Tilt Considerations
For flux tilt considerations, refer to Subsection 4.3.2.2.7.
4.4.3 Description
of the Thermal and Hydraulic Design of the Reactor Coolant System
The thermal and hydraulic design of the reactor coolant system is described in this
subsection.
4.4.3.1 Plant Configuration Data
The descriptive summary of the reactor coolant system is given in Section 5.1. That
overview describes the reactor coolant pressure boundary and the reactor coolant
equipment used for the various coolant requirements encountered in both normal
and abnormal operations. The engineered safety functions are described in
Chapter 6.0 with system details and analysis shown there. The reactor
recirculation loops are described in detail in Subsection G.2.3 of Appendix G; The
main steam and feedwater systems are treated in Section 5.4. Plant configuration
data are included in these chapters.
Table 4.4-7 provides the flow path length, height, liquid level, minimum elevations, and minimum flow areas for each major flow path volume within the reactor vessel
and recirculation loops of the reactor coolant system. Table 4.4-8 provides the
lengths and sizes of all safety injectio n lines to the reactor coolant system.
LSCS-UFSAR 4.4-14 REV. 14, APRIL 2002 4.4.3.2 Operating Restrictions on Pumps See Subsection G.2.2 of Appendix G.
4.4.3.3 Power-Flow Operating Map
See Subsection G.2.3 of Appendix G.
4.4.3.4 Temperature-Power Operating Map (PWR)
Not applicable.
4.4.3.5 Load-Following Characteristics
See Subsection G.2.4 of Appendix G.
4.4.3.6 Thermal and Hydraulic Characteristics Summary Table
A summary of the thermal and hydraulic char acteristics of the reactor coolant system for the initial core and the initial cores of other reactors of similar design is included in
Table 4.4-1.
4.4.4 Evaluation
The thermal-hydraulic design of the reactor core and reactor coolant system is based
upon an objective of no fuel damage during normal operation or during abnormal
operational transients. This design objective is demonstrated by analysis in the
following sections.
4.4.4.1 Critical Heat Flux
Table 4.4-1 provides data on maximum heat flux, average heat flux, heat transfer areas, and other parameters affecting heat transfer of the initial core. The concept of
critical heat flux has been used in the determination of operationally significant power
distribution constraints. These are given in terms of the linear heat generation rate and minimum critical power ratio as discussed in the following subsections.
4.4.4.2 Core Hydraulics
See Subsection G.2.3 of Appendix G.
4.4.4.3 Influence of Power Distribution
The design constraints imposed by the maximum average planar linear heat
generation rate, the core power density, and the local peaking factor limit the gross LSCS-UFSAR 4.4-15 REV. 18, APRIL 2010 peaking factor (radial x axial). There are many combinations of radial and axial peaking factors that satisfy this design constraint, but each will have a different effect
on the MCPR. In general, the MCPR decreas es as the radial peaking (bundle power) increases and as the axial peak location moves to the top of the core. For example, for
a 1.96 gross factor, a flat (1.0) axial and a 1.96 radial would give a relatively low CPR, whereas a 1.0 radial and a 1.96 axial peaked in the bottom of the core would give a
relatively high CPR. These extremes are obviously not suited to design because they
are not representative of realistic reactor behavior. Therefore, the design radial
peaking factor is selected higher than that likely to be encountered in reactor
operation, and the combination of this radial with the design axial profile is also more
limiting than that expected during operating conditions.
4.4.4.4 Core Thermal Response
The thermal response of the core evaluated for expected transient conditions is
covered in Chapter 15. All expected abnormal operational transients are
conservatively evaluated to ensure that the integrity of the vessel and fuel is not
compromised. These transients are analyzed at varying power and flow conditions
within the analyzed power-to-flow map.
4.4.4.5 Analytical Methods
The analytical methods, thermodynamic data, and hydrodynamic data used in
determining the thermal and hydraulic characteristics of the core are similar to those
used throughout the nuclear power industry.
Core thermal-hydraulic analyses are performed with the aid of a digital computer
program. This program models the reacto r core through a hydraulic description of orifices, lower tie-plates, fuel rods, fuel rod spacers, upper tie-plates, fuel channel, and
the core bypass flow paths.
The methods discussed in section 4.4.4.5.1 through 4.4.4.5.3 describe the analytical
methods for GE. However, the descriptions below are typical for the nuclear industry.
These descriptions apply generally to AREVA methods. Further detailed descriptions of AREVA methods can be found in References 14, 17, 25 and 26.
4.4.4.5.1 Reactor Model
The orifice, lower tie-plate, fuel rod spacers, and upper tie-plate are hydraulically
represented as being separate, distinct local losses of zero thickness. The fuel channel cross section is represented by a square section with enclosed area equal to the
unrodded cross-sectional area of the actual fu el channel. The fuel channel assembly consists of three basic axial regions. The first and most important is the active fuel
region which consists of the fuel rods, nonfueled rods, and fuel-rod spacers. The
second is the nonfueled region consisting of nonfueled rods and the upper tie-plate.
LSCS-UFSAR 4.4-16 REV. 14, APRIL 2002 The third region represents the unrodded portion of the fuelchannel above the upper tie-plate. The active fuel region is considered in independent axial segments
or nodes over which fuel thermal properties are assumed constant and coolant
properties are assumed to vary linearly. The code can handle 12 fuel channel types and 10 types of bypass flow paths. In normal analyses the fuel assemblies are
modeled by four channel types--a hot centra l orifice region channel type, an average central orifice region channel type, a hot peripheral orifice region type and an
average peripheral orifice region type.
Usually there is one fuel assembly representing each of the hot types. The average types then make up the balance of the core.
The computer program iterates on flow through each flow path (fuel assemblies and
bypass paths) until the total differential pr essure (plenum to plenum) across each path is equal, and the sum of the flows thro ugh each path equals the total core flow.
Orificing is selected to optimize the core flow distribution between orifice regions as
discussed in Subsection 4.4.2.5. The core design pressure is determined from the
required turbine throttle pressure, the steamline pressure drop, steam dryer
pressure drop, and the steam separator pre ssure drop. The core inlet enthalpy is determined from the reactor and turbine heat balances. The core power
distribution is determined as per Subsection 4.4.2.3. The required core flow is then
determined by applying the procedures of this section and specifications such that
the thermal limits of Reference 11 are satisfied and the nominal expected bypass
flow fraction is approximately 10%. The results of applying these methods and
specifications are:
- a. flow for each bundle type,
- b. flow for each bypass path,
- c. core pressure drop,
- d. fluid property axial distribution for each bundle type, and
- e. CPR calculations for each bundle type.
4.4.4.5.2 System Flow Balances
The basic assumption used by the code in performing the hydraulic analysis is that the flow entering the core will divide itself between the fuel bundles and the bypass flow paths such that each assembly and bypass flow path experience the same
pressure drop.
LSCS-UFSAR 4.4-17 REV. 14, APRIL 2002 The bypass flow paths considered are described in Table 4.4-9 and shown in Figure 4.2-2. Due to the large flow area , the pressure drop in the bypass region above the core plate is essentially all elevation head. Thus, the sum of the core
plate differential pressure and the bypass region elevation head is equal to the core
differential pressure.
The total core flow less the control rod cooling flow enters the lower plenum through
the jet pumps. A fraction of this passe s through the various bypass paths. The remainder passes through the orifice in the fuel support (experiencing a pressure
loss) where more flow is lost through the fit-up between the fuel support and the lower tie-plate into the bypass region. The majority of the flow continues through
the lower tie-plate (experiencing a pressure loss) where some flow is lost through
the flow path defined by the fuel channel and lower tie-plate, and restricted by the finger springs, into the bypass region.
The flow through the bypass flow paths are expressed by the form:
Full scale tests have been performed to establish the flow coefficients for the major
flow paths. These tests simulate actual plant configurations which have several
parallel flow paths and, therefore, the fl ow coefficients for the individual paths could not be separated. However, analytical models of the individual flow paths
were developed as an independent check of the tests. The models were derived for
actual BWR design dimensions and considered the effects of dimensional variations.
These models predicted the test results when the "as-built" dimensions were
applied. When using these models for hydraulic design calculations, nominal
drawing dimensions are used. This is done to yield the most accurate prediction of
the expected bypass flow. With the larg e number of components in a typical BWR core, deviations from the nominal dimensions will tend to statistically cancel, resulting in a total bypass flow best represented by that calculated using nominal
dimensions.
The balance of the flow enters the fuel bundle from the lower tie plate and passes
through the fuel rod channel spaces. A small portion of the in-channel flow enters
the non-fueled rod through orifice holes just above the lower tie-plate. This flow, normally referred to as the water-rod flow, remixes with the active coolant channel
flow below the upper tie-plate.
4.4.4.5.3 System Heat Balances
Within the fuel assembly, heat balances on the active coolant are performed
nodally. Fluid properties are expressed as the bundle average at the particular
node of interest and are based on Reference 7. In evaluating fluid properties a
constant pressure model is used.
()6-4.4 P C PCP CW 2 3 4 C 221 1++=
LSCS-UFSAR 4.4-18 REV. 14, APRIL 2002 The core power is divided into two parts: an active coolant power and a bypass flow power. The bypass flow is heated by neutron-slowing down and gamma heating in the
water and by heat transfer through the channe l walls. Heat is also transferred to the bypass flow from structures and control elements which are themselves heated by gamma absorption and by (n, ) reactions in the control material. The fraction of total reactor power deposited in the bypass region is very nearly 2%. A similar phenomena occurs with the fuel bundle to the active coolant and the water rod flows. The net
effect is that approximately 96% of the core power is conducted through the fuel
cladding and appears as heat flux.
The power is allocated to the individual fuel bundles using a relative power factor.
The power distribution along the length of th e fuel bundle is specified with axial power factors which distribute the bundle's power among the axial nodes. A nodal location
power or peaking factor is used to establish the peak heat flux at each nodal location.
Relative, axial, and local peaking factors are more thoroughly discussed in Subsection
4.3.2.
The relative (radial) and axial power distributions when used with the bundle flow
determine the axial coolant property distribution resulting in sufficient information to
calculate the pressure drop components with in each fuel assembly type. Once the equal pressure drop criterion has been sati sfied, the critical bundle power (the power which would result in critical quality existing at some point in the bundle using the
correlation expressed in References 1 and 11) is determined by an iterative process for
each fuel type.
In applying the above methods to core design, the number of bundles (for a specified
core thermal power) and bundle geometry (8 x 8, rod diameter, etc.) are selected based
on power density and linear heat generation rate limits.
4.4.4.5.4 Uncertainties in Design Analyses
The effects of uncertainties in design values and on calculational results are accounted
for in the statistical analysis on which the MCPR limits are based.
4.4.4.6 Reactor Stability Analysis
4.4.4.6.1 Introduction
There are many definitions of stability, but for feedback processes and control systems
it can be defined as follows: a system is stable if, following a disturbance, the
transient settles to a steady, noncyclic state.
A system may also be acceptably safe even if oscillatory, provided that any limit cycle of the oscillations is less than a prescribed magnitude. Instability then, is either a
continual departure from a final steady-state value or greater-than-prescribed limit
cycle about the final steady-state value.
LSCS-UFSAR 4.4-19 REV. 14, APRIL 2002 The mechanism for instability can be explained in terms of frequency response.
Consider a sinusoidal input to a feedback control system which for the moment has
the feedback disconnected. If there were no time lags or delays between input and
output, the output would be in phase with the input. Connecting the output so as to
subtract from the input (negative feedback or 180º out-of-phase connection) would
result in stable closed loop operation. However, natural laws can cause phase shift
between output and input and should the phase shift reach 180º, the feedback
signal would be reinforcing the input signal rather than subtracting from it. If the
feedback signal were equal to or larger than the input signal (loop gain equal to one
or greater), the input signal could be disconnected and the system would continue to
oscillate. If the feedback signal were less than the input signal (loop gains less than
one), the oscillations would die out.
The design of the BWR is based on the premise that power oscillations can be
readily detected and suppressed.
4.4.4.6.2 Description
Three types of stability considered in the design of boiling water reactors are (1)
reactor core (reactivity) stability, (2) channel hydrodynamic stability, and (3) total
system stability. Reactivity feedback instability of the reactor core could drive the
reactor into power oscillations. Hydrodynamic channel instability could impede
heat transfer to the moderator and drive the reactor into power oscillations. The
total system stability considers control system dynamics combined with basic
process dynamics. The criteria is demonstrated if it is analytically demonstrated
that no divergent oscillation develops within the system as a result of calculated
step disturbances of any critical variable, such as steam flow, pressure, neutron
flux, and recirculation flow, or that the divergent oscillation can be detected and suppressed.
Stability is expressed in terms of two compatible parameters. First is the decay
ratio x 2/x 0 , designated as the ratio of the magnitude of the second overshoot to the first overshoot resulting from a step perturbation. A plot of the decay ratio is a
graphic representation of the physical responsiveness of the system, which is readily evaluated in a time-domain analysis. Second is the damping coefficient n, the definition of which corresponds to the pole pair closest to the j axis in the s-plane for the system closed loop transfer function. This parameter also applies to the frequency-domain interpretation. The damping coefficient is related to the decay ratio as shown in Figure 4.4-1.
4.4.4.6.3 Solution Description for Thermal-Hydraulic Stability
BWR cores may exhibit thermal-hydraulic instabilities in certain portions of the
core power and recirculation flow operating domain. The instabilities and the
solutions devised to detect and suppress them are discussed in Reference 22 and 23.
LSCS-UFSAR 4.4-20 REV. 17, APRIL 2008 LSCS has adopted the solution Option III, designated as the Oscillation Power Range Monitor (OPRM). The OPRM complies with GDC-12, as discussed in
Section 3.1.2.2.3.
The overall design philosophy of the OPRM is to generate an alarm in the control
room if it detects core instabilities (based on period-based algorithm only), and to generate an automatic suppression system trip if the instabilities reach an
amplitude that could threaten the fuel safety limits.
The overall objective of the oscillation detection algorithm is to reliably detect expected instabilities at a low magnitude such that mitigation can occur well before the MCPR Safety Limit is exceeded, while avoiding spurious trips during expected neutron flux transients. The algorithm is based on the detection of the three known characteristics that BWR neutron flux oscillations exhibit. These characteristics are the amplitude or absolute magnitude, growth rate, and periodic behavior. Only the period based detection algorithm is used in the safety analysis. The other algorithms provide defense in depth and additional protection against unanticipated oscillations. Details of the algorithm can be found in References 22 and 23.
The OPRM consists of a micoprocessor that analyzes signals from LPRMs. Since LPRMs are evenly distributed throughout the reactor core, they are capable of
responding to any neutron flux oscillations that can create an MCPR concern.
Individual LPRMs readily respond to a wide variety of normal operating maneuvers
and expected events, and are also subject to electrical interference. For these
reasons, each OPRM may use multiple LPRMs as a means of maintaining a strong
response to a neutron flux oscillation while minimizing the susceptibility to false
signals associated with a single LPRM, or may utilize a detection algorithm designed to achieve the same objective. The OPRM is automatically bypassed at
high flow or low power conditions, where core instabilities are unlikely to occur, to
avoid spurious actuation.
4.4.4.6.4 Stability Criteria
The following discussion on stability is based on the original design bases, which did not assume an inherent tendency towards oscillations. They are presented here
for historic perspective. The new design, in compliance with the NRC Generic
Letter 94-02, is based on the detection and suppression methodology, and is
discussed above in Section 4.4.4.6.3.
Stability criteria are established to demonstrate compliance with the requirements
set forth in 10CFR50 Appendix A, General Design Criterion (GDC) 12.
These stability compliance criteria consider potential limit cycle response within the
limits of safety system and/or operator in tervention and the OPRM assures that for BWR fuel designs this operating mode does not result in specified acceptable LSCS-UFSAR 4.4-21 REV. 17, APRIL 2008 fuel design limits being exceeded. The onset of power oscillations for which corrective actions are necessary is reliably and readily detected and suppressed by
operator actions and/or automatic system functions.
To ensure compliance of the GE BWR design with GDC 12 requirements, the following stability acceptance criteria have been established.
(1) Neutron flux limit cycles which oscillate up to the 120% APRM high neutron flux scram set point or up to the LPRM upscale
alarm trip (without initiating scram) prior to operator
mitigating action, shall not result in exceeding specified
acceptable fuel design limits.
(2) The individual channels shall be designed and operated to be hydrodynamically stable or more stable than the reactor core for
all expected operating conditions.
Calculations which predict that core-wide limit cycles will not occur (decay ratio
< 0.8) also demonstrate compliance with GDC-12.
This criteria is presently used for LSCS two recirculation loop operation. For single
recirculation loop operation, the plant is monitored per General Electric SIL-380 (Reference 8).
These criteria shall be satisfied for all attainable conditions of the reactor that may
be encountered in the course of plant operation. For stability purposes, the most
severe conditions to which these criteria will be applied correspond to natural
circulation flow at a power corresponding to the extrapolated APRM rod block
intercept condition.
The licensing basis is to generate a trip signal during oscillations of sufficiently low
amplitude to provide margin to the MCPR safety limits for all expected modes of BWR oscillations. The OPRM oscillation recognition algorithm is intended to
discriminate between true stability-related neutron flux oscillations and other flux
variations that may be expected during plant operation. Extensive evaluation of
operating plant data is done to determine the combination of algorithm and OPRM
setpoints, which meet the design objectives. The final algorithm/setpoint design is
subjected to in-plant testing with the trip function disabled.
The OPRM assures that for BWR fuel designs, this operating mode does not result
in specified acceptable fuel design limits being exceeded. The onset of power
oscillations for which corrective actions are necessary is reliably and readily
detected and suppressed by operator actions and/or automatic system functions.
LSCS-UFSAR 4.4-22 REV. 17, APRIL 2008 4.4.4.6.5 Expected Oscillation Modes
The OPRM is capable of responding to the expected modes of BWR stability-related
oscillations. The expected oscillation modes are as follows (Reference 13, Section 6.1):
- Core-wide, in which the average neutron flux in all fuel assemblies oscillates
in phase.
- First Order Side-by-Side or a regional oscillation where the neutron flux on
one side of the reactor oscillates 180 o out of phase with the flux on the other side.
- First Order Precession a regional oscillation where the axis of zero oscillation amplitude rotates azimuthally, or the two reactor regions of peak oscillation
amplitude shift from one location to another at a frequency lower than the
oscillation frequency.
Other modes of oscillation are not expected in a BWR.
4.4.4.6.6 Analysis Approach
The total system stability analysis evaluates the relative stability of the total
system, from time responses generated by applying step changes to the input variables to the total system stability model. The observed time response of an
output variable of a high order dynamic system represents a superposition of the
system's several response modes. The relative intensity of each particular mode in
the time response is determined by the zeroes (the roots of the numerator) of the
transfer function relating a given output variable to a particular input. Therefore, in judging the relative stability of the sy stem, the observer should separate the distinct modes in the time response and apply the stability criterion to each modal
response. The approach used here, of disturbing one input variable and applying
the stability criterion to the resulting system response is a good approximation to
modal separation. It is particularly applicable in calculating stability since, as a
system tends toward instability a single oscillatory mode tends to dominate the
observed time response (Reference 9).
LSCS-UFSAR 4.4-23 REV. 17, APRIL 2008 Reference 15 describes the process used to calculate a conservative final MCPR value for an anticipated stability-related oscillation. It involves the determination of
initial MCPR by a cycle-specific evaluation and the calculation of hot bundle
oscillation magnitude. The licensing crite rion is met when the final MCPR is greater than the MCPR safety limit. Appr opriate reload parameters are checked every cycle to determine the initial MCPR. This methodology provides a
conservative means of demonstrating with a high probability and confidence that
the MCPR safety limits will not be violated for anticipated oscillations. The use of the MCPR safety limit to provide protection against possible fuel damage is
exceedingly conservative (Reference 24, Section 4.5.2).
4.4.4.6.7 Mathematical Model
This mathematical model applies to the initial core analysis. The mathematical
model representing the core examines the linearized reactivity response of a reactor
system with density-dependent reactivity feedback caused by boiling. In addition, the hydrodynamics of various hydraulically coupled reactor channels or regions are
examined separately on an axially multin oded basis by grouping various channels that are thermodynamically and hydraulically similar. This interchannel
hydrodynamic interaction or coupling exists through pressure variations in the inlet plenum, such as can be caused by disturbances in the flow distribution between
regions or channels. This approach provides a reasonably accurate, three-dimensional representation of the reactor's hydrodynamics.
The core model, shown in block diagram fo rm in Figure 4.4-2, solves the dynamic equations that represent the reactor core in the frequency domain. From the
solution of these dynamic equations, the reactivity and individual channel
hydrodynamic stability of the boiling water reactor is determined for a given reactor
flow rate, power distribution, and total power. This gives the most basic
understanding of the inherent core behavior (and hence the system behavior) and is
the principal consideration in evaluating the stable performance of the reactor. As LSCS-UFSAR 4.4-24 REV. 17, APRIL 2008 new experimental or reactor operating data are obtained, the model is refined to improve its capability and accuracy.
The plant model considers the entire reactor system, neutronics, heat transfer, hydraulics, and the basic processes, as well as associated control systems such as
the flow controller, pressure regulator, feedwater controller, etc. Although the
control systems may be stable when analyzed individually, final control system
settings must be made in conjunction with the operating reactor so that the entire
system is stable. The plant model yields results that are essentially equivalent to
those achieved with the core model and allows the addition of the controllers, which
have adjustable features permitting the attainment of the desired performance.
The plant model solves the dynamic equations that present the BWR system in the
time domain. The variables, such as steam flow and pressure, are represented as a
function of time. The extensiveness of this model (Reference 10) is shown in block diagram form in Figure 4.4-3. Many of the blocks are extensive systems in
themselves. The model is periodically refined as new experimental or reactor
operating data are obtained to improve its capability and accuracy.
4.4.4.6.8 Initial Core Analysis Results
The results of the two recirculation pump operation core and channel stability
analysis is given in the Reload Licensing Package for each cycle. The plant stability
analysis is performed only for the initial core and is described below.
The plant stability analysis was performed by assuming that the reactor is initially
operating at the most sensitive condition, corresponding to natural circulation flow
and a power level at the rod block limit. The nuclear system is then subjected to
step disturbances from control rods, pressure regulator setpoint, and level controller
setpoint. These time responses are shown in Figures 4.4-6 through 4.4-8. It is clear
that the decay ratio is less than the stability criterion.
For expected normal operating modes, the time response of each of the important
variables of the reactor system (neutron flux, pressure, and steam flow) to small step disturbances can be underdamped, but must analytically show a decay ratio of
less than 0.25 in order to satisfy the operational design guide limit. Using final
design parameters each of the following disturbances are analytically imposed, one
at a time, using the model previously described for time domain analysis:
- a. a pressure setpoint change of at least 5 psi,
- b. a control rod position change equivalent to a local power change of at least 5% of point (of the magnitude of power at the time of the
disturbance),
LSCS-UFSAR 4.4-25 REV. 17, APRIL 2008 c. a load demand change of at least 5% of point, and d. a reactor water level setpoint change of at least 6 inches.
Using actual design parameters, calculated responses of important nuclear system variables to step disturbances from control rod reactivity, pressure regulator
setpoint, level controller setpoint, and turbine load setpoint are tested for rated
power-flow conditions and at the nominal power corresponding to the lower end of
the automatic power-flow control path.
Results of the analysis for 105% rated power and 100% rated flow are shown in
Figures 4.4-9 through 4.4-12. It is evident that the response meets the stability
criterion. Figures 4.4-13 through 4.4-16 show the results of analysis at the low
limit of the automatic flow-control range.
It is concluded that for all normal operating points over the flow-control range the
decay ratio of the total system response s is less than one-fourth, good dynamic performance is expected, and the ratio conforms with the stability criterion.
4.4.5 Testing
and Verification
See Subsection G.4.3 of Appendix G.
The OPRM, which is installed to detect and suppress thermal-hydraulic Instabilities, is extensively tested using available data from several BWR plants.
After installation, the plant is operated for a period of time with the OPRM trip function disabled while OPRM performance is monitored for susceptibility to spurious trips. The OPRM trip function is enabled following approval of the associated Technical Specification.
4.4.6 Instrumentation
Requirements See Subsections 7.7.3.2 and 7.6.3.4 of Chapter 7.
4.4.6.1 Loose Parts Monitoring System (Deleted)
LSCS-UFSAR 4.4-26 REV. 17, APRIL 2008 This page intentionally left blank.
LSCS-UFSAR 4.4-27 REV. 17, APRIL 2008
4.4.7 References
- 1. "General Electric BWR Thermal Analysis Basis (GETAB): Data, Correlation, and Design Application," January 1977 (NEDE-10958-PA
and NEDO-10958A).
- 2. "Core Flow Distribution in a Modern Boiling Water Reactor as Measured in Monticello," NEDO-10299, AEC Topical Report
NEDO-10299, January 1971.
- 3. R. C. Martinelli and D. E. Nelson, "Prediction of Pressure Drops During Forced Connection Boiling of Water," ASME Trans., 70, pp. 695-702, 1948.
- 4. C. J. Baroozy, "A Systematic Correlation for Two-Phase Pressure Drop," Heat Transfer Conference, Preprint No. 37, AICLE, Los
Angeles, 1966.
- 5. W. H. Jens and P. A. Lottes, "Analysis of Heat Transfer, Burnout, Pressure Drop, and Density Data for High Pressure Water," USAEC
Report - 4627, 1972.
LSCS-UFSAR 4.4-28 REV. 18, APRIL 2010
- 6. S. Levy et al., "Experience with BWR Fuel Rods Operating Above Critical Flux," Nucleonics , April 1965.
- 7. 1967 International Standard Steam Water Properties.
- 8. General Electric Service Information Letter (SIL) No. 380, "BWR Core Thermal Hydraulic Stability," Rev. 1, dated February 10, 1984.
- 9. Zadeh and Desoer, "Linear System Theory," McGraw-Hill Book Co., 1963.
- 10. "Analytical Methods of Plant Transient Evaluations for General Electric Boiling Water Reactor," NEDO-10802, General Electric
Company, BWR Systems Department, February 1973.
- 11. "General Electric Standard Application for Reactor Fuel," NEDE-P-A, (Latest approved revision).
- 12. GE Document, "SAFER/GESTR-LOCA, Loss-of-Coolant Accident Analysis, LaSalle County Station Units 1 & 2," NEDC-31510P, as
amended & revised.
- 13. Cycle specific Fuel Design Report (AREVA fuel only).
- 14. Exxon Nuclear Methodology for Boiling Water Reactors: THERMEX Thermal Limits Methodology, Summary Description , XN-NF-80-19(A), Volume 3, Revision 2, Exxon Nuclear Company, Inc., Richland, WA (January 1987).
- 15. Generic Mechanical Design Criteria for BWR Fuel Design , ANF-89-98(P)(A), Revision 1 and Supplement 1, Advanced Nuclear
Fuels Corporation, Richland, WA (May 1995).
- 16. Advanced Nuclear Fuels Critical Power Methodology for Boiling Water Reactors, ANF-524(P)(A), Revision 2 and Supplements, Advanced
Nuclear Fuels Corporation, November 1990.
- 17. ANFB Critical Power Correlation , ANF-1125 (P)(A) and Supplements 1 and 2, Advanced Nuclear Fuels Corporation, April 1990; ANFB Critical Power Correlation Application for Co-Resident Fuel , EMF-1125(P)(A), Supplement 1, Appendix C, Siemens Power Corporation, August 1997;
and ANFB Critical Power Correlation Determination of ATRIUM-9B Additive Constant Uncertainties , ANF-1125(P)(A) Supplement 1, Appendix E, Siemens Power Corporation, September 1998.
LSCS-UFSAR 4.4-29 REV. 18, APRIL 2010
- 18. Siemens Power Corporation Methodology for Boiling Water Reactors:
Evaluation and Validation of CASMO-4 / MICROBURN-B2 EMF-2158(P)(A) Revision 0, Siemens Power Corporation, October 1999.
- 19. Exxon Nuclear Methodology for Boiling Water Reactors - Neutronic Methods for Design and Analysis , XN-NF-80-19(P)(A) Volume 1 and Supplements 1 and 2, Exxon Nuclear Company, March 1983.
- 20. EXEM BWR-2000 ECCS Evaluation Model EMF-2361(P)(A)
Revision 0, AREVA NP Inc., May 2001.
- 21. MICROBURN-B2 Based Impact of Failed / Bypassed LPRMs and TIPs, Extended LPRM Calibration Interval, and Single Loop Operation Measured Radical Bundle Power Uncertainty , EMF-2493(P)
Revision 0, Siemens Power Corporation, December 2000.
- 22. NEDO-31960, "BWR Owners Group Long-Term Stability Solutions Licensing Methodology," June 1991.
- 23. NEDO 31960, "BWR Owners Group Long-Term Stability Solutions Licensing Methodology," Supplement 1, March 1992.
- 24. NEDO-32465-A, "BWR Owners' Group Reactor Stability Detect and Suppress Solution Licensing Basis Methodology and Reload
Application," August 1996.
- 25. SPCB Critical Power Correlation , EMF-2209(P)(A) Revision 3, AREVA NP, September 2009.
- 26. Application of Siemens Power Corporation's Critical Power Correlation to Co-Resident Fuel , EMF-2245(P)(A), Revision 0, Siemens Power Corporation, August 2000.
LSCS-UFSAR TABLE 4.4-1 (SHEET 1 OF 2) TABLE 4.4-1 REV. 13 THERMAL AND HYDRAULIC DESIGN CHARACTERISTICS OF THE REACTOR CORE (INITIAL CORE DATA) 238-732 218-592 218-560 251-764 251-784 251-764 BWR/6 BWR/6 ZPS-1 WPPSS NP No. 2 BWR/6 LSCS GENERAL OPERATING CONDITIONS Reference design thermal output, MWt 3579 2894 2436 3323 3833 3323 Power level for engineered safety features, MWt 3758 3039 2550 3489 4025 3489 Steam flow rate, at 420° F final feedwater temperature, millions lb/hr 15.396 12.451 10.477 14.295 16.488 14.166 Core coolant flow rate, millions
lb/hr 105.0 84.5 78.5 108.5 113.5 108.5 Feedwater flow rate, millions
lb/hr 15.358 12.42 10.448 14.256 16.488 14.127 System pressure, nominal in steam dome, psia 1040 1040 1020 1020 1040 1020 System pressure, nominal core design, psia 1055 1055 1035 1035 1055 1035 Coolant saturation temperature at core design pressure, °F 551.1 551.1 548.8 548.8 551.1 548.8 Average power density, kW/liter 56 56 50.51 51.2 56.0 48.17 Specific power, kW/kg (U total) 25.9 25.9 23.7 23.7 25.9 23.7 Maximum thermal output, kW/ft 13.4 13.4 13.4 13.4 13.4 13.4 Average thermal output, kW/ft 6.04 6.04 5.45 5.45 6.04 5.33 Core total heat transfer area, ft 2 73,409 59,369 55,401 75,582 78,624 74,871 Maximum heat flux, Btu/hr-ft 2 354,000 354,000 354,000 354,000 354,000 361,000 Average heat flux, Btu/hr-ft 2 159,550 159,550 143,900 143,920 159,550 143,740 LSCS-UFSAR TABLE 4.4-1 (SHEET 2 OF 2) TABLE 4.4-1 REV. 13 238-732 218-592 218-560 251-764 251-784 251-764 BWR/6 BWR/6 ZPS-1 WPPSS NP No. 2 BWR/6 LSCS GENERAL OPERATING CONDITIONS Core inlet enthalpy, at 420° F FFWT, Btu/lb 527.8 527.8 527.4 527.6 528.1 527.5 Core inlet temperature, at 420° F FFWT, °F 533.0 533.0 532.6 532.8 533.3 532.8 Core maximum exit voids within assemblies, % 76 76 75 75 76 76 Core average void fraction, active coolant 0.428 0.429 0.418 0.415 0.427 0.418 Active coolant flow area per
assembly, in 2 15.50 15.50 15.50 15.50 15.50 15.82 Core average inlet velocity, ft/sec 7.2 7.2 7.0 7.1 7.2 6.77 Maximum inlet velocity, ft/sec 7.6 7.6 7.4 7.5 7.6 7.2 Total core pressure drop, psi 25.7 25.5 27.3 27.5 25.8 24.8 Core support plate pressure drop, psi 21.3 21.1 22.9 23.1 21.4 19.61 Average orifice pressure drop Central region, psi 8.6 8.5 11.2 11.4 8.7 8.13 Peripheral region, psi 17.3 17.2 19.6 19.8 17.5 16.66 Maximum channel pressure loading, psi 14.5 14.5 13.7 13.7 14.6 12.84 TYPICAL POWER PEAKING FACTOR Maximum relative assembly power 1.40 1.40 1.40 1.40 1.40 1.40 Local peaking factor 1.13 1.13 1.24 1.15 1.13 1.15 Axial peaking factor 1.40 1.40 1.40 1.40 1.40 1.40 Gross peaking factor 1.96 1.96 1.96 1.96 1.96 1.96 Total peaking factor 2.22 2.22 2.43 2. 2.22 2.25 LSCS-UFSAR TABLE 4.4-2 TABLE 4.4-2 REV. 13 TYPICAL VOID DISTRIBUTION (INITIAL CORE)
NODE CORE AVERAGE (AVERAGE NODE VALUE) MAXIMUM CHANNEL (END OF NODE VALUE)
Bottom 1 0.000 0.0 2 0.001 0.032 3 0.018 0.122 4 0.065 0.230 5 0.136 0.325 6 0.212 0.401 7 0.281 0.462 8 0.341 0.511 9 0.391 0.552 10 0.433 0.587 11 0.469 0.616 12 0.499 0.641 13 0.525 0.662 14 0.547 0.681 15 0.566 0.696 16 0.582 0.708 17 0.595 0.719 18 0.606 0.728 19 0.616 0.736 20 0.624 0.742 21 0.631 0.748 22 0.637 0.753 23 0.643 0.757 Top 24 0.647 0.761
Core average value = 0.419 M aximum exit value = 0.761 Active fuel length = 150 inches LSCS-UFSAR TABLE 4.4-2a TABLE 4.4-2a REV. 4 - APRIL 1988 AXIAL POWER DISTRIBUTION USED TO GENERATE VOID AND QUALITY DISTRIBUTIONS (TYPICAL)
AXIAL NODE POWER-FACTOR Bottom of core 1 0.54 2 0.83 3 1.02 4 1.17 5 1.26 6 1.33 7 1.37 8 1.39 9 1.40 10 1.39 11 1.38 12 1.34 13 1.29 14 1.21 15 1.10 16 0.99 17 0.89 18 0.79 19 0.71 20 0.64 21 0.58 22 0.52 23 0.46 Top of core 24 0.40
LSCS-UFSAR TABLE 4.4-3 (SHEET 1 OF 2) TABLE 4.4-3 REV. 13 FLOW QUALITY DISTRIBUTION (TYPICAL)
- Core average value = 0.074 Maximum exit value = 0.281 Active fuel length = 150 inches NODE CORE AVERAGE (AVERAGE NODE VALUE) MAXIMUM CHANNEL (END OF NODE VALUE)
BOTTOM 1 0.00 0.00 2 0.000 0.001 3 0.000 0.006 4 0.002 0.017 5 0.006 0.032 6 0.013 0.049 7 0.022 0.067 8 0.032 0.085 9 0.044 0.103 10 0.053 0.121 11 0.063 0.139 12 0.073 0.157 13 0.083 0.173 14 0.093 0.189 15 0.101 0.203 16 0.109 0.216 17 0.117 0.228 18 0.123 0.238 19 0.129 0.248 20 0.134 0.256 21 0.138 0.263
LSCS-UFSAR TABLE 4.4-3 (SHEET 2 OF 2)
TABLE 4.4-3 REV. 16, APRIL 2006 NODE CORE AVERAGE (AVERAGE NODE VALUE) MAXIMUM CHANNEL (END OF NODE VALUE) 22 0.142 0.270 23 0.146 0.276 TOP 24 0.150 0.281
- These flow quality distribution valu es are typical for the initial core.
The GE9 and GE14 fuel has an active fuel length of 150 inches. The
ATRIUM-9B and ATRIUM-10 fuel have an active fuel length of 149.0 inches. This design characteristic difference in combination with changes in power distribution and reactor core state produce different flow quality distributions. These differences are included in transient and core design methodology.
LSCS-UFSAR TABLE 4.4-4 TABLE 4.4-4 REV. 4 - APRIL 1988 CORE FLOW DISTRIBUTION (TYPICAL)
OFFICE ZONEDESCRIPTION CENTRAL HOT CENTRAL AVERAGE PERIPHERAL HOT PERIPHERAL AVERAGE Relative Assembly Power 1.4 1.04 0.95 0.70 Relative Assembly Flow 0.93 1.06 0.55 0.57
LSCS-UFSAR TABLE 4.4-5 TABLE 4.4-5 REV. 0 - APRIL 1984 TYPICAL RANGE OF TEST DATA MEASURED PARAMETER TEST CONDITIONS ADIABATIC TESTS Spacer single-phase loss coefficient N Re* = 0.5 x 10 5 to 3.5 x 10 5 Lower tie plate + orifice
single-phase loss
coefficient T = 100 to 500°F Upper tie plate single-phase friction factor Spacer two-phase loss
coefficient P = 800 to 1400 psia Two-phase friction
multiplier G = 0.5 x 10 6 to 1.5 x 10 6 lb/h-ft 2 X - 0 to 40%
DIABATIC TESTS Heated bundle pressure drop P = 800 to 1400 psia G = 0.5 x 10 6 to 1.5 x 10 6 lb/h-ft 2 ___________________
- Reynolds Number
LSCS-UFSAR TABLE 4.4-6 TABLE 4.4-6 REV. 4 - APRIL 1988
THIS PAGE LEFT INTENTIONALLY BLANK.
LSCS-UFSAR TABLE 4.4-7 TABLE 4.4-7 REV. 0 - APRIL 1984 REACTOR COOLANT SYSTEM GEOMETRICAL DATA FLOW PATH LENGTH (in.) HEIGHT AND LIQUID LEVEL (in.) ELEVATION OF BOTTOM OF EACH VOLUME* (in.) MINIMU M FLOW AREAS (ft 2) A. Lower Plenum 216 216 216 -172.5 71.5 B. Core 164 164 164 44 142.0 C. Upper Plenum and Separators 178 178 208 49.5 D. Dome (Above Normal Water Level) 312 312 386.0 343.5 E. Downcomer Area 321 321 321 -51.0 79.5 F. Recirculation Loops and Jet Pumps (one loop) 108.5 ft (one loop) 403 -394.5 132.5 in 2
- Reference point is recirculat ion nozzle outlet centerline.
LSCS-UFSAR TABLE 4.4-8 TABLE 4.4-8 REV. 0 - APRIL 1984 LENGTHS AND SIZES OF SAFETY INJECTION LINES LINE OD (inches)
LINE LENGTH (feet
)I. HPCS Line A. Pump discharge to valve 16 146.0 B. Inside containment to RPV 12 101.5 Total 247.5 II. LPCI Lines A. Loop A
- 1. Pump discharge to valve* 18/12 182.0 2. Inside containment to RPV 12 101.5 Total 283.5 B. Loop B
- 1. Pump discharge to valve* 18/12 388.5
- 2. Inside containment to RPV 12 84.5 Total 473.0 C. Loop C
- 1. Pump discharge to valve* 18 344.0 2. Inside containment to RPV 12 77.0 Total 421 0 III. LPCS Line A. Pump discharge to valve* 16 282.5 B. Inside containment to RPV 12 84.5 Total 367.0
___________________
- Valve located as near as possible to outside of containment wall.
LSCS-UFSAR TABLE 4.4-9 TABLE 4.4-9 REV. 0 - APRIL 1984 BYPASS FLOW PATHS FLOW PATH DESCRIPTION DRIVING PRESSURE NUMBER OF PATHS 1a. Between Fuel Support and the Control Rod Guide Tube (Upper Path) Core Plate Differential One/Control Rod 1b. Between Fuel Support and the Control Rod Guide Tube (Lower Path) Core Plate Differential One/Control Rod 2. Between Core Plate and the Control Rod Guide Tube Core Plate Differential One/Control Rod
- 3. Between Core Support and the Incore Support Instrument Guide Tube Core Plate Differential One/Instrument 4. Between Core Plate and Shroud Core Plate Differential One
- 5. Between Control Rod Guide Tube and Control Rod Drive Housing Core Plate Differential One/Control Rod 6. Between Fuel Support and Lower Tie-Plate Channel Wall Differential Plus Lower Tie-Plate Differential One/Channel 7. Control Rod Drive Coolant Independent of of Core One/Control Rod
- 8. Between Fuel Channel and Lower Tie-Plate Channel Wall Differential One/Channel 9. Holes in Lower Tie-Plate Lower Tie-Plate/ Bypass Region Differential Two/Assembly
LSCS-UFSAR 4.5-1 REV. 13 4.5 REACTOR MATERIALS
4.5.1 Control
Rod System Structural Materials
4.5.1.1 Material Specifications The following material listing applies to the control rod drive mechanism supplied for this application. The position in dicator and minor nonstructural items are
omitted. a. Cylinder, Tube and Flange Assembly Flange ASME SA 182 Grade F304 Plugs ASME SA 182 Grade F304 Cylinder ASTM A269 Grade TP 304 Outer Tube ASTM A269 Grade TP 304 Tube ASTM A351 Grade CF-3 Spacer ASTM A351 Grade CF-3
- b. Piston Tube Assembly Piston Tube ASTM A479 Grade XM-19 Stud ASTM A276 Type 304 Head ASME SA 182 Grade F304 Ind. Tube ASME SA 312 Type 316 Cap ASME SA 182 Grade F304.
- c. Drive Assembly Coupling Spud Inconel X-750 Index Tube ASTM A479 Grade XM-19 Piston Head Armco17-4 PH Coupling ASME SA 312 Grade TP 304 or ASTM A511 Grade MT 304 Magnet Housing ASME SA 312 Grade TP 304 or ASTM A511 Grade MT 304.
- d. Collet Assembly Collet Piston ASTM A269 Grade TP 304 or ASME SA 312 Grade TP 304 Finger Inconel X-750 Retainer ASTM A260 Grade TP 304 or ASTM A511 Grade MT 304 Guide Cap ASTM A269 Grade TP 304.
LSCS-UFSAR 4.5-2 REV. 13
- e. Miscellaneous Parts Stop Piston ASTM A276 Type 304 Connector ASTM A276 Type 304 O-Ring Spacer ASME SA 240 Type 304 Nut ASME SA 193 Grade B8 Barrel ASTM A269 Grade TP 304 or ASME SA 312 Grade TP 304 or
ASME SA 240 Type 304 Collet Spring Inconel X-750 Ring Flange ASME SA 182 Grade F304.
The materials listed under ASTM specific ation number are all in the annealed condition (with the exception of the outer tube in the cylinder, tube and flange assembly), and their properties are re adily available. The outer tube is approximately 1/8 hard, and has a te nsile of 90,000/125,000 psi, yield of 50,000/85,000 psi, and minimum elongation of 25%.
The coupling spud, collet fingers and collet spring are fabricated from Inconel X-750 in the annealed or equalized condition, and heat treated to produce a tensile of 165,000 psi minimum, yield of 105,000 ps i minimum and elongation of 20%
minimum. The piston head is Armco 17-4 PH in condition H-1100, with a tensile of 140,000 psi minimum, yield of 115,000 ps i minimum and elongation of 15%
minimum.
These are widely used materials, whose properties are well known. All have been successfully used for the past 10 to 15 years in similar drive mechanisms. The parts are readily accessible for inspec tion, and replaceable if necessary.
4.5.1.2 Special Materials
No cold worked austenitic stainless steel s with a yield strength greater than 90,000 psi are employed in the control rod drive system. Hardenable martensitic stainless steels are not used. Armco 17-4 PH (precipitation hardened stainless steel) is used for the piston head. This material is aged to the H-1100 condition to produce resistance to stress corrosion cracking in the BWR environments. Armco 17-4 PH (H-1100) has been successfully used for the past 10 to 15 years in BWR drive mechanisms.
4.5.1.3 Processes, Inspections and Tests All austenitic stainless steel used in the control rod drive system is solution annealed material with one exception, the outer tube in the cylinder, tube, and
flange assembly (Subsection 4.5.1.1). Proper solution annealing is verified by LSCS-UFSAR 4.5-3 REV. 13 testing per ASTM-A262, "Recommended Prac tices for Detecting Susceptibility to Intergranular Attack in Stainless Steels." Two special processes are employed which subject selected components to temperatures in the sensitization range:
- a. The cylinder (cylinder, tube and flange assembly) and the retainer (collet assembly) are hard surfaced with Colmonoy 6.
- b. The following components are nitrided to provide a wear resistant surface:
- 1. tube (cylinder, tube and flange assembly),
- 2. piston tube (piston tube assembly), 3. index tube (drive line assembly), and
- 4. collet piston and guide cap (collet assembly).
Colmonoy hard surfaced components have performed successfully for the past 10 to 15 years in drive mechanisms. Nitrided components have accumulated 8 years of BWR service. It is normal practice to remove some control rod drives at each refueling outage. At this time, both the Colmonoy hard surfaced parts and nitrided surfaces are accessible for visual examination. In addition, dye penetrant examinations have been performed on nitrided surfaces of the longest service drives. This inspection program is adequate to detect any incipient defects before they could become serious enough to cause operating problems.
4.5.1.4 Control of Delta Ferrite Content
All Type 308 weld metal is purchased to a specification which requires a minimum of 5% delta ferrite. This amount of ferrite is adequate to prevent any microfissuring (hot cracking) in austenitic stainless steel welds.
4.5.1.5 Protection of Materials Duri ng Fabrication, Shipping and Storage
All the control rod drive parts listed previo usly (Subsection 4.5.1.1) are fabricated under a process specification which limits contaminants in cutting, grinding and tapping coolants and lubricants. It also restricts all other processing materials (marking inks, tape, etc.) to those which are completely removable by the applied cleaning process. All contaminants are then required to be removed by the appropriate cleaning process prior to any of the following:
- a. any processing which increases part temperature above 200° F, LSCS-UFSAR 4.5-4 REV. 13
- b. assembly which results in decre ase of accessibility for cleaning, or c. release of parts for shipment.
The specification for packaging and shipping the control rod drive provides the following.
The drive is rinsed in hot deionized water and dried in preparation for shipment.
The ends of the drive are then covered with a vapor-tight barrier with desiccant.
Packaging is designed to protect the drive and prevent damage to the vapor barrier. The planned storage period considered in the design of the container and packaging is 2 years. This packaging has been qualified and in use for a number of years. Periodic audits have indicated satisfactory protection.
Site or warehouse storage specifications require inside heated storage comparable to level B of ANSI 45.2.2.
4.5.2 Reactor
Internals Materials 4.5.2.1 Material Specifications
Materials used for steam dryer and core structure are as follows:
Plate, Sheet and Strip ASTM A240 Type 304 Bolts ASTM A193 Grade B8 Nuts ASTM A194 Grade 8 Forgings ASTM A182 Grade F304 Bar ASTM A276 Type 304 Bar ASTM A479 Type 304 Pipe ASTM A312 Grade TP 304 Tube ASTM A269, A249, or A213 Grade TP 304 Pipe Fittings ASTM A403 Grade WPW 304 or WP 304 Pipe Fittings (cast) ASTM A351 Grade CF8
The following materials are employed in other reactor internal structures:
- a. Steam Separator. All materials are Type 304, 304L, or 316L stainless steel Plate, Sheet and Strip ASTM A240, Type 304 LSCS-UFSAR 4.5-5 REV. 14, APRIL 2002 Forgings ASTM A182, Grade F304 Bars ASTM A479 Type 304 Pipe ASTM A312 Grade TP 304 Tube ASTM A269 Grade TP 304 Bolting Material ASTM A193 Grade B8 Nuts ASTM A194 Grade 8 Castings ASTM A351 Grade CF8
- b. Jet Pump Assemblies. The components in the jet pump assemblies are a riser, inlet, mixer, diffuser, adaptor, and brackets. All these components are fabricated with Type 304 stainless steel to the following specifications:
Castings ASTM A351 Grade CF8 Bars ASTM A276 Type 304 Bolts ASTM A193 Grade B8 or B8M Sheet and Plate ASTM A240 Type 304 Tubing ASTM A269 Grade TP 304 Pipe ASTM A358 Type 304 and ASTM A312 Grade TP304 Weld Coupling ASTM A403 Grade WP304 Forgings ASTM A182 Grade F304 Auxiliary Wedges The frames ar e fabricated from Type 304, 304L, 316, or 316L stainless steel.
The sliding components are fabricated from XM-19 or Alloy X-750.
Slip Joint Clamps The clamp frames are fabricated per ASTM A-182 Grade F XM-19. The sub-components are fabricated per ASTM B-637 UNS N07750 Type 3. Due to damage repaired during L1R08, the following unique features are associated with Unit 1 jet pump 9.
- The damaged "Stelllite-6" hard faced surface on the restrainer bracket pad was removed.
- Two auxiliary wedges are located on the riser restrainer bracket. The frames are fabricated from ASTM A-240 or A-479 Type 304 stainless steel (0.02%
max.
LSCS-UFSAR 4.5-5a REV. 16, APRIL 2006
- carbon) and the sliding component is fabricated from Alloy X-750 in accordance with ASTM B-637 UNS N077550 Type 3;
- The replacement inlet mixer wedge is fabricated from ASTM A-240 or A-479 Type 304 stainless steel (0.02% max ca rbon). Both of the wedge bearing surfaces are hard faced with "Stellite-21".
- In L2R10, inlet-mixer wedges and mounting hardware fabricated from Alloy X-750 and solution heat treated 300 series austenitic stainless steel (0.02% max. carbon) materials were installed in all of the Unit 2 jet pumps.
- During L1R11, jet pump riser brace clamps were installed on Unit 1 jet pumps 5/6 and 9/10 to mitigate crack indications by structurally replacing the upper and lower riser brace yoke to riser pipe welds designated as RS-8 and RS-9. The clamp components are fabricated from ASME SA-479/ASTM A479, ASME SA-240/ASTM A240, or ASME SA-182/ASTM A182 Type 316 stainless steel. The bolting components are fabricated from ASME SA-479/ASTM A479, or ASME SA-240/ASTM A240 Type XM-19 stainless steel.
The ratchet springs and nuts are fabricated from ASME SB-670/ASTM B-637 Grade UNS N07750, Type 3 Alloy X-750.
LSCS-UFSAR 4.5-6 REV. 15, APRIL 2004 Identification and justification for using materials in the jet pump assemblies which are not included in Appendix I to Section III of ASME B&PV Code are provided as follows: a. The inlet mixer adaptor casting, the wedge casting, bracket casting adjusting screw, and the diffuser collar casting are Type 304 hard surfaced with Stellite 6 for slip fit joints.
- b. The adaptor is a bimetallic component made by welding a Type 304 forged ring to a forged Inconel 600 ring, made to Specification ASTM B166.
- c. The inlet contains a pin, insert, and beam made of Inconel X-750 to Specification ASTM B637 Grade 688 or UNS N07750 Type 3 (beam), and ASTM A370 Grade E 38 and E55 (pin and insert).
- d. The jet pump beam bolt is stainless steel Type 316L.
- e. The jet pump beam keeper, s crews, plate and pins are 304L, XM-19, or X-750.
4.5.2.2 Controls on Welding
All welding of the reactor internals is performed in accordance with the ASME Section IX B&PV Code. Interpass temperature does not exceed 370° F. Processes used are GTAW, SMAW, GMAW, and SAW. All welds except intermittent and tack welds are
examined by liquid penetrant in accordance with ASME Section III. All welding filler material has a minimum of 5% ferrite as determined by the Schaeffler diagram.
4.5.2.3 Nondestructive Examination of Wrought Seamless Tubular Products
Wrought seamless tubular products were supplied in accordance with the applicable ASTM/ASME material specifications. These specifications require a hydrostatic test on each length of tubing. No special NDT was performed on the tubes.
4.5.2.4 Fabrication and Processing of Austenitic Stainless Steel
All materials have been solution heat treated and either water or air quenched.
Where an air cool was used, a sample of each heat and heat treatment lot was tested in accordance with ASTM A262 practice A or E. There was no heating above 800° F after the final heat treatment, except for thermal cutting or welding.
4.5.2.5 Regulatory Guide Conformance Assessment
This information is addressed in Appendix B of the FSAR.
LSCS-UFSAR 4.6-1 REV. 13 4.6 FUNCTIONAL DESIGN OF REACTIVITY CONTROL SYSTEMS
4.6.1 Information
for Control Rod Drive Systems (CRDS)
4.6.1.1 Control Rod Drive System Design 4.6.1.1.1 Design Bases
4.6.1.1.1.1 General Design Bases
4.6.1.1.1.1.1 Safety Design Bases
The control rod drive mechanical system meets the following safety design bases:
- a. Design provides for a sufficiently rapid control rod insertion so that no fuel damage results from any abnormal operating transient.
- b. Design includes positioning devices, each of which individually supports and positions a control rod.
- c. Each positioning device:
- 1. prevents its control rod from initiating withdrawal as a result of a single malfunction; collet piston stuck in upper position or stuck open withdraw valve will allow drive to continue withdrawal if initiating signal already given (Subsection 4.6.2.3);
- 2. is individually operated so that a failure in one positioning device does not affect the operation of any
other positioning device;
- 3. is individually hydraulically energized when rapid control rod insertion (scram) is signaled so that failure of power sources external to the positioning device does not prevent other positioning devices' control rods from being inserted; and 4. is locked to its control rod to prevent undesirable separation.
LSCS-UFSAR 4.6-2 REV. 14, APRIL 2002 4.6.1.1.1.1.2 Power Generation Design Basis The control rod system drive design provid es for positioning the control rods to control power generation in the core.
4.6.1.1.2 Description The control rod drive system (CRDS) controls gross changes in core reactivity by incrementally positioning neutron absorbing control rods within the reactor core in response to manual control signals. It is also required to quickly shut down the reactor (scram) in emergency situations by rapidly inserting withdrawn control rods into the core in response to a manual or automatic signal. The control rod drive system consists of locking piston, control rod drive mechanisms, and the CRD
hydraulic system (including hydraulic control units, interconnecting piping, instrumentation, and electrical controls).
4.6.1.1.2.1 Control Rod Drive Mechanisms
The CRD mechanism (drive) used for positioning the control rod in the reactor core is a double-acting, mechanically latched, hydraulic cylinder using water as its operating fluid. (See Figures 4.6-1, 4.6-2, 4.6-3, and 4.
6-4.) The individual drives are mounted on the bottom head of the reactor pressure vessel. The drives do not interfere with refueling and are operative even when the head is removed from the reactor vessel. The drives are also readily accessible for inspection and servicing. The bottom location makes maximum utilization of the water in the reactor as a neutron shield and gives the least possible neutron exposure to the drive components. Using water from the condensate storage tank as the operating fluid eliminates the need for special hydraulic fluid. Drives are able to utilize simple piston seals whose leakage does not contaminate the reactor water and does cool the drive mechanisms and their seals.
The drives are capable of inserting or withdrawing a control rod at a slow, controlled rate, as well as providing rapid insertion when required. A mechanism on the drive locks the control rod in 6-inch increments of stroke over the length of the core.
A coupling spud at the top end of the drive index tube (piston rod) engages and locks into a mating socket at the base of the control rod. The weight of the control rod is sufficient to engage and lock this coupling. Once locked, the drive and rod form an integral unit that must be manually unl ocked by specific procedures before components can be separated.
The drive holds its control rod in distinct latch positions until the hydraulic system actuates movement to a new position. Withdrawal of each rod is limited by the seating of the rod in its guide tube. Withdrawal to the overtravel limit can be LSCS-UFSAR 4.6-3 REV. 18, APRIL 2010 accomplished only if the rod and drive ar e uncoupled and will re sult in a control room alarm.
The individual rod indicators, grouped in one large core map control panel display, correspond to relative rod locations in the core. For display purposes the control rods are considered in groups of four adjacent rods centered around a common core volume. Each group is monitored by fo ur LPRM strings (Subsection 7.7.6). Rod groups at the periphery of the core may have less than four rods.
A Rod Select Display and a Status Display are located below the core map display. The rod select display is a touch-screen LC D that provides the operational interface used to select and perform the movement of a control rod. The status display is a touchscreen LCD display mounted directly below the core map display. The status display is capable of providing the same indications as the rod select display. The status display also serves as a back-up to the flat panel touchscreen rod select display in the event of component failure. A selected rod is indicated on all three displays.
4.6.1.1.2.2 Drive Components
Figure 4.6-2 illustrates the operating principle of a drive. Figures 4.6-3 and 4.6-4 illustrate the drive in more detail. The main components of the drive and their functions are described in the following paragraphs.
4.6.1.1.2.2.1 Drive Piston The drive piston is mounted at the lower end of the index tube. This tube functions as a piston rod. The drive piston and index tube make up the main moving assembly in the drive. The drive piston operates between positive end stops, with a hydraulic cushion provided at the upper end only. The piston has both inside and outside seal rings and operates in an annular space between an inner cylinder (fixed piston tube) and an outer cylinder (drive cy linder). Because the type of inner seal used is effective in only one direction, the lower sets of seal rings are mounted with one set sealing in each direction.
A pair of nonmetallic bushings prevents metal-to-metal contact between the piston assembly and the inner cylinder surface. The outer piston rings are segmented step-cut seals with expander springs ho lding the segments against the cylinder wall. A pair of split bushings on the outside of the piston prevents piston contact with the cylinder wall. The effective piston area for downtravel, or withdrawal, is approximately 1.2 in 2 vs. 4.1 in 2 for uptravel, or insertion. This difference in driving area tends to balance the control rod weight and assures a higher force for insertion than for withdrawal.
LSCS-UFSAR 4.6-4 REV. 13 4.6.1.1.2.2.2 Index Tube The index tube is a long hollow shaft made of nitrided Type 304 stainless steel. Circumferential locking grooves, spaced every 6 inches along the outer surface, transmit the weight of the control rod to the collet assembly.
4.6.1.1.2.2.3 Collet Assembly
The collet assembly serves as the index tube locking mechanism. It is located in the upper part of the drive unit. This as sembly prevents the index tube from accidentally moving downward. The assembly consists of the collet fingers, a return spring, a guide cap, a collet housing (part of the cylinder, tube, and flange), and the collet piston. LaSalle is the first domestic facility which contains the redesigned collet retainer tube. The collet retainer tube is fabricated from cast American Society for Testing and Materials A 351 CF-3 alloy with Colmonoy hardfacing, and
the index tube and piston tube ar e fabricated from XM-19 alloy.
Locking is accomplished by fingers mounted on the collet piston at the top of the drive cylinder. In the locked or latched position the fingers engage a locking groove in the index tube.
The collet piston is normally held in the latched position by a force of approximately 150 pounds supplied by a spring. Metal piston rings are used to seal the collet
piston from reactor vessel pressure. Th e collet assembly will not unlatch until the collet fingers are unloaded by a short, automatically sequenced, drive-in signal. A
pressure, approximately 180 psi above reactor vessel pressure, must then be applied to the collet piston to overcome spring force, slide the collet up against the conical surface in the guide cap, and spread the fingers out so they do not engage a locking
groove. A guide cap is fixed in the upper end of th e drive assembly. This member provides the unlocking cam surface for the collet fingers and serves as the upper bushing for the index tube.
If reactor water is used during a scram to supplement accumulator pressure, it is drawn through a filter on the guide cap.
4.6.1.1.2.2.4 Piston Tube
The piston tube is an inner cylinder, or column, extending upward inside the drive piston and index tube. The piston tube is fixed to the bottom flange of the drive and remains stationary. Water is brought to the upper side of the drive piston through this tube. A series of orif ices at the top of the tube provides progressive water shutoff to cushion the drive piston at the end of its scram stroke.
LSCS-UFSAR 4.6-5 REV. 13 4.6.1.1.2.2.5 Stop Piston A stationary piston, called the stop piston, is mounted on the upper end of the piston tube. This piston provides the seal between reactor vessel pressure and the space above the drive piston. It also functions as a positive end stop at the upper limit of control rod travel. A stack of spring washers just below the stop piston helps absorb the final mechanical shock at th e end of control rod travel. The piston rings are similar to the drive piston outer rings. A bleed-off passage to the center of the piston tube is located between the two pairs of rings. This arrangement allows seal leakage from the reactor vessel (during a scram) to be bled directly to the discharge line. The lower pair of seals is used only during the cushioning of the drive piston at the upper end of the stroke.
The center tube of the drive mechanism forms a well to contain the position indicator probe. This probe is an aluminum extrusion attached to a cast aluminum housing. Mounted on the extrusion are he rmetically sealed, magnetically operated, position indicator switches. Each switch is sheathed in a braided glass sleeve, and the entire probe assembly is protected by a thin-walled stainless steel tube. The switches are actuated by a ring magnet located at the bottom of the drive piston.
The drive piston, piston tube, and indicator tube are all of nonmagnetic stainless steel, allowing the individual switches to be operated by the magnet as the piston passes. One switch is located at each position corresponding to an index tube groove, thus allowing indication at each la tching point. An additional switch is located at each midpoint between latching points to indicate the intermediate positions during drive motion. Thus, indica tion is provided for each 3 inches of travel. Duplicate switches are provided for the full-in and full-out postions. One additional switch (an overtravel switch) is located at a position below the normal full-out position. Because the limit of do wntravel is normally provided by the control rod itself as it reaches the backseat position, the drive can pass this position and actuate the overtravel switch only if it is uncoupled from its control rod. A convenient means is thus provided to verify that the drive and control rod are coupled after installation of a drive or at any time during plant operation.
4.6.1.1.2.2.6 Flange and Cylinder Assembly
A flange and cylinder assembly is made up of a heavy flange welded to the drive cylinder. A sealing surface on the upper face of this flange forms the seal to the
drive housing flange. The seals contain reactor pressure and the two hydraulic control pressures. Teflon coated, stainless steel rings are used for these seals. The drive flange contains the integral ball, or two-way, check (ball-shuttle) valve. This valve directs either the reactor vessel pressure or the driving pressure, whichever is higher, to the underside of the drive piston. Reactor vessel pressure is admitted to this valve from the annular space between the drive and drive housing through passages in the flange.
LSCS-UFSAR 4.6-6 REV. 13 Water used to operate the collet piston passes between the outer tube and the cylinder tube. The inside of the cylinder tube is honed to provide the surface required for the drive piston seals.
Both the cylinder tube and outer tube are welded to the drive flange. The upper ends of these tubes have a sliding fit to allow for differential expansion.
The upper end of the index tube is threaded to receive a coupling spud. The coupling (Figure 4.6-1) accommodates a small amount of angular misalignment between the drive and the control rod. Six spring fingers allow th e coupling spud to enter the mating socket on the control rod. A plug then enters the spud and prevents uncoupling.
4.6.1.1.2.2.7 Lock Plug
Two means of uncoupling are provided. With the reactor vessel head removed, the lock plug can be raised against the spring force of approximately 50 pounds by a rod
extending up through the center of the control rod to an unlocking handle located above the control rod velocity limiter. The control rod, with the lock plug raised, can then be lifted from the drive.
The lock plug can also be pushed up from below, if it is desired to uncouple a drive without removing the reactor pressure vessel head for access. In this case, the central portion of the drive mechanism is pushed up against the uncoupling rod assembly, which raises the lock plug and allows the coup ling spud to disengage the socket as the drive piston and index tube are driven down.
The control rod is heavy enough to force the spud fingers to enter the socket and push the lock plug up, allowing the spud to enter the socket completely and the plug to snap back into place. Therefore, the drive can be coupled to the control rod using only the weight of the control rod. However, with the lock plug in place, a force in excess of 50,000 pounds is required to pull the coupling apart.
4.6.1.1.2.3 Materials of Construction Factors that determine the choice of construction materials are discussed in the following subsections.
4.6.1.1.2.3.1 Index Tube
The index tube must withstand the locking and unlocking action of the collet fingers. A compatible bearing combinat ion must be provided that is able to withstand moderate misalignment forces.
The reactor environment limits the choice of materials suitable for corrosion resistance. The column and tensile loads LSCS-UFSAR 4.6-7 REV. 13 can be satisfied by an anne aled AISI-300 series stainless steel. The wear and bearing requirements are provided by Malc omizing the complete tube. To obtain suitable corrosion resistance, a carefully controlled process of surface preparation is employed.
4.6.1.1.2.3.2 Coupling Spud The coupling spud is made of Inconel-750 that is aged for maximum physical strength and the required corrosion resistance. Because misalignment tends to cause chafing in the semispherical contact area, the part is protected by a thin chromium plating (Electrolized). This plating also prevents galling of the threads attaching the coupling spud to the index tube.
4.6.1.1.2.3.3 Collet Fingers
Inconel-750 is used for the collet fingers, which must function as leaf springs when cammed open to the unlocked position. Colmonoy 6 hard facing provides a long wearing surface, adequate for design life, to the area contacting the index tube and unlocking cam surface of the guide cap.
4.6.1.1.2.3.4 Seals and Bushings
Graphitar 14 is selected for seals and bushings on the drive piston and stop piston. The material is inert and has a low friction coefficient when water lubricated.
Because some loss of Graphitar strength is experienced at higher temperatures, the drive is supplied with cooling wate r to hold temperatures below 250
° F. The Graphitar is relatively soft, which is adva ntageous when an occasional particle of foreign matter reaches a seal. The resulting scratches in the seal reduce sealing efficiency until worn smooth, but the drive design can tolerate considerable water leakage past the seals into the reactor vessel.
4.6.1.1.2.3.5 Summary
All drive components exposed to reactor vessel water are made of AISI-300 series stainless steel except the following:
- a. Seals and bushings on the drive piston and stop piston are Graphitar 14.
- b. All springs and members requiring spring action (collet fingers, coupling spud, and spring washer s) are made of Inconel-750.
- c. The ball check valve is a Haynes Stellite cobalt-base alloy.
- d. Elastomeric O-ring seals are ethylene propylene.
LSCS-UFSAR 4.6-8 REV. 13 e. Collet piston rings are Haynes 25 alloy.
- f. Certain wear surfaces are hard-faced with Colmonoy 6.
- g. Nitriding by a proprietar y new Malcomizing process and chromium plating are used in certain areas where resistance to abrasion is necessary.
- h. The drive piston head is made of Armco 17-4PH.
Pressure-containing portions of the drives are designed and fabricated in accordance with requirements of Sect ion III of the ASME Boiler and Pressure Vessel Code.
4.6.1.1.2.4 Control Rod Drive Hydraulic System
The control rod drive hydraulic system (Drawing Nos. M-100 and M-146) controls the pressure and flow to and from the driv es through hydraulic control units (HCU).
The water discharged from the drives during a scram flows through the HCU's to the scram discharge volume. The water discharged from a drive during a normal control rod positioning operation flows through the HCU into the exhaust header, a reverse flow then occurs from the exhaust header through the insert/exhaust directional solenoid valves (121) into the latched CRD's. There are as many HCU's as the number of control rod drives.
4.6.1.1.2.4.1 Hydraulic Requirements The CRD hydraulic system design is sh own in Drawing Nos. M-100 and M-146 and Figures 4.6-5 and 4.6-6. The hydraulic requirements, identified by the function they perform, are as follows:
- a. An accumulator hydraulic charging pressure of approximately 1400 to 1500 psig is required. Fl ow to the accumulators is required only during scram reset or system startup.
- b. Drive pressure of approximately 250 psi above reactor vessel pressure is required. A flow rate of approximately 4 gpm to insert a control rod and 2 gpm to withdraw a control rod is required.
- c. Cooling water to the drives is required at approximately 15 psi above reactor vessel pressure and at a flow rate of 0.20 to 0.34 gpm per drive unit. (Cooling water can be interrupted for short periods without damaging the drive.)
LSCS-UFSAR 4.6-9 REV. 14, APRIL 2002 d. The scram discharge volume is sized to receive and contain all the water discharged by the drives during a scram; a minimum volume of 3.34 gallons per drive is required.
- e. The CRD System provides approximately 0.05 gpm to the condensing chambers reference legs for the narrow range, wide range, and fuel zone reactor vessel level instrumentation (UFSAR Section 7.7.1.2.2).
4.6.1.1.2.4.2 System Description The CRD hydraulic systems provide the required functions with the pumps, filter, valves, instrumentation, and piping shown in Drawing Nos. M-100 and M-146 and described in the following paragraphs.
Duplicate components are included, where necessary, to ensure continuous system operation if an inservice component requires maintenance.
The control rod drive hydraulic system also supplies a purge flow to the reactor water cleanup pumps to prevent settling of sediment in the base of each of the two pumps. This flow is taken from the char ging water header and becomes part of the RWCU process fluid once it enters the pump. It is not returned to the CRD
hydraulic system. (Drawings M-97 an d M-143, Sheet 1, and M-100 and M-146, Sheet 1). This purge flow is not required for operation of the pumps.
4.6.1.1.2.4.2.1 Supply Pump One supply pump pressurizes the system.
The condensate system is the normal source of water from the hotwell reject lin
- e. However, during shutdown conditions, the pump suction is from the condensate st orage tank. One spare pump is provided for standby. A discharge check valve prevents backflow through the nonoperating pump. A portion of the pump discharge fl ow is diverted through a minimum flow bypass line to the condensate st orage tank. This flow is controlled by an orifice and is sufficient to prevent immediate pump damage if the pump discharge is inadvertently closed. An additional recirculation line is provided for the supply pumps. This line provides a means of maintaining the pump manufacturer's recommended minimum flow, during unit ou tage time periods when CRD system flow demand is minimal. Flow in this line is controlled by a severe service manual control valve, which is closed during normal plant operation. This line is used concurrently with the previously mentioned minimum flow bypass line to the condensate storage tank.
Condensate water is processed by two filters in the system. The pump suction filter is a cleanable element type with a 25-micron absolute rating. The drive water filter downstream of the pump is a cleanable element type with a 50-micron absolute rating. A differential pressure indicator and control room alarm monitor the filter element as it collects foreign material.
LSCS-UFSAR 4.6-10 REV. 13 4.6.1.1.2.4.2.2 Accumulator Charging Pressure Accumulator charging pressure is established by the discharge pressure of the system supply pump. During scram the scram inlet (and outlet) valves open and permit the stored energy in the accumulato rs to discharge into the drives. The resulting pressure decrease in the charging water header allows the CRD supply pump to run out (i.e., flow rate to increase substantially) into the control rod drives via the charging water header. The fl ow sensing system upstream of the accumulator charging header detects high flow and closes the flow control valve. This action maintains increased flow through the charging water header.
Pressure in the accumulator charging header is monitored in the control room with a pressure indicator and a low/high pre ssure alarm. An automatic scram is initiated when the charging water header pressure drops below 1157 psig for more than approximately 10 seconds.
The automatic scram on low pressure in the charging water header is not active in the run mode because the accumulators are not required for scram at operating pressures. The automatic scram is also no t active in the shutdown mode since no control rods may be withdrawn in this mode. In all other modes, the automatic scram on low charging-water-header pressure remains active.
During normal operation the flow control valve maintains a constant system flow rate. This flow is used for drive flow, drive cooling, and system stability.
4.6.1.1.2.4.2.3 Drive Water Pressure Drive water pressure required in the dr ive header is maintained by the drive pressure control valve, which is manually adjusted from the control room. A flow rate of approximately 6 gpm (the sum of the flow rate required to insert and withdraw a control rod) normally passe s from the drive water pressure stage through two solenoid-operated stabilizing valves (arranged in parallel) and then goes into the cooling water line. The fl ow through one stabilizing valve equals the drive insert flow; that of the other stabilizing valve equals the drive withdrawal flow. When operating a drive, the required flow is diverted to that drive by closing the appropriate stabilizing valve. Thus, flow through the drive pressure control valve is always constant.
Flow indicators in the drive water head er and in the line downstream from the stabilizing valves allow the flow rate thro ugh the stabilizing valves to be adjusted when necessary. Diff erential pressure between the reactor vessel and the drive pressure stage is indicated in the control room.
LSCS-UFSAR 4.6-11 REV. 13 4.6.1.1.2.4.2.4 Cooling Water Header The cooling water header is located downstream from the drive pressure control valve. When not moving a CRD, all system flow returns to vessel through the cooling water header.
The flow through the flow control valve is virtually constant. Therefore, once adjusted, the drive pressure control va lve maintains the required pressure independent of reactor pressure. Changes in setting of the pressure control valves are required only to adjust for changes in the cooling requirements of the drives, as their seal characteristics change with time. A flow indicator in the control room monitors cooling water flow. A differential pressure indicator in the control room indicates the difference between reactor vessel pressure and drive cooling water pressure. Although the drives can function without cooling water, seal life is shortened by long term exposure to reactor temperatures. The temperature of each drive is recorded in the control room, an d excessive temperatures are annunciated.
4.6.1.1.2.4.2.5 Return Line
The H 2O discharged from the HCU during a normal control rod positioning operation is discharged back to the RPV through the insert/exhaust directional solenoid valves of adjoining HCUs.
4.6.1.1.2.4.2.6 Scram Discharge Volume
The scram discharge volume consists of header piping which connects to each HCU and drains into an instrument volume. The header piping is sized to receive and
contain all the water discharged by the drives during a scram, independent of the instrument volume. Each header pipe is designed with a hydrolazing port having 3/4" threaded plugs to allow the lines to be flushed occasionally, to prevent radiation build-up. During normal plant operation the scram discharge volume is empty and vented to atmosphere through its open vent and drain valves. When a scram occurs, upon a signal from the safety circuit, these vent and drain valves are closed to conserve reactor water. Lights in the control room indicate the position of these
valves. During a scram, the scram di scharge volume partly fills with water discharged from above the drive pistons. While scrammed, the control rod drive seal leakage from the reactor continues to flow into the scram discharge volume until the discharge volume pressure equals the reactor vessel pressure. A check valve in each HCU prevents reverse flow from the scram discharge header volume to the drive. When the initial scram signal is cleared from the reactor pr otection system, the scram discharge volume signal is overridden with a keylock override switch, and the scram discharge volume is drained and returned to atmospheric pressure.
LSCS-UFSAR 4.6-12 REV. 15, APRIL 2004 Remote manual switches in the pilot valve solenoid circuits allow the discharge volume vent and drain valves to be tested without disturbing the reactor protection system. Closing the scram discharge volume valves allows the outlet scram valve seats to be leak tested by timing the accumulation of leakage inside the scram discharge volume.
There are two instrument volumes associat ed with the scram discharge volume. Four level switches and two analog trip systems connected to each instrument volume to monitor the volume for abnormal water level. Each analog trip system consists of a transmitter and a trip unit. The level switches are set at three different levels. At the lowe st level, a level switch actuates to indicate that the volume is not completely empty during post scram draining or to indicate that the volume starts to fill through leakage accumulation at other times during reactor operation. At the second level, one leve l switch produces rod withdrawal block to prevent further withdrawal of any cont rol rod when leakage accumulates to approximately half the capacity of the instrument volume. The remaining two level switches and the trip units are interconnected with the reactor protection system (RPS) trip channels and will initiate a reactor scram should water accumulation fill the instrument volume. The liquid level switches are float type and transmitters are differential pressure type. Each di fferential pressure transmitter/trip unit combinations are powered from separate ESS Division sources that are independent of the Reactor Protection system power supply.
Redundant Vent & Drain Valves, placed in series, are located in the vent and drain piping for the scram discharge volume.
This system configuration addresses the co ncerns identified in IE Bulletin No. 80-
- 17. 4.6.1.1.2.4.3 Hydraulic Control Units
Each hydraulic control unit (HCU) furnishes pressurized water on signal to a drive unit. The drive then positions its cont rol rod as required. Operation of the electrical system that supplies scram and normal control rod positioning signals to the HCU is described in Subsection 7.7.2.
Operation of the electrical system which supplies ATWS signals to the HCU is described in Subsection 7.6.5.
The basic components in each HCU are: manual, pneumatic, and electrical valves; an accumulator; related piping; electrical connections; filters; and instrumentation (Drawing Nos. M-100 and M-146 and Figure 4.6-7).
The components and their functions are described in the following paragraphs.
4.6.1.1.2.4.3.1 Insert Drive Valve
The insert drive valve is solenoid-operated and opens on an insert signal. The valve supplies drive water to the bottom side of the main drive piston.
LSCS-UFSAR 4.6-13 REV. 18, APRIL 2010 4.6.1.1.2.4.3.2 Insert Exhaust Valve
The insert exhaust valve also opens by so lenoid on an insert signal. The valve discharges water from above the drive piston to the exhaust water header.
4.6.1.1.2.4.3.3 Wi thdraw Drive Valve
The withdraw drive valve is solenoid-opera ted and opens on a withdraw signal. The valve supplies drive water to the top of the drive piston.
4.6.1.1.2.4.3.4 Withdraw Exhaust Valve The solenoid-operated withdraw exhaust valve opens on a withdraw signal and discharges water from below the main drive piston to the exhaust header. It also serves as the settle valve. The valve op ens following any normal drive movement (insert or withdraw) to allow the control rod and its drive to settle back into the
nearest latch position.
4.6.1.1.2.4.3.5 Speed Control Valves The speed control valves regulate the control rod insertion and withdrawal rates during normal operation. They are manually adjustable flow control valves used to regulate the water flow to and from the volume beneath the main drive piston. A correctly adjusted valve does not require readjustment except to compensate for changes in drive seal leakage.
4.6.1.1.2.4.3.6 Scram Pilot Valves
The scram pilot valves are operated from the reactor protection system trip system. Either a single scram pilot valve with dual solenoid operated pilot assemblies or two single scram pilot valve assemblies control both the scram inlet valve and the scram exhaust valve. The scram pilot valve (either with dual solenoid operated pilot assemblies or with a single pilot solenoid assembly) are solenoid-operated, normally energized valves. On loss of electrical sign al to the scram pilot valve solenoids, such as the loss of external a-c power, the inlet port(s) close and the exhaust port(s) open on both scram pilot valve solenoids. Th e scram pilot valves (Drawing M-100 and M-146) are arranged so that the trip system signal must be removed from both scram pilot valve solenoids before air pressure can be discharged from the scram valve operators. This prevents the inadvertent scram of a single drive in the event of a failure of one of the scram pilot valve solenoids.
LSCS-UFSAR 4.6-13a REV. 18, APRIL 2010 4.6.1.1.2.4.3.7 Scram Inlet Valve
The scram inlet valve opens to supply pressurized water to the bottom of the drive piston. This quick opening globe valve is op erated by an internal spring and system pressure. It is closed by air pressure applied to the top of its diaphragm operator.
LSCS-UFSAR 4.6-14 REV. 18, APRIL 2010 A position indicator switch on this valve pr ovides indication in the control room as soon as the valve starts to open.
As the scram inlet valve and the scram exhaust valve start to open, position indication switches on the valves initiate "valve open" indication in the main control room. 4.6.1.1.2.4.3.8 Scram Exhaust Valve
The scram exhaust valve opens slightly before the scram inlet valve, exhausting water from above the drive piston. The exhaust valve opens faster than the inlet valve because of a high air pressure spring setting in the valve operator. Otherwise the valves are similar.
4.6.1.1.2.4.3.9 Scram Accumulator
The scram accumulator stores sufficient energy to fully insert a control rod at lower vessel pressures. At higher vessel pressures the accumulator pressure is assisted or supplanted by reactor vessel pressure. The accumulator is a hydraulic cylinder with a free-floating piston. The piston separates the water on top from the nitrogen below. A check valve in the accumulator charging line prevents loss of water pressure in the event supply pressure is lost.
During normal plant operation, the accumulator piston is seated at the bottom of its cylinder. Loss of nitrogen decreases th e nitrogen pressure, which actuates a pressure switch and sounds an alarm in the control room. To ensure that the accumulator is always able to produce a sc ram, it is continuously monitored for water leakage. A float-type level switch actuates an alarm if water leaks past the piston barrier and collects in the accumulator instrumentation block.
4.6.1.1.2.4.3.10 Alternate Rod Insertion Scram Valves
The alternate rod insertion (ARI) scram valves are redundant to the existing RPS scram backup valves C11-F110A&B, and scram discharge volume vent and drain pilot valves C11-F379 & F387. The ARI valves provide an alternate means of initiating control rod insertion during an ATWS event. The ARI valves have direct current solenoid dual coil operators. The valves are provided with position switches to indicate valve open/closed status in the main control room. The valves perform three functions during an ATWS trip:
- 1. Block the instrument air supply line to the pilot scram valves.
LSCS-UFSAR 4.6-15 REV. 13 3. Exhaust air header to the scram discharge volume vent and drain valves, permitting these valves to close.
4.6.1.1.2.5 Control Rod Drive System Operation The control rod drive system performs rod insertion, rod withdrawal, and scram.
These operational functions are described as follows.
4.6.1.1.2.5.1 Rod Insertion Rod insertion is initiated by a signal from the operator to the insert valve solenoids. This signal causes both insert valves to open. The insert drive valve applies reactor pressure plus approximately 90 psi to the bottom of the drive piston. The insert exhaust valve allows water from above the drive piston to discharge to the exhaust header.
As is illustrated in Figure 4.6-3, the lock ing mechanism is a ratchet-type device and does not interfere with rod insertion. The speed at which the drive moves is determined by the flow through the insert speed control valve, which is set for approximately 4 gpm for a shim speed (non scram operation) of 3 in/sec. During normal insertion, the pressure on the downstream side of the speed control valve is 90 to 100 psi above reactor vessel pressure. However, if the drive slows for any reason, the flow through and pressure drop across the insert speed control valve will decrease; the full differential pressure (260 psi) will th en be available to cause continued insertion. With 260-psi differential pressure acting on the drive piston, the piston exerts an upward force of 1040 pounds.
4.6.1.1.2.5.2 Rod Withdrawal Rod withdrawal is, by design, more involved than insertion. The collet finger (latch) must be raised to reach the unlocked position (Figure 4.6-3). The index tube notches and the collet fingers are shaped so that the downward force on the index tube holds the collet fingers in place. The index tube must be lifted before the collet fingers can be released. This is done by opening the drive insert valves (in the manner described in the preceding paragr aph) for approximately 1 second. The withdraw valves are then opened, applying driving pressure above the drive piston and opening the area below the piston to the exhaust header. Pressure is simultaneously applied to the collet piston. As the piston raises, the collet fingers
are cammed outward, away from th e index tube, by the guide cap.
The pressure required to release the latch is set and maintained at a level high enough to overcome the force of the latch return spring plus the force of reactor pressure opposing movement of the collet piston. When this occurs, the index tube is unlatched and free to move in the withdraw direction. Water displaced by the drive piston flows out through the withdraw speed control valve, which is set to give LSCS-UFSAR 4.6-16 REV. 18, APRIL 2010 the control rod a shim speed of 3 in/sec. The maximum control rod drive withdrawal speed is 6.0 in/sec when the Operating Limit MCPR established in the Core Operating Limits Report (COLR) is set greater than or equal to the value corresponding to a RWE - at Power analysis for an "unblocked" condition (References 6 and 7). Otherwise, the maximum control rod drive withdrawal speed is 3.6 in/sec. See subsection 15.4.2.3 fo r additional details. The entire valving sequence is automatically controlled and is initiated by a single operation of the rod withdraw switch.
4.6.1.1.2.5.3 Scram
During a scram the scram pilo t valves and scram valves ar e operated as previously described. With the scram valves open, accumulator pressure is admitted under the drive piston, and the area over the drive piston is vented to the scram discharge volume.
The large differential pressure (initially approximately 1500 psi and always several hundred psi, depending on reactor vessel pressure) produces a large upward force on the index tube and control rod. This force gives the rod a high initial acceleration and provides a large margin of force to overcome any possible friction. After the initial acceleration is achieved, the drive continues at a nearly constant velocity. This characteristic provides a high initial rod insertion rate. As the drive piston nears the top of its stroke, the piston seals close off the large passage (buffer orifices) in the stop piston tube, and the drive slows.
Prior to a scram signal the accumula tor in the hydraulic control unit has approximately 1450-1510 psig on the water side, and >
980 and <1200 psig on the nitrogen side. As the inlet scram valve opens, the full water side pressure is
available at the control rod drive acting on a 4.1 in 2 area. As CRD motion begins, this pressure drops to the gas-side pressure less line losses between the accumulator and the CRD. At low vessel pressures, the accumulator completely discharges with a resulting gas-side pres sure of approximately 575 psig. Reactor pressure provides the force necessary to scram the reactor when reactor pressure exceeds scram accumulator pressure.
The control-rod-drive accumulators are required to scram the control rod when the reactor pressure is low. When the reactor pressure is low, the accumulator retains sufficient stored energy to ensure the complete insertion of the control rod in the required time. The accumulator is not required in order to scram the control rod in time when the reactor is close to or at fu ll operating pressure. In this instance, the reactor pressure alone will scram the control rod in the required time. However, the accumulator does provide an additional energy boost to the reactor pressure in providing scram action at vessel pressu res less than accumulator pressures.
LSCS-UFSAR 4.6-17 REV. 14, APRIL 2002 The control rod drive system, with accumu lators, was designed to meet the scram time requirements specified in Technical Specification.
4.6.1.1.2.6 Instrumentation
The general functional requirements for the control rod drive are discussed in Subsection 4.6.1.1.2.4.1.
4.6.1.2 Control Rod Drive Housing Supports
4.6.1.2.1 Safety Objective
The control rod drive (CRD) housing su pports prevent any significant nuclear transient in the event a drive housing brea ks or separates from the bottom of the reactor vessel.
4.6.1.2.2 Safety Design Bases The CRD housing supports meet the following safety design bases:
- a. Following a postulated CRD housing failure, control rod downward motion is limited so that any resulting nuclear transient cannot be sufficient to cause fuel damage.
- b. The clearance between the CRD housings and the supports is sufficient to prevent vertical cont act stresses caused by thermal expansion during plant operation.
4.6.1.2.3 Description The CRD housing supports are shown in Fi gure 4.6-8. Horizontal beams are installed immediately below the bottom of the reactor vessel, between the rows of CRD housings. The beams are supported by brackets welded to the steel form liner of the drive room in the reactor support pedestal.
Hanger rods, approximately 10-feet long and 1-3/4-inches in diameter, are supported from the beams on stacks of disc springs. These springs compress approximately 2 inches under the design load.
LSCS-UFSAR 4.6-18 REV. 13 The support bars are bolted between the bottom ends of the hanger rods. The spring pivots at the top, and the beveled, loose-fitting ends on the support bars prevent substantial bending moment in th e hanger rods if the support bars are overloaded.
Individual grids rest on the support bars between adjacent beams. Because a single-piece grid would be difficult to handle in the limited work space and because it is necessary that control rod drives, position indicators, and incore instrumentation components be accessible for inspection and maintenance, each grid is designed for inplace assembly or disassembly. Each grid assembly is made from two grid plates, a clamp, and a bolt. The top part of the clamp guides the grid to its correct position directly below the respective CRD housing that it would support in the postulated accident.
When the support bars and grids are installe d, a gap of approximately 1-1/2 inch at room temperature is provided between th e grid and the bottom contact surface of the control rod drive flange. During system heatup, this gap is reduced by a net downward expansion of the housings with respect to the supports. In the hot operating condition, the gap is reduced approximately 1/4 inch.
In the postulated CRD housing failure, the CRD housing supports are loaded when the lower contact surface of the CRD flange contacts the grid. The resulting load is then carried by two grid plates, two support bars, four hanger rods, their disc springs, and two adjacent beams.
The American Institute of Steel Construction (AISC) Manual of Steel Construction , "Specification for the Design, Fabricatio n and Erection of Structural Steel for Buildings," was used in designing the CRD housing support system. However, to
provide a structure that absorbs as much energy as practical without yielding, the allowable shear, tension and bending stresses used 1.5 times the AISC allowable stresses.
For purposes of mechanical design, the postulated failure resulting in the highest forces is an instantaneous circumferential separation of the CRD housing from the reactor vessel, with an internal pressu re of 1086 psig (react or vessel operating pressure) acting on the area of the separated housing. The weight of the separated housing, control rod drive, and blade, plus the pressure of 1086 psig acting on the area of the separated housing, gives a fo rce of approximately 32,000 pounds. This force is multiplied by an impact factor that conservatively assumes the housing travels through a 1-1/2 inch gap before it contacts the supports. The total force of approximately 120,000 pounds is then treated as a static load in design.
LSCS-UFSAR 4.6-19 REV. 14, APRIL 2002 All CRD housing support subassemblies are fabricated of commonly available structural steel, except for the following items:
Material a. grid bars ASTM-A-441, b. disc springs Schnorr, Type BS-125-71-8, and c. hex bolts and nuts ASTM-A-307.
4.6.2 Evaluations
of the CRDS
4.6.2.1 Failure Mode and Effects Analysis Engineering standards for electrical and physical separation, a design with high safety factors, and the unitary design approach for the CRD modules using ASME standards have each contributed toward an effective and proven CRDS for the control and safe shutdown of BWR's designed by GE. An analysis of failure modes and effects has not been completed for the LSCS units because the CRDS design has a proven history beginning with Dresde n-1. Further analytical evaluations are believed to be of less value than the accrual of real operating data and the incorporation of generic improvements based on actual experience. LSCS utilized this approach in lieu of FMEA.
4.6.2.2 Protection from Common Mode Failures
Based on NEDO-10189, NEDO-10349, and NEDO-20626, General Electric concludes that the complete failure of the BWR control rod scram system due to common mode failure is of such extremely low probability that no change in BWR design to account for the event is warranted.
EGC does not believe the ATWS to be a credible event; nevertheless, the LSCS design includes three provisions to assist shutdown in this unlikely event: tripping of the recirculation pumps, scram discharge volume upgrades, and the addition of alternate rod insertion (ARI) and main steam isolation valve closure modifications. These modifications adequately prevent and, additionally, contribute to the
mitigation of ATWS events.
4.6.2.3 Safety Evaluation
4.6.2.3.1 Control Rod Drives 4.6.2.3.1.1 Evaluation of Scram Time
The rod scram function of the control rod drive system provides the negative reactivity insertion required by safety design basis in Subsec tion 4.6.1.1.1.1.1, LSCS-UFSAR 4.6-20 REV. 1 Item c, part 1. The scram time shown in the description is adequate as shown by the transient analyses of Chapter 15.0.
4.6.2.3.1.2 Analysis of Malfunction Relating to Rod Withdrawal
There are no known single malfunctions that cause the unplanned withdrawal of even a single control rod; providing initiating signal has not been given (Subsections 4.6.1.1.1.1.1, Item c, part 1, and 4.6.2.3.1.2.10). However, if multiple malfunctions are postulated, studies show that an unplanned rod withdrawal can occur at withdrawal speeds that vary with the combination of malfunctions postulated. In all cases the subsequent withdrawal speeds are less than that assumed in the rod drop accident analysis as discussed in Chapter 15.0. Therefore, the physical and radiological consequences of such rod withdrawals are less than those analyzed in the rod drop accident.
4.6.2.3.1.2.1 Drive Housin g Fails at Attachment Weld
The bottom head of the reactor vessel has a penetration for each control rod drive location. A drive housing is raised into position inside each penetration and fastened by welding. The drive is raised into the drive housing and bolted to a flange at the bottom of the housing. Th e housing material is seamless, Type 304 stainless steel pipe with a minimum tens ile strength of 75,000 psi. The basic failure considered here is a complete circumferential crack through the housing wall
at an elevation just below the J-weld.
Static loads on the housing wall include the weight of the drive and the control rod, the weight of the housing below the J-weld, and the reactor pressure acting on the
6-inch diameter cross-sectional area of the housing and the drive. Dynamic loading results from the reaction force during drive operation.
If the housing were to fail as described, the following sequence of events is foreseen.
The housing would separate from the vessel. The control rod, drive, and housing would be blown downward against the support structure by reactor pressure acting on the cross-sectional area of the housing and the drive. The downward motion of the drive and associated parts would be determined by the gap between the bottom of the drive and the support structure and by the deflection of the support structure under load. In the current design, maximum deflection is limited to 3.65 inches. If the collet were to remain latched, no further control rod ejection would occur (Reference 4); the housing would not drop far enough to clear the vessel penetration.
Reactor water would leak at a rate of approximately 220 gpm through the 0.03-inch diametral clearance between the housing and the vessel penetration.
If the basic housing failure were to occur while the control rod is being withdrawn (this is a small fraction of the total drive operating time) and if the collet were to stay unlatched, the following sequence of events is foreseen. The housing would LSCS-UFSAR 4.6-21 REV. 1 separate from the vessel. The drive and housing would be blown downward against the control rod drive housing support.
Calculations indicate that the steady-sta te rod withdrawal velocity would be 0.3 ft/sec. During withdrawal, pressure under the collet piston would be approximately 250 psi greater than the pressure over it.
Therefore, the collet would be held in the unlatched position until driving pressure was removed from the pressure-over port.
4.6.2.3.1.2.2 Rupture of Hydraulic Line(s) to Drive Housing Flange
There are three types of possible rupture of hydraulic lines to the drive housing flange: (1) pressure-under line break; (2) pressure-over line break; and (3) coincident breakage of both of these lines.
4.6.2.3.1.2.2.1 Pressu re-Under Line Break
For the case of a pressure-under line break, a partial or complete circumferential opening is postulated at or near the point where the line enters the housing flange. Failure is more likely to occur after anot her basic failure wherein the drive housing or housing flange separates from the reactor vessel. Failure of the housing, however, does not necessarily lead directly to failure of the hydraulic lines.
If the pressure-under line were to fail and if the collet were latched, no control rod withdrawal would occur. There would be no pressure differential across the collet piston and, therefore, no tendency to unlatch the collet. Consequently, the associated control rod could not be inserted or withdrawn.
The ball check valve is designed to seal off a broken pressure-under line by using reactor pressure to shift the check ball to its upper seat. If the ball check valve were prevented from seating, reactor water would leak to the atmosphere. Because of the broken line, cooling water could not be supplied to the drive involved. Loss of cooling water would cause no immediate dama ge to the drive. However, prolonged exposure of the drive to temperatures at or near reactor temperature could lead to deterioration of material in the seals. High temperature would be indicated to the operator by the thermocouple in the position indicator probe. A second indication would be high cooling water flow.
If the basic line failure were to occur while the control rod is being withdrawn, the hydraulic force would not be sufficient to hold the collet open, and spring force normally would cause the collet to latch and stop rod withdrawal. However, if the collet were to remain open, calculations indicate that the steady-state control rod withdrawal velocity would be 2 ft/sec.
LSCS-UFSAR 4.6-22 REV. 15, APRIL 2004 4.6.2.3.1.2.2.2 Pressure-Over Line Break The case of the pressure-over line breakage considers the complete breakage of the line at or near the point where it enters the housing flange. If the line were to break, pressure over the drive piston would drop from reactor pressure to atmospheric pressure. Any significant re actor pressure (approximately 600 psig or greater) would act on the bottom of the drive piston and fully insert the drive. Insertion would occur regardless of the operational mode at the time of the failure. After full insertion, reactor water would leak past the stop piston seals. This leakage would exhaust to the atmosphere through the broken pressure-over line.
The leakage rate of 1000 psi reactor pressu re is estimated to be 4 gpm nominal but not more than 10 gpm, based on experiment al measurements. If the reactor were hot, drive temperature would increase. This situation would be indicated to the reactor operator by the drift alarm, by th e fully inserted driv e, by a high drive temperature (indicated on a recorder in th e control room), and by operation of the drywell sump pump.
4.6.2.3.1.2.2.3 Simultaneous Breakage of the Pressure-Over and Pressure-Under Lines
For the simultaneous breakage of the pressure-over pressure-under lines, pressures above and below the drive piston would drop to zero, and the ball check valve would close the broken pressure-under line. Reactor water would flow from the annulus outside the drive, through the vessel ports, and to the space below the drive piston.
As in the case of pressure-over line breakage, the drive would then insert at a speed dependent on reactor pressure. Full insertion would occur regardless of the operational mode at the time of failure. Reactor water would leak past the drive seals and out the broken pressure-over line to the atmosphere, as described previously. Drive temperature would incr ease. Indication in the control room would include the drift alarm, the fully-inserted drive, the high drive temperature on a recorder in the control room, and operation of the drywell sump pump.
4.6.2.3.1.2.3 All Drive Flange Bolts Fail in Tension Each control rod drive is bolted to a flange at the bottom of a drive housing. The flange is welded to the drive housing.
The CRD mechanism is bolted to the CRD housing flange by 8 bolts. Each bolt ha s significantly high load carrying capacity compared to the actual load.
If a progressive or simultan eous failure of all bolts were to occur, the drive would separate from the housing. The control rod and the drive would be blown downward against the support structure. Impact velocity and support structure loading would be slightly less than that for drive housing failure because reactor pressure would LSCS-UFSAR 4.6-23 REV. 18, APRIL 2010 act on the drive cross-sectional area only and the housing would remain attached to the reactor vessel. The drive would be is olated from the cooling water supply. Reactor water would flow downward past the velocity limiter piston, through the large drive filter, and into the annular space between the thermal sleeve and the drive. For worst-case leakage calculations, the large filter is assumed to be deformed or swept out of the way so it would offer no significant flow restriction. At a point near the top of the annulus, wher e pressure would have dropped to 350 psi, the water would flash to steam and cause ch oke-flow conditions. Steam would flow down the annulus and out the space between the housing and the drive flanges to the atmosphere. Steam formation would limit the leakage rate to approximately 840 gpm.
If the collet were latched, control rod ejection would be limited to the distance the drive can drop before coming to rest on the support structure. There would be no tendency for the collet to unlatch because pressure below the collet piston would drop to zero. Pressure fo rces, in fact, exert 1435 pounds to hold the collet in the latched position.
If the bolts failed during control rod withdrawal, pressure below the collet piston would drop to zero. The collet, with 1650 pounds return force, wo uld latch and stop rod withdrawal.
4.6.2.3.1.2.4 Weld Joining Flange to Housing Fails in Tension
The failure considered is a crack in or near the weld that joins the flange to the housing. This weld extends through the wall and completely around the housing. The flange material is forged, Type 304 stainless steel, with a minimum tensile strength of 75,000 psi. The housing materi al is seamless, Type 304 stainless steel pipe, with a minimum tensile strength of 75,000 psi. The conventional, full-penetration weld of Type 308 stainless steel has a minimum tensile strength approximately the same as that for the parent metal. The design pressure and temperature are 1250 psig and 575
° F. Reactor pressure acting on the cross-sectional area of the drive, the weight of the control rod, drive, and flange, and the dynamic reaction force during drive operation result in a maximum tensile stress at the weld of approximately 6000 psi.
If the basic flange-to-housing joint failure occurred, the flange and the attached drive would be blown downward against the support structure. The support structure loading would be slightly less than that for drive housing failure because reactor pressure would act only on the drive cross-sectional area. Lack of differential pressure across the collet piston would cause the collet to remain latched and limit control rod motion to ap proximately 3.65 inches. Downward drive movement would be small and, therefore, most of the drive would remain inside the housing. The pressure-under and pressu re-over lines are flexible enough to withstand the small displacement and rema in attached to the flange. Reactor LSCS-UFSAR 4.6-24 REV. 18, APRIL 2010 water would follow the same leakage path described above for the flange-bolt failure, except that exit to the atmosphere would be through the gap between the lower end of the housing and the top of the flange. Water would flash to steam in the annulus surrounding the drive. Th e leakage rate would be approximately 840 gpm.
If the basic failure were to occur during co ntrol rod withdrawal (a small fraction of the total operating time) and if the colle t were held unlatched, the flange would separate from the housing. The drive an d flange would be blown downward against the support structure. The calculated steady-state rod withdrawal velocity would
be 0.13 ft/sec. Because pressure-under and pressure-over lines remain intact, driving water pressure would continue to the drive, and the normal exhaust line restriction would exist. The pressure below the velocity limiter piston would drop below normal as a result of leakage fr om the gap between the housing and the flange. This differential pressure across the velocity limiter piston would result in a net downward force of approximately 70 po unds. Leakage out of the housing would greatly reduce the pressure in the annulus surrounding the drive. Thus, the net downward force on the drive piston would be less than normal. The overall effect of these events would be to reduce rod withdrawal to approximately one-half of normal
speed. With a 560-psi differential across the collet piston, the collet would remain
unlatched; however, it should relatch as soon as the drive signal is removed.
4.6.2.3.1.2.5 Housing Wall Ruptures
This failure is a vertical split in the drive housing wall just below the bottom head of the reactor vessel. The flow area of the hole is considered equivalent to the annular area between the drive and the thermal sleeve. Thus, flow through this annular area, rather than flow through the hole in the housing, would govern leakage flow. The housing is made of Ty pe 304 stainless steel seamless pipe, with a minimum tensile strength of 75,000 psi. The maximum hoop stress of 11,900 psi results primarily from the reactor design pr essure (1250 psig) acting on the inside of the housing.
If such a rupture were to occur, reactor water would flash to steam and leak
through the hole in the housing to the atmosphere at approximately 1030 gpm.
Choke-flow conditions would exist as described previously for the flange-bolt failure. However, leakage flow would be greater because flow resistance would be less; that is, the leaking water and steam would not have to flow down the length of the housing to reach the atmosphere. A critica l pressure of 350 psi causes the water to flash to steam.
No pressure differential across the collet piston would tend to unlatch the collet; but the drive would insert as a re sult of loss of pr essure in the drive housing causing a pressure drop in the space above the drive piston.
LSCS-UFSAR 4.6-25 REV. 18, APRIL 2010 If this failure occurred during control rod withdrawal, drive withdrawal would stop, but the collet would remain unlatched. The drive would be stopped by a reduction of the net downward force action on the drive line. The net force reduction would occur when the leakage flow of 1030 gpm reduces the pressure in the annulus outside the drive to approximately 540 psig , thereby reducing the pressure acting on top of the drive piston to the same value. A pressure differential of approximately 710 psi would exist across the collet piston and hold the collet
unlatched as long as the operator held the withdraw signal.
4.6.2.3.1.2.6 Flange Plug Blows Out
To connect the vessel ports with the bottom of the ball check valve, a hole of 3/4-inch diameter is drilled in the drive flange. The outer end of this hole is sealed with a plug of 0.812-inch diameter and 0.25-inch thickness. A full-penetration, Type 308 stainless steel weld holds the plug in place. The postulated failure is a full circumferential crack in this weld and subsequent blowout of the plug.
If the weld were to fail, the plug were to blow out, and the collet remained latched, there would be no control rod motion.
There would be no pressure differential across the collet piston acting to unlatch the collet. Reactor water would leak past the velocity limiter piston, down the a nnulus between the drive and the thermal sleeve, through the vessel ports and drilled passage, and out the open plug hole to the atmosphere at approximately 320 gpm.
Leakage calculations assume only liquid flows from the flange. Actually, hot reactor water would flash to steam and choke-flow conditions would exist. Thus, the expected leakage rate would be lower than the calculated value. Drive temper ature would increase and initiate an alarm in the control room.
If this failure were to occur during control rod withdrawal and if the collet were to stay unlatched, calculations indicate that control rod withdrawal speed would be
approximately 0.24 ft/sec. Leakage from the open plug hole in the flange would cause reactor water to flow downward past the velocity limiter piston. A small differential pressure across the piston would result in an insignificant driving force of approximately 10 pounds, tending to increase withdraw velocity.
A pressure differential of 295 psi across the collet piston would hold the collet unlatched as long as the driving signal was maintained.
Flow resistance of the exhaust path from the drive would be normal because the ball check valve would be seated at the lower end of its travel by pressure under the drive piston.
LSCS-UFSAR 4.6-26 REV. 18, APRIL 2010 4.6.2.3.1.2.7 Drive Pressure Control Va lve Closure (Reactor Pressure, 0 psig)
The pressure to move a drive is generated by the pressure drop of practically the full system flow through the drive pressure control valve. This valve is a motor-operated valve with a normally closed, stan dby manually operated valve in parallel. The motor-operated valve is adjusted to a fixed opening, to develop a normal pressure (260 psig in excess of normal reactor pressure) on the upstream side of the motor-operated valve. In the event of mo tor-operated valve failure, this valve can be isolated (upstream and downstream gate valves) and its functi on replaced by the manually operated standby valve.
If the flow through the drive pressure contro l valve were to be stopped, as by a valve closure or flow blockage, the drive pressure would increase to the shutoff pressure of the supply pump. The occurrence of this co ndition during withdrawal of a drive at zero vessel pressure will result in a drive pressure increase from 260 psig to no more than 1700 psig. Calculations indicate that the drive would accelerate from a nominal 3 in/sec to approximately 6 in/sec. A pressure differential of 1670 psi across the collet piston would hold the collet unlatched. Flow would be upward, past the velocity limiter piston, but retarding force would be negligible. Rod movement would stop as soon as the driving signal was removed.
4.6.2.3.1.2.8 Ball Check Valve Fails to Close Passage to Vessel Ports
Should the ball check valve sealing the passage to the vessel ports be dislodged and prevented from reseating following the insert portion of a drive withdrawal sequence, water below the drive piston would return to the reactor through the vessel ports and the annulus between the drive and the housing rather than through the speed control valve. Because the flow resistance of this return path would be lower than normal, the calculated withdrawal speed would be 2 ft/sec. During withdrawal, differential pressure across the collet piston would be approximately 40 psi. Therefore, the collet would tend to latch and would have to stick open before continuous withdrawal at 2 ft/sec, could occur. Water would flow upward past the velocity limiter piston, generating a small retarding force of approximately 120 pounds.
4.6.2.3.1.2.9 Hydraulic Cont rol Unit (HCU) Valve Failures
Various failures of the valves in the HCU can be postulated, but none could produce differential pressures approaching those described in the preceding paragraphs and
none alone could produce a high velocity withdrawal. Leakage through either one or both of the scram valves produces a pressure that tends to insert the control rod rather than to withdraw it. If the pressure in the scram discharge volume should exceed reactor pressure following a scram, a check valve in the line to the scram discharge header prevents this pressure from operating the drive mechanisms.
LSCS-UFSAR 4.6-27 REV. 18, APRIL 2010 4.6.2.3.1.2.10 Collet Fingers Fail to Latch When the drive withdraw signal is remove d, the drive continues to withdraw at a fraction of normal speed. Without some in itiating signal there is no known means for the collet fingers to become unlocked. If the drive withdrawal valve fails to close following a rod withdrawal, it would have the same effect as failure of the collet fingers to latch in the index tube. Beca use the collet fingers remain locked until they are unloaded, accidental opening of the drive withdrawal valve does not unlock them. 4.6.2.3.1.2.11 Withdrawal Speed Control Valve Failure
Normal withdrawal speed is determined by differential pressures in the drive and is set for a nominal value of 3 in/sec. Withdr awal speed is maintain ed by the pressure regulating system and is independent of reactor vessel pressure. Tests have shown that accidental opening of the speed control valve to the full-open position produces a velocity of approximately 6 in/sec.
The control rod drive system prevents rod withdrawal and it has been shown above that only multiple failures in a drive uni t and in its control unit could cause an unplanned rod withdrawal.
4.6.2.3.2 Scram Reliability of CRDS
High scram reliability is the result of a nu mber of features of the CRD system. For example:
- a. Two sources of scram energy are used to insert each control rod when the reactor is operating: accumulator pressure and reactor vessel pressure.
- b. Each drive mechanism has its own scram valves and scram pilot valves. Alternatively each drive mechanism may have a single pilot valve with dual solenoid operated pilot assemblies in place of two scram pilot valves. With either scram pilot valve configuration, only one drive can be affected if a scram valve fails to open. Two pilot solenoid s are provided for each drive.
Both pilot solenoids must be de-e nergized to initiate a scram of that drive mechanism.
- c. The reactor protection system and the HCU's are designed so that the scram signal and mode of operation override all others.
- d. The alternate rod insertion (ARI) system provides an alternate means of exhausting the scram air header and closing the vent LSCS-UFSAR 4.6-27a REV. 18, APRIL 2010 and drain valves of the scram discharge volume, thereby providing an additional reactor scram mechanism which is diverse, redundant and independent of the reactor protection system.
LSCS-UFSAR 4.6-28 REV. 13
- e. The collet assembly and index tu be are designed so they will not restrain or prevent control rod insertion during scram.
- f. The scram discharge volume is monitored for accumulated water and will scram the reactor before the volume is reduced to a point that could interfere with a scram.
4.6.2.3.2.1 Reliability Analysis
A reliability analysis was performed to de monstrate that the ARI design meets the design failure rate criteria of 10
-6 failures to actuate per reactor-year (reference 5).
The probability of spurious actuation was shown to be more than a factor of 10 less likely than the probability of failure to ac tuate. The basis for demonstrating the 10 6 criteria was the complete electrical in dependence of the ARI system from the electrical portion of the reactor protection system (RPS) including power supplies. When determining the overall electrical system failure probability (ARI and RPS), the independence results in an overall failure probability well beyond any practical
means of engineering judgement (~10-11 failures to actuate per demand). Note that the mechanical portion of the CRD is unchanged by the ARI modification and now becomes the limiting factor in the overall scram system reliability. Hence, the ARI modification provides a conservative means of demonstrating adequate ATWS prevention for the expected ATWS initiators.
The charging water header pressure is monitored with a low pressure alarm to provide warning to control room operators of an impending reactor scram due to low charging-water-header pressure.
The scram assures that sufficient energy remains in the accumulators to shut down the reactor.
4.6.2.3.2.2 Control Rod Support and Operation
As described previously, each control rod is independently supported and controlled as required by safety design bases.
4.6.2.3.3 Control Rod Drive Housing Supports
4.6.2.3.3.1 Safety Evaluation
Downward travel of the CRD housing and its control rod following the postulated housing failure equals the sum of these distances: (1) the compression of the disc springs under dynamic loading, and (2) the initial gap between the grid and the
bottom contact surface of the CRD flange. If the reactor were cold and pressurized, the downward motion of the control rod would be limited to the spring compression LSCS-UFSAR 4.6-29 REV. 13 (approximately 2 inches) plus a gap of approx imately 1-1/2 inch. If the reactor were hot and pressurized, the gap would be reduced approximately 1/4 inch and the spring compression would be slightly less than in the cold condition. In either case, the control rod movement following a housing failure is substantially limited below one drive notch movement (6 inches). Sudden withdrawal of any control rod through a distance of one drive notch at an y position in the core does not produce a transient sufficient to damage any radioactive material barrier.
The CRD housing supports are in place during power operation and when the nuclear system is pressurized. If a control rod is ejected during shutdown, the reactor remains subcritical because it is designed to remain subcritical with any one control rod fully withdrawn at any time.
At plant operating temperature, a gap of approximately 1-1/4 inch exists between the CRD housing and the supports. At lower temperatures the gap is greater. Because the supports do not contact any of the CRD housing except during the postulated accident condition, vertical contact stresses are prevented.
4.6.3 Testing
and Verification of the CRDS 4.6.3.1 Control Rods
4.6.3.1.1 Testing and Inspection
The tests performed on control rods plus their related surveillance program are covered in Subsection 4.6.3.2.
4.6.3.2 Control Rod Drives
4.6.3.2.1 Testing and Inspection
4.6.3.2.1.1 Development Tests
The development drive (one prototype) te sting to date included more than 5000 scrams and approximately 100,000 latching cycles. One prototype was exposed to simulated operating conditions for 5000 ho urs. These tests demonstrated the following:
- a. The drive easily withstands the forces, pressures, and temperatures imposed.
- b. Wear, abrasion, and corrosion of the nitrided Type 304 stainless parts are negligible. Mechanical performance of the nitrided surface is superior to that of materials used in earlier operating reactors.
LSCS-UFSAR 4.6-30 REV. 13
- c. The basic scram speed of the drive has a satisfactory margin above minimum plant requirements at any reactor vessel pressure.
- d. Usable seal lifetimes in ex cess of 1000 scram cycles can be expected.
4.6.3.2.1.2 Factory Quality Control Tests
Quality control of welding, heat treatment, dimensional tolerances, material verification, and similar factors is ma intained throughout the manufacturing process to ensure reliable performance of the mechanical reactivity control components. Some of the quality control tests performed on the control rods, control rod drive mechanisms, and hydraulic control units are listed as follows:
- a. Control rod absorber tube tests:
- 1. Material integrity of the tubing and end plug is verified by ultrasonic inspection.
- 2. The boron-10 fraction of the boron content of each lot of boron-carbide is verified.
- b. Control rod drive mechanism tests:
- 2. Electrical components are ch ecked for electrical continuity and resistance to ground.
- 3. Drive parts that cannot be visually inspected for dirt are flushed with filtered water at high velocity. No significant foreign material is permitted in effluent water.
- 4. Seals are tested for leakage to demonstrate correct seal operation.
- 5. Each drive is tested for sh im motion, latching, and control rod position indication.
LSCS-UFSAR 4.6-31 REV. 13 6. Each drive is subjected to scram timing tests as required by Technical Specifications to verify correct scram performance.
- c. Hydraulic control unit tests:
- 1. Hydraulic systems are hydrostatically tested in accordance with the applicable code.
- 2. Electrical components and systems are tested for electrical continuity and resistance to ground.
- 3. Correct operation of the accumulator pressure and level switches is verified.
- 4. The unit's ability to perfor m its part of a scram is demonstrated.
- 5. Correct operation and adjustment of the insert and withdrawal valves is demonstrated.
4.6.3.2.1.3 Operational Tests
After installation, all rods and drive mech anisms can be tested through their full stroke for operability.
During normal operation, each time a control rod is withdrawn, the operator can observe the incore monitor indications to verify that the control rod is following the drive mechanism. All control rods that are partially withdrawn from the core can be tested for rod-following by inserting or withdrawing the rod and returning it to its original position, while the operator observes the incore monitor indications.
To make a positive test of control rod to control rod drive coupling integrity, the operator can withdraw a control rod to th e end of its travel and then attempt to withdraw the drive to the overtravel position. Failure of the drive to overtravel demonstrates rod-to-drive coupling integrity.
Hydraulic supply subsystem pressures can be observed from instrumentation in the control room. Scram accumulator pressures can be observed on the nitrogen pressure gauges.
4.6.3.2.1.4 Acceptance Tests
The information in this subsection is bein g maintained for historical purposes only, as it is related to pre-startup testing.
LSCS-UFSAR 4.6-32 REV. 18, APRIL 2010 Criteria for acceptance of the individual control rod drive mechanisms and the associated control and protection systems will be incorporated in specifications and test procedures covering three distinct phases: (1) preinstallation, (2) after installation prior to startup, and (3) during startup testing.
The preinstallation specification will define criteria and acceptable ranges of such characteristics as seal leakage, friction, and scram performance under fixed test conditions which must be met before the component can be shipped.
The after-installation, prestartup tests include normal and scram motion and are primarily intended to verify that piping, valves, electrical components, and instrumentation are properly installed. The test specifications will include criteria and acceptable ranges for drive speed, times settings, scram valve response times, and control pressures. These tests are intended more to document system condition
than as tests of performance.
As fuel is placed in the reactor, the startup test procedure will be followed. The tests in this procedure are intended to determine that the initial operational characteristics meet the limits of the specifications over the range of primary coolant temperatures and pressures from ambient to operating. The detailed specifications and procedures have not as yet been prepared but will follow the general pattern established for such specifications and procedures in BWR's presently under construction and in operation.
4.6.3.2.1.5 Surveillance Tests
The surveillance requirements (SR) for the control rod drive system are recommended as follows:
- a. Sufficient control rods shall be withdrawn, following a refueling outage when core alterations are performed, to demonstrate with the technical specification design margin that the core can
be made subcritical at any time in the subsequent fuel cycle with the strongest operable control rod fully withdrawn and all other operable rods fully inserted.
- b. Each partially or fully withdrawn control rod shall be exercised as defined in the Technical Specifications. When any control rod is immovable as a result of excessive friction or mechanical interference, a determination must be made and appropriate action taken.
The monthly control rod exercise test serves as a periodic check against deterioration of the control rod system and also verifies LSCS-UFSAR 4.6-33 REV. 13 the ability of the control rod drive to scram because if a rod can be moved with drive pressure, it will scram since higher pressure is applied during scram.
The frequency of exercising the control rods under the conditions of three or more control rods valved out of service provides even further assurance of the reliability of the remaining control rods.
- c. The coupling integrity shall be verified for each withdrawn control rod as follows:
- 1. when the rod is first withdrawn, observe any indicated response of the nuclear instrumentation; and
- 2. when the rod is fully withdrawn the first time, observe that the drive will not go to the overtravel position.
Observation of a response from the nuclear instrumentation during an attempt to withdraw a control rod indicates indirectly that the rod and drive are coupled. The overtravel position feature provides a positive check on the coupling integrity, for
only an uncoupled drive can reach the overtravel position.
- d. During operation, accumulator pressure and level at the normal operating value are verified.
Experience with control rod drive systems of the same type indicates that weekly verification of accumulator pressure and level is sufficient to assure operability of the accumulator portion of the control rod drive system.
- e. After each major refueling outage, each operable control rod shall be subjected to scram time tests from the fully withdrawn
position.
Experience indicates that the scram times of the control rods do not significantly change over the time interval between refueling outages. A test of the scram times at each refueling outage is sufficient to identify any significant lengthening of the scram times. Routine accumulator surveillance is performed to authenticate the discharge pressure of the CRD pump and its associated hydraulic accumulator. Accu mulator hydraulic pressure retention above the analysis valu e of 1157 psig is observed after a CRD pump trip to assure scram action via charging-water-LSCS-UFSAR 4.6-34 REV. 14, ARPIL 2002 header pressure supplied from the accumulator. The 1157 psig value for this CRD-accumulator auto scram was selected because it exceeds the analytical point where the control rod maximum insertion times were defined.
4.6.3.3 Control Rod Drive Housing Supports
4.6.3.3.1 Testing and Inspection
CRD housing supports are removed for inspection and maintenance of the control rod drives. The operational condition du ring which CRD housing supports can be removed is controlled by the Technical Specifications. When the support structure is reinstalled, it is inspected for correct assembly with particular attention to maintaining the correct gap between the CRD flange lower contact surface and the grid. 4.6.4 Information for Combined Performance of Reactivity Systems
4.6.4.1 Vulnerability to Common Mode Failures Protection of the CRDS from common mode failures is described in Subsection 4.6.2.2, and in GE's "BWR Scram System Reliability Analysis," dated September 30, 1976 (Proprietary) which was pr ovided to Mr. D. F. Ross (NRC) by Mr. E. A. Hughes (GE) by letter of the same date. The evaluation of the ECCS and SLCS against common mode failures is presented in Section 6.3 and Subsection 9.3.5 respectively. In additi on, no balance-of-plant failure will prevent reactivity shutdown. Therefore, no commo n mode failures need be considered in Chapter 15.0.
4.6.4.2 Accidents Taking Credit for Two or More Reactivity Control Systems
There are no postulated accidents evaluated in Chapter 15.0 that take credit for two or more reactivity control systems preventing or mitigating the accident.
4.6.5 Evaluation
of Combined Performance
As indicated in Subsection 4.6.4.2, credit is not taken for multiple reactivity control systems for any postulated accidents in Chapter 15.0.
LSCS-UFSAR 4.6-35 REV. 18, APRIL 2010
4.6.6 References
- 1. J. P. Fritz, "Testing of Cruciform Control Rods for BWR/6," NEDO-10565, GE APED, April 1972.
- 2. R. J. Benche, "Visual and Photographic Examination of Dresden 1 High Exposure Control Rod B87," NEDO-10541, April 1972.
- 3. R. G. Stirn et al., "Rod Drop Accident Analysis for Large Boiling Water Reactors," NEDO-10527, General Electric Co., Atomic Power Equipment Department, March 1972.
- 4. J. E. Benecki, "Impact Testin g on Collet Assembly for Control Rod Drive Mechanism 7RD B144A," General Electric Company, Atomic Power Equipment Department, APED-5555, November 1967.
- 5. "Reliability Evaluation Analysis - Unit 2 Alternate Rod Insertion System", COM-0249-R-003, February 1983.
- 6. 51-9121141-000, "Licensing Impa cts of Control Rod Withdrawal Speeds for LaSalle Reactors,"
AREVA NP Inc., September 2009.
- 7. SPC Document, "Exxon Nuclear Methodology for Boiling Water Reactors - Neutronic Methods for Design and Analysis," XN-NF-80-19(P)(A), Volume 1 and Supplements 1 and 2, Exxon Nuclear Company, Richland, WA, March, 1983.
111------......1lO 59-----51 on,r II III I--ODOODO 000'DO ODD 0 I I NUMBER OF FUEL ASSEMBLIES 764 NUMBER OF CONTROL RODS 185 NUMBER OF LPRM STRINGS 43 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FI GURE 4.1-1 CORE ARRANGEMENT REV.a-APRIL 1984 111------......1lO 59-----51 on,r II III I--ODOODO 000'DO ODD 0 I I NUMBER OF FUEL ASSEMBLIES 764 NUMBER OF CONTROL RODS 185 NUMBER OF LPRM STRINGS 43 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FI GURE 4.1-1 CORE ARRANGEMENT REV.a-APRIL 1984
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DIM.INCHES DIM.IDE NT.DIM.INCHES LSCS-UFSAR 1:::1 r-tr=========-!
II, 000000001 100000000 00000000 00000000 I oo6"OOOOI 00000000 ,..
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DIM.INCHES DIM.IDENT.
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AOc.1,.1 e c8@8@l@@)e r".0".11 O,l,jtl."OOee@l e (!D@)ee,*This value is based on 100 mil channels.Channel thickness can be 80 or 100 mil.**This data is based on GE original equipment control blades.Different control blade design are also utilized.LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-2 CORE CELL GE 8X8R FUEL TYPE REV.13 LSCS-UFSAR ooooBooooT r 00000000 0 00000 000 w PO oua w'0000 000 EO 000IH 00 QoOO 0 000 p 00 000 c 00000000 ooodEL.L£T 0 0 00000000ooooooO-FUEL ROD gg88ggg I@-WATER ROOOOOOOOI-TIE ROO 00000000 I 5
DIM.I.D.DIM.INCHES DIM.LD.DIM.INCHES***This value is based on 100 mil channels.Channel thickness vary from 80 to 100 mil.This data is based on GE original equipment control blades.Different control blade designs are also utilized.LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYS1S REPORT FIGURE 4.1-2a CORE CELL GE 8X8NB FUEL TYPE REV.13 LSCS-UFSAR ooooBooooT r 00000000 0 00000 000 w PO oua w'0000 000 EO 000IH 00 QoOO 0 000 p 00 000 c 00000000 ooodEL.L£T 0 0 00000000ooooooO-FUEL ROD gg88ggg I@-WATER ROOOOOOOOI-TIE ROO 00000000 I 5
DIM.I.D.DIM.INCHES DIM.LD.DIM.INCHES***This value is based on 100 mil channels.Channel thickness vary from 80 to 100 mil.This data is based on GE original equipment control blades.Different control blade designs are also utilized.LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYS1S REPORT FIGURE 4.1-2a CORE CELL GE 8X8NB FUEL TYPE REV.13 LSCS-UFSAR ATRlUM-9B DESIGN 4-1 I!p o B I-"I-I)Ii...I I L W'o00000000 I rf'0 0 d bocoocY 000000000\,..;.. I 000000000 I 00 boo Mol 0008 000 1 0 00 000 000 we 000 I 000 we 000 000 000 I 000 000 000000000 I 000000000 I 0000 F 000000 E OO-d I J)I ,f" K1,1'00000000&r;;.l.o FUEL ROD I 000000000 A I 000000000 I 0008 000 we WATER I 000 we 000 CHANNEL I 000 000 I I 000000000 I..H I 000000000 c...I/I I l.l====:::J S
.1 I CHANNEL FUEL ROO PELLET WATER CHANNEL I DIM.1.0.B I C D I E 1 F C H I I I DIM INCHES 0.080 I 5.278 I 0.J80 lit*1...1...lit...-T...I CONTROL BLADE'I BUNDLE LATTICe: CElL-.DIM.I.D.J I K I L l.t j N I 0 P 0>j..,>JA:
S I DIM INCHES 1.58 1 4.875 I 0.250 I lit I lit I...*'0.281T 12.00*See fu,ference 25.**This value is bHS("d on 80 mil channels.Channel thickness can be 80 or 100 mil.lk-\SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-2b CORE CELL FANP ATRlUM-9B FU8L REV.15, APRIL 2004 LSCS-UFSAR ATRlUM-9B DESIGN 4-1 I!p o B I-"I-I)Ii...I I L W'o00000000 I rf'0 0 d bocoocY 000000000\,..;.. I 000000000 I 00 boo Mol 0008 000 1 0 00 000 000 we 000 I 000 we 000 000 000 I 000 000 000000000 I 000000000 I 0000 F 000000 E OO-d I J)I ,f" K1,1'00000000&r;;.l.o FUEL ROD I 000000000 A I 000000000 I 0008 000 we WATER I 000 we 000 CHANNEL I 000 000 I I 000000000 I..H I 000000000 c...I/I I l.l====:::J S
.1 I CHANNEL FUEL ROO PELLET WATER CHANNEL I DIM.1.0.B I C D I E 1 F C H I I I DIM INCHES 0.080 I 5.278 I 0.J80 lit*1...1...lit...-T...I CONTROL BLADE'I BUNDLE LATTICe: CElL-.DIM.I.D.J I K I L l.t j N I 0 P 0>j..,>JA:
S I DIM INCHES 1.58 1 4.875 I 0.250 I lit I lit I...*'0.281T 12.00*See fu,ference 25.**This value is bHS("d on 80 mil channels.Channel thickness can be 80 or 100 mil.lk-\SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-2b CORE CELL FANP ATRlUM-9B FU8L REV.15, APRIL 2004 LSCS*UFSAR,I B we CHANNEL o p<STC0160:.DGN CHANNEL FUEL ROD PELLET WATER CHANNEL DIM LD.A I B I C D I E I F G H I I DIM INCHES 0.100 I 5.278 I 0.:38*I*I***I*CONTROL BLADE BUNDLE LATTICE CELL DIM LD.J I K ILM I N I 0 P Q I S DIM INCHES 1.58 I 4.875 I 0.260*I*I**0.261 I 12.0*See Reference 26 LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-2c CORE CELL FANP ATRIUM-IO FUEL REV.15, APRIL 2004 LSCS*UFSAR,I B we CHANNEL o p<STC0160:.DGN CHANNEL FUEL ROD PELLET WATER CHANNEL DIM LD.A I B I C D I E I F G H I I DIM INCHES 0.100 I 5.278 I 0.:38*I*I***I*CONTROL BLADE BUNDLE LATTICE CELL DIM LD.J I K ILM I N I 0 P Q I S DIM INCHES 1.58 I 4.875 I 0.260*I*I**0.261 I 12.0*See Reference 26 LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-2c CORE CELL FANP ATRIUM-IO FUEL REV.15, APRIL 2004 LSCS-UFSAR C ar:rI:1DIR o:.l TE FuelRcxl o
--1006.06.666 Dfi)OeOOeOeo 30000000000 500001)000 00°0-0000 000__0 0000000000 LatgeCmtralW amrRa:l 90410_00410(10 Notes: 1)View of bundle lattice looking down from top.2)Channel fastener is atA1 corner.LASAlLE COUNTY STATION UPDATED FINAL SAFETY AJ."lALYSIS REPORT FIGURE 4.1-2d GE14 LATTICE ARRANGEMENT REV.16, APRIL 2006 LSCS-UFSAR C ar:rI:1DIR o:.l TE FuelRcxl o
--1006.06.666 Dfi)OeOOeOeo 30000000000 500001)000 00°0-0000 000__0 0000000000 LatgeCmtralW amrRa:l 90410_00410(10 Notes: 1)View of bundle lattice looking down from top.2)Channel fastener is atA1 corner.LASAlLE COUNTY STATION UPDATED FINAL SAFETY AJ."lALYSIS REPORT FIGURE 4.1-2d GE14 LATTICE ARRANGEMENT REV.16, APRIL 2006 LSCS - UFSAR
Rev.20, APRIL2014 LSCS*UFSAR CHANNIL'ASTINI" ASlEMILY"UII..lUND""'A'1..HANOL'IIUIL"OO'NTI"'" I S'ACI"'UNUM flUIL CHANNIL"'IHQ GETTER\.OWl..Ttl'&.ATI II'HOI" 5"tI'ING'U., rTY!'.,.....'0',.I..LI'f HOSIPlICI lIu.L"OO LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1*3 (fYPICAL)FUEL ASSEMBLY (GE 8X8R SHOWN)REV.13 LSCS*UFSAR CHANNIL'ASTINI" ASlEMILY"UII..lUND""'A'1..HANOL'IIUIL"OO'NTI"'" I S'ACI"'UNUM flUIL CHANNIL"'IHQ GETTER\.OWl..Ttl'&.ATI II'HOI" 5"tI'ING'U., rTY!'.,.....'0',.I..LI'f HOSIPlICI lIu.L"OO LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1*3 (fYPICAL)FUEL ASSEMBLY (GE 8X8R SHOWN)REV.13 LAROE CENTRAL WATER ROD FUEL ROD LOWER TIE PLATE LSCS-UFSAR CHANNEL LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-3a FUEL ASSEMBLY GE 8x8NB FUEL TYPE REV.13 LAROE CENTRAL WATER ROD FUEL ROD LOWER TIE PLATE LSCS-UFSAR CHANNEL LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-3a FUEL ASSEMBLY GE 8x8NB FUEL TYPE REV.13 W<i...J a..\:::!I-et:: w a..a..::J a o a::: W<<....J a..\:::!I-<<LSCS-UFSAR l.l.J U s: w a 00 oz u\:::!ox I-..JCO l'*.+t*;0 f.1.........0 o..............*0 0-1......t....j.........u 3: o....JeD et::...J::J.<<LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1*8b FUEL ASSEMBLY FANP ATRIUM*9B FUEL REV.15,.('\PRIL 2004 W<i...J a..\:::!I-et:: w a..a..::J a o a::: W<<....J a..\:::!I-<<LSCS-UFSAR l.l.J U s: w a 00 oz u\:::!ox I-..JCO l'*.+t*;0 f.1.........0 o..............*0 0-1......t....j.........u 3: o....JeD et::...J::J.<<LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1*8b FUEL ASSEMBLY FANP ATRIUM*9B FUEL REV.15,.('\PRIL 2004 FRAMATOME-ANP LSCS-UFSAR
-'WATER CHANNEL:I': ULTRAFLOW 111 , SPACER L:.IY_.*IY"-LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-3c FUEL ASSEMBLY FA,"J"P ATRIUM-9B FUEL REV.15, APRIL 2004 FRAMATOME-ANP LSCS-UFSAR
-'WATER CHANNEL:I': ULTRAFLOW 111 , SPACER L:.IY_.*IY"-LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-3c FUEL ASSEMBLY FA,"J"P ATRIUM-9B FUEL REV.15, APRIL 2004 LSCS*UFSAR
//;.7FRAMATOME ANPPART*LENGTH FUEL ROD/'lOWER FUEL/'A FlOO ADAPTER ,/A SPRING SLEEVE ATRIUM TII 10 WATER CHANNEL FUa ASSEMal.)'
FOM IIOIUNG WATER REACTOR.LASALLE COUNTY STATION UPDATED FINAL SAFETY A.l"lALYSIS REPORT FIGURE 4.I-3d FUEL ASSEMBLY FilliP ATRIUM-I0 FUEL REV.15, APRIL 2004 LSCS*UFSAR
//;.7FRAMATOME ANPPART*LENGTH FUEL ROD/'lOWER FUEL/'A FlOO ADAPTER ,/A SPRING SLEEVE ATRIUM TII 10 WATER CHANNEL FUa ASSEMal.)'
FOM IIOIUNG WATER REACTOR.LASALLE COUNTY STATION UPDATED FINAL SAFETY A.l"lALYSIS REPORT FIGURE 4.I-3d FUEL ASSEMBLY FilliP ATRIUM-I0 FUEL REV.15, APRIL 2004 LSCS-UFSAR Spacer Fue/Rod Upper Tie Plate Debris Filter Lower Tieplate Spacer Two Large Central Water Rods Fourteen Part-Length Rods T lower Tleplate (it:..., 1-i I!
v i (1 i LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-3e GE14 Fuel Bundle (Typical)REV.16, APRIL 2006 LSCS-UFSAR Spacer Fue/Rod Upper Tie Plate Debris Filter Lower Tieplate Spacer Two Large Central Water Rods Fourteen Part-Length Rods T lower Tleplate (it:..., 1-i I!
v i (1 i LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-3e GE14 Fuel Bundle (Typical)REV.16, APRIL 2006 143 tn.
HANDLE Io iJ SHEATH..JI....1-..."-----BLAD EVELOCITY LIMITER COUPLING SOCKET-------LA SALLE COU NT.Y STATION UPDATED FINAL SAFETY ANALYSIS REPOkT FIGURE 4.1-4 GENERAL ELECTRIC CONTROL ROD ASSEMBLY Rrr.5-APRIL 1989 143 tn.
HANDLE Io iJ SHEATH..JI....1-..."-----BLAD EVELOCITY LIMITER COUPLING SOCKET-------LA SALLE COU NT.Y STATION UPDATED FINAL SAFETY ANALYSIS REPOkT FIGURE 4.1-4 GENERAL ELECTRIC CONTROL ROD ASSEMBLY Rrr.5-APRIL 1989 LSCS-UFSAR 6.5 ,n0t"'I(I (;'1(1 0 , t, II)SHEATH----...-.IJ-..,.....
143 In Coupling Release Handle COUPLING SOCKET--......_\...,--------SLAOE LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-4a GENERAL ELECTRIC ORIGINAL EQUIPMENT CONTROL ROD ASSEMBLY REV.14, APRIL 2002 I LSCS-UFSAR 6.5 ,n0t"'I(I (;'1(1 0 , t, II)SHEATH----...-.IJ-..,.....
143 In Coupling Release Handle COUPLING SOCKET--......_\...,--------SLAOE LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-4a GENERAL ELECTRIC ORIGINAL EQUIPMENT CONTROL ROD ASSEMBLY REV.14, APRIL 2002 I LSCS-UFSAR lIathium Plate SHEATH COUPLING RELEASE HANDLE (LOWER)COUPLING SOCKET----------,.
()()()()()()Cl!J UPPER HANDLE NEUTRON ABSORBEF RODS-3 AT TIP ARE VELOCITY LIMITER LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1*4b GENERA.L ELECTRIC TYPlCAL DURALIFE 215 CONTROL ROD ASSEMBLY REV.14.APRIL 2002 I LSCS-UFSAR lIathium Plate SHEATH COUPLING RELEASE HANDLE (LOWER)COUPLING SOCKET----------,.
()()()()
()()Cl!J UPPER HANDLE NEUTRON ABSORBEF RODS-3 AT TIP ARE VELOCITY LIMITER LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1*4b GENERA.L ELECTRIC TYPlCAL DURALIFE 215 CONTROL ROD ASSEMBLY REV.14.APRIL 2002 I LSCS-UFSAR LASALLE COUNTY ST A nON UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.14c GENERAL ELECTRIC TYPICAL MARATHON CONTROL ROD ASSEMBLY REV.14.APRIL 2002 LSCS-UFSAR LASALLE COUNTY ST A nON UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.14c GENERAL ELECTRIC TYPICAL MARATHON CONTROL ROD ASSEMBLY REV.14.APRIL 2002
"-J_....WET sTEAM-RETURNING WATER STEAM WATER MIXTURE\i.WATER LEVEL RETURNING WATER TURNING VANES (INLET NOZZLE)STANDPIPE CORE DISCHARGE LA SALLE COUNTY STATION UPDATED FINAL"SAFETY ANALYSIS REPORT FIGURE 4.1-5 STEAM SEPARATOR REV.0-APRIL 1984"-J_....WET sTEAM-RETURNING WATER STEAM WATER MIXTURE\i.WATER LEVEL RETURNING WATER TURNING VANES (INLET NOZZLE)STANDPIPE CORE DISCHARGE LA SALLE COUNTY STATION UPDATED FINAL"SAFETY ANALYSIS REPORT FIGURE 4.1-5 STEAM SEPARATOR REV.0-APRIL 1984 STEAM DRYE R SKIRT VANES COLLECTING TROUGH LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-6 STEAM DRYER REV.a-APRIL 1984 STEAM DRYE R SKIRT VANES COLLECTING TROUGH LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-6 STEAM DRYER REV.a-APRIL 1984
- LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-7 STEAM DRYER PANEL REV.0-APRIL 1984***LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.1-7 STEAM DRYER PANEL REV.0-APRIL 1984 CEN11!RINQ SPRING COHTAOl ROD GM TO#'GWOE LA SALLE COUNTY STATION UPDATED FrNAL SAFETY ANALYSIS REPORT FIGURE 4.2-1 SCHEMATIC OF FOUR BUNDLE CELL ARRANGEMENT REV.a-APRIL 1984 SCHEMATIC OF REACTOR ASSEMBLY SHOWING THE LEAKM}E FLOW PATHS34 NOTE: PERIPHERAL FUEL SUPPORTS ARE WeLOl'iD INTO THE COAE SUPPORT PLATE.FOR THESE/BUNDLES.PATH NUMBERS 1.2, 5, AND 7 DO NOT EX 1ST.6 FUEL SUPPORT IN-cORE GUIDE ruBe...I SHROUD CONTROL GUiDE TUBE CHANNEL o LOWER TIE PLATE 7 I.CONTROL ROD GUIDE TUBe*FUEL SUPPORT 2.CONTROL ROO GUIDE TUSE*CORE SUPPORT PLATE 3.CORE SUP!"Ol'tT PLATE*INCOAe Tuae 4.COl'll!SUPPa"T PLATE--SHROUD 5.CONTROL ROO GUIOE TUBE*DRIVE HOUSING G.FuEL SUPPORT'l..OWEA He PLATE 7.CONTAOL ROD DRIVE COOl.ING WATER a CHANNEL-LOWER TIE PLATE 9.LOI'VER TIE PLATE HOl.ES{lWO/A$SEMBLYj CONTROL ROD DRIVE HOUSING LA SALLE COUNTY STATION VPDiHED f HUlL SAfETY ,l\NAl YSI S REPORT FIGURE 4.2-2 BYPASS FLOW PATHS0-APRIL 1984 ,
,.FIGURE 4.2-3 FUEL BUNDLE eXeR BP8X8R YJEL TY?ES LA SALLE COUNTY STATION UPDA7ED FiNAL SAFETY ANAL vsIS REPORT II I , ii" I I1---------------..",;1
[i i I X I-<:):2 eo\.\I i.J a.JC lU LoU II!.:.).....<Ii Il.*;: C III 0;>a:III....J U"" o:t:>LL...I.....<-...1-""r""-.-,....flI'......w (I Z II:...z i*I I<<x...l-J.....::--.::;: '"'""'" f""1"'"'4-APRIL 1988 EXPANSION SPR:NG SPACER LOCKING TAe WASHER LARGE'W,A.TER ROD LowER TIE PLATS j},...----------.,1 LA SALLE COU NTY STATION"!UPDATED FINAL SAFETY ANALYSIS REPORT 1.f J---------------...1, I l i FIGUR£4.2-3aIl!I GE I i4- 1988 LSCS-OFBil]!I Improved lie plote c:luigl"l.
Ir"--Increosed plen...'" volumein Gooolinio fods-+Circular ferrylt high per10rmllnce 5poeer Laroe woter tube[1..3'4 inl:n O.D./0.040 ineh thiclenessl os spoeer ccptur.IfF I...-Rel:Juced tube 0.0.1I If thfu UTP tlq IJ I,K I 11,1 inch of std.aEeS v II 0.591 inch W/R tl.lbel rI
...-:::;:
Helium r-,..preprest",rizatiol"l level 15 A nu.Godolinio stu/ldown Zone' I r-O.64tO inch..."&I&1 V rod 0 red l:utcllio 0.411 inch pd.t o.o.*I 1,..--0.483 inch fuel rod 0.0.'0,410 ineh pellet length.---_....JLIlll.
12 inch NA TU illcrn,k.t 01 top and Ij inch at bottom LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE: 4.2-3b\I ,!Fuel Bundle GE 8X8NB Fuel Type REV,8-APRJl 1992 o o a:: Io&J...<-I Q.Io&J i=<t LSCS-UFSAR l.U U:>w o 0<':>oz a::S2 u I--ICOIWl0----u::I...t t..**i..,_00*..............0 0--...t't**+..t..+*.<LASALLE COUNTY STATION UPDATED FIN.AL SAFETY ANALYSIS REPORT FIGURE 4.2-3c FUEL BUNDLE FANP ATRIUM*9B TYPE REV.15, APRIL 2004 LSCS-UFSAR v U i'III c Ll\.SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.2-:3d FUEL BUNDLE FANP ATRIUM*10 TYPE REV.15, APRIL 2004 I LSCS-UFSAR J I')x....T iIol.......)oIl;""";:-
"'" I ('""'.....T I Spacer Lower Tieplate Upper Tie Plate Two Large Central Water Rods FoUrteen Part-Length Rods Spaoer Debris Filter Lower Tleplate LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.2-3e GE14 Fuel Bundle (Typical)REV.16, APRIL 2006 LSCS - UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT Figure 4.2-3f GNF2 Fuel Bundle (Typical)
Rev. 20, APRIL 2014 LUCS-UFSAR REV. 14, APRIL 2002
Page Intentionally Left Blank LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.2-4 DELETED GUIDf TUBE'JE LOC In CONT R OL"'aD DR HOUSING GR1FICf./RflL.LER$/1 f,JEL$I.JPFORT CASTING COPE SUPPOFlT PLATE_--.I----+--.-
CONTROL rwo 144,0, STROKE COLJPUNG LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.2-5 CONTROL ROD VELOCITY LIMITER REV.0-APRIL 1984 LSCS-UFSAR"i--",
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.\LASALLE COUNTY STATIONUPDATEDFINAL SAFETY ANALYSIS REPORT FIGURE 4.2-5a FABRICAST VELOCITY LHVllTER REV.14.APRIL 2002 IoSCS"UFS"'Ut 1l1Q iL 10)'4.....II: l::J""I.....*E i" z......is*g(,)!I.ASALJ.r:
(:OUNTY ST,\TIOr-:
UPDATED FI)J:\I, SAFETY ANAl.YSIS REPORT FIGURE'l.2-6
('1'il'ICAL)
TEMI'ERATURE VB HEAT FLUX-DOL BX8R FUEL TY?E HEV.1.1 LSCS-U'SAJl 1llC1O LAS.".LLE COUNT'{STATTON UPDATED SAFEr):".'J"ALYSIS REPORT FIerRF.4,2*7 mL'ICAL)
TF.MPF:RATl!Rf:
VS.HEAT PLUX-Uf'£8X8R FeEL 1'\"PB to ,-.-----------------------------------,_l 1.101 A$$LWlJI'TIONS; IH ALL ADDS IN THE SI,IHOLf Ar'tHE SAME n:UPI!"AT1JItE I2l ACDS Al!A.CH IHOtCATEI:I 100 L:\SALLE COUNTY STATION UPDATED FfNN, S,\FE'ry ANALYSrS R!';f>ORT E:.'-lERGY RELEASE AS 1\
OF Trr,m r,TYPICAL)
REV, 13 LSCS-UFSAR THIS PAGE INTENTIONALLY LEFT BLANK LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.3-1 a UNIT 1 CORE LOADING MAP FIGURE 4.3-1a REV. 10 -APRIL 1994 LSCS-UFSAR THIS PAGE INTENTIONALLY LEFT BLANK LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.3-1 b UNIT 2 CORE LOADING MAP FIGURE 4.3-1b REV. 10 - APRIL 1994 REV. hllhfidNI)l 4 w A: RTL =9~8 0 3 Y W a s X
- L LA SALLE COUNTY STArIQN UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.3-2 L K- AS A FUNMON OF EXPOSURE AT VARIOUS VOID FRACTIONS, HIGH ENRICHMENT, DOMINANT FUEL TYPE ~=y ~~~
rv d NQf.J.~b~b~ ~'IO.tb+ C~Z~d tv c~ c c~ c~ v G NCYil~tf5dd Wt.}lt+' ~Z~d PZ'u . L - APRIL 1988 N0il:)VtfJ W01'15ezil FIGURE 4.3-3 tv 3 x W eti m a a c C ^. v NQh1~11tid Wd1V~ g~~ft LA SALLE C4UN'1'Y S'TJ~kT10N UPDATED FINAL SAFETY ANALYSIS R£PGR'f ATOM FRACTION AS A FUNCTION OF EXPOSURE, HIGH ENRICHMEIT, DOMINANT FUEL. TYPE. 4 , 014 10I DS { TYP=u:_: i Swoilotri~zl N ISSl~ 9-+ETd. 4 - APYIL 198,8 FIGURE 4.3-4 :u LA SALLE COUNTY STATION UPDATED FINAL SAF Y ANALYSIS REPORT FISSION FRACTION AS A FUNCTION OF EXPOSURE, HIGH ENRICHMENT, DOMINANT FUEL TYPE, 4 04 VOIDS (.7IPIC 11 f r.SCS-UFSAR 4 R 9 7 9 EXPOSUR9 (Gw4h) 9 I-ARALLE COUNTY STATION UPDATED FINAL SAFETY A 1ALYM REPORT vlc~fRrl 4.,3-5 (TYPICAL)
NEL"i'RON G&NERATION TItiIE VS_ EXPMURE AT 40 PERCENT VOIDS RED'_ 13 1Q 1e LSCSXFSA R. 15 14 10 C?' = ENAICHMENT FU IL 1J3 &d"ICWtt3" RUEL 0.711 FUEL (NATilrt^L w 5 i FJtpo"! (aftht 10 I,AS4,L1X COU. N" N STATI ON L'PDMED FINAL SAFb'TY ANALYSIS REPORT FIGURE 4.3-6 (TI'PICAL)
DELAYED NEUTRON FRACTION VS. EXPOSURE 4T 40 PERCENT VOIDS REV. 13 LSCS-UFSAR THIS PAGE INTENTIONALLY LEFT BLANK FIGURE 4.3-7 REV. 4 - APRIL 1988 LSCS-UFSAR THIS PAGE INTENTIONALLY LEFT BLANK FIGURE 4.3-8 REV. 4 - APRIL 1988 LSCS-UFSAR THIS PAGE INTENTIONALLY LEFT BLANK FIGURE 4.3-9 REV. 4 - APRIL 1988 LSCS-UFSAR THIS PAGE INTENTIONALLY LEFT BLANK FIGURE 4.3-10 REV. 4 - APRIL 1988 LSCS-UFSA-P 0 FIGURE 4.3-11 RE'. 13 LASA2.LR COUNTY STATION N N W J K K YJ Q a UPDATED FINAL SAFETY ANALYSIS REPORT BFGftvh-INC OF CYCLE AND END OF CYCLE CORE 81 ,'ERAGE XXIAL POWER - ?G4 CORE, BWIV4 AND $WEilsi x 4 d w-Z w 1 6 -t2 ,I SCS-ITFSAR 40 50 PERCENT volas lalml LASALLE COVINT1' STATION UPDATED FINAL SAFETY ANALYSTS REPORT FIOUkh 4_3-tl MODERATOR VOID RE!tCTMTY COEFFICIENT AT FOC-l INITIAL CYCLE RIM 13 LSC~J11{ p 1qQ i40 I20D Ww 2000 AVEAACE fUEL TEWA CRATURE PCi 2400 2804 e b Gwett LASMIE WUN'1Y hI'ATION UNP,1TEU I?INAL SAH'Fl'Y ANALYSIS RKI}r}lid' FIGURE 4.3-1.1 DOPPLER REACH VITY CQIT'FICI I~NT ,1S A FIJN(C(UN OF VU EL FXPOSURR AN 1) AVF.RACl? Fi"KLTEMNRRJi`!'URE. A1' ,W AVER-AGE vUm caN'rvNT OF mm i iju l N'N li~(a (kII:N`i' (INITIAL CY(,I,i , ,)
2.00 1.50 1.00 us 0.50 v 0.00 ISCS. U FS_AR Cold Shutdown Margin Cyde Exposure, +GWD/MT 0.0 - 5.0 10.0 15.0 20.0 LASALLE COUNTY STATION tTP1)!lTKf)
PINAL SctPP.TY ANALYESIS
?,PORT FIGURE 4.3-14 E?LWFIX OF A COLD SH1ITROWN 34ARGIN CURVE REV. 13 43 . 19 ..--..--- 25 07 02 t}d 10 14 1B 22 26 30 34 38 4? 46 50 54 58 - IA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.3- l S CONTROL ROD ASSIGNMENTS FOR GROUPS I THROUGH 4 (SEQUENCE A) RrV. 0 -- APRTL 1984 a~0~o~o~a~a~a~a 0 mammon iaiaioiaiaioioi 39 27 23 REV. 0 - APRIL 1984 42 06 10 14 15 22 26 30 34 35 42 46 50 54 S,a LA SALLE C"OUNTY STATION UPDATLD FTNAL SAFETY ANALYSIS REPORT FIGURE 4.3-16 CONTROL ROD ASSIGNMENTS FOR GROUPS a THROUGH 1 {7 (SEQUERC A ) ON o~a~no a~o~a on mmummmus =-SOUND REV. 0 - APRIL 1984 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT CONTROL_ ROIL ASSIGNMEN7S FOR GROUPS I THROUGH 4 (SEQUENCE B) sioioioioiaioio 59 r 3 .. $!) i '1 t . 51 3 4 3 41 r 2 1 2 43 --3 .. 3 4 39 -1 2 1 15 -.. 3 ... 3 3.1 -2 1 1
- 21 -.. 3 4 3 23 -t 2 1 19 -3 4 3 .. lS '-2 1 2 11 4 3 (J1 2 1 OJ L 3 .. 06 III l4 13 21 26
- 0 -A?R!L 1984 " 3 :J 2 1 2 i 3 .4. 3 , 2 1 2 M 4 l 4 J '1. 1 '1 1 3 " 3 c 1 2 1 2 l 4 3. 4-l 1 2 t .. 1 4 3 1 1-1 2 ..J l ... 3 2 1 2 .. 3 " 30 35 42 45 50 54 sa LA SALLE COuNTY ST ATION UPDATED FIHAL SAFETY ANALYSIS REPORT FIGURE 4.3 .. 17 CONTROL ROD ASSIGNf.1£N7S FOR GROUPS 1 THROUGH 4 (ScQUEi',CE B}
43 39- 35 27 19 ---.-....- FIGURE 4.3-13 Q 06 10 14 18 22 2f, 30 34 38 42 46 50 54 sa LA SAILt-E COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT COUTROL ROD ASSIGNMENTS FOR GROUPS B THROUGH 10 (SEQUENCE B ) REV. 0 - A-PRIL 1984 oia Iwo mmmummmummonsam aoioiIm TID CONTROL FRACTION LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.3-1D HOT OPERATING EOO-1 SCRAM REACTIVITY REV. 0 -- APRIL 1984 i.SCS-ITFSAX X 4 t "xV '"AisonmNaNvx LA.SALLE COUNTY STATION UPDATED FFN.~L SAFETY ANALYSIS REPORT PIGUR8 4.3-20 ti t W XENON REAC11VITY BUILDUP AND BURNOUT AFTER SHL"TDOWN (TYPICAL}
RED'. 13 REV. 19, APRIL 2012AT3323MWt LSCS-UFSAR
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LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4,3-21 TYPICAL RADIAL POWER DISTRIBUTION LSCS-UFSA R REV. 19, APRIL 2012AZIMUTHALFASTFLUXDISTRIBUTIONFIGURE4.3-21aUPDATEDFINALSAFETYANALYSISREPORTLASALLECOUNTYSTATION REV. 19, APRIL 2012AT3323MWt LSCS-UFSAR ,K,
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4:1 .0 LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.3-22 TYPICAL AXIAL POWER DISTRIBUTION LSCS-UFSA R REV. 19, APRIL 2012AXIALFASTFLUXDISTRIBUTIONFIGURE4.3-22aUPDATEDFINALSAFETYANALYSISREPORTLASALLECOUNTYSTATION 1.0 0.02 001 D.D!0,02 iOS a IG Of;C:AV l';ATIO X2lXo uo10 LA SALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-1 DAMPING COEFFICIENT VS, DECAY RATIO (SECOND ORDER SYSTEMS)REV.0-APRIL 1984 REACT!VITY...PERTURBATI01>l REACTOR K!NETrCS t
__....,.......NEUTRON flUX RESPONSE TOrAL REACTOR REACT IIIL T't FEEDBACK TOTAL[NDIVIDUAl CHANNEL TYPE REACTIVITY FEEDBACK FROM OTHER CHANNEL TYPES+REACTIVITY TO POWER TRANSFER FUNCTION AT CONSTANT INLET FLOW REACTIVITY TO flOW TRANSFER FUNCTION AT CONSTANT POWER ROIl TO POWER TRANSFER fUNCTION TO QTH[R CHANNELS LA COUNTY STATION UPDATID FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-2 HYDRODYNAMIC ANO CORE STABILITY MODEL REV.0-APRIL 1984
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LSCS-UFSAR FIGURE 4.4-4 REV. 4 - APRIL 1988
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LSCS-UFSAR FIGURE 4.4-5 REV. 4 - APRIL 1988
THIS PAGE INTENTIONALLY LEFT BLANK LSCSUFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE INITIAL CORE 10 PSI PRESSURE REGULATOR SETPOINT STEP AT 51.5%RATED POWER (NATURAL CIRCULATION)
Rev.14.APRIL 2002 LSCS-UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-7 INITIAL CORE 10 CENT ROD REACTIVITY STEP AT 51.5%RATED POWER (NATURAL CIRCULATION)
Rev.14, APRIL 2002 15CS UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-8 INITIAL CORE 6-INCH WATER LEVEL SETPOINT STEP AT 51.5%RATED POWER (NATURAL CIRCULATION)
Rev.14, APRIL 2002 LSCS-UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4A9 INITIAL CORE 10 PSI PRESSURE REGULATORY SETPOINT AT 105%RATED POWER AND 100%RATED FLOW Rev.14, APRIL 2002 LSCS UFSAI<LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-10 INITIAL CORE 10 CENT ROD REACTIVITY STEP AT 105%RATED POWER AND 100%Ri\TED FLOW Rev.14, APRIL 2002 LSCS*UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-11 INITIAL CORE 10%LOAD DEMAND STEP AT 105%RATED POWER AND 100%RATED FLOW Rev.14, APRIL 2002 LSCS UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4,4-12 INITIAL CORE 6-INCH WATER SETPOlNT STEP AT 105%RATED POWER AND 100%RATED FLOW Rev.14, APRIL 2002 LSCS-UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-13 INITIAL CORE 10 PSI PRESSURE REGULATOR SETPOINT STEP AT 68%POWER AND 50%RATED FLOW Rev.14, APRIL 2002 LSCS-UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-14 INITIAL CORE 10 CENT ROD REACTIVITY STEP AT 68%POWER AND 50%RATED FLOW Rev.14, APRIL 2002 LSCS-UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-15 INITIAL CORE 10%LOAD DEMAND STEP AT 68%POWER AND 50%RATED FLOW Rev.14, APRIL 2002 LSCS-UFSAR LASALLE COUNTY STATION UPDATED FINAL SAFETY ANALYSIS REPORT FIGURE 4.4-16 INITIAL CORE 6-INCH WATER LEVEL SETPOINT STEP AT 68%POWER AND 50%RATED FLOW Rev.14.APRIL 2002
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